JWES−AE−9607

ReportononlntegritytheIntemationalof

NuclearComponents

May8，1996

1996年8月

社団法人

Workshop

日本溶接協会

原子力研究委員会

一てπ一

気
・だ響

翼鼠蕊魍饗謹駕撫鍛」魍ρ9Workshopの会場にて（中央が矢川教授、その右がKim教授）

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挨拶するDr，Yoon（KINS副所長）

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挨拶する矢川教授

帽一一国圃噂一隔

一一

暫Al：

96，5．7〜8

李剣：KINS，。

挨拶するKim教授

鹿島委員の講演K口NSの入ロで

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網暑馨1324。・・蛍、a5ヨー1費一．臼…E寒蓼馨二騨0㎝β、

使用済み燃料の乾式貯蔵施設とその説明図

序

世界のエネルギー事情は大きく変わりつつある。特に、アジア地域では2010年の需要見直しが、1992年の2倍にも達するという急カーブを描く。この需要激増に対して、この地域での原子力発電が、新しい供給源となることは間違いない。

これまで、我が国の原子力開発は欧米との協力のもとで行われることが主であったが、原子力発電の安全確保などの問題については、地理的つながりの強い地域内での共通課題
として問題を共有し、かつ協力していくことが重要である。

我が国は、これまで長年にわたって原子力発電の良好な発電実績を有し、またその安全技術について国際的に高い評価を受けてきた。この優れた安全運転の実績は、改めて言うまでもなく、高信頼性の材料・機器設計のハードウエア面と、高度な運転管理技術・保守
技術・検査技術・教育訓練などのソフトウエア面の両者がベースとなっている。このような背景のもと、今後、アジア地区での原子力開発がますます活発化するにあたって、我が国からの原子力発電機器・材料の輸出も盛んになっていくことが予想される。

このような状況にあって、原子力機器・材料に関わっている研究者・技術者の交流も極めて重要度を増しつつある。今回のワークショップの目的はまさにこの点にあり、原子力発電の安全確保の最重要事項の1つである機器コンポーネントの健全性に関する討議を、我が国と韓国の研究者・技術者を一同に会して行うことにした。
本報告書は、このワークショップ、さらには付随して得られた情報をまとめたものである。

ワークショップ開催にあたっては日韓両国の関係者の献身的御努力があった。この場を借りて厚く御礼申し上げたい。（矢川）Wednesday, May 8
9:00 Registration

9:20 Opening Ceremony
Welcoming Address
Dr. Young.S Eun(Vice-President, KlNS)
Prof. Y.J.Kim(Professor, SungKyunKwan Univ.)
Prof. G. Yagawa(Professor, Univ. of Tokyo)

Session A [Meeting Room #208][Chairman, Prof. Y.J.Kim]

1 10:OO Application of Leak-Before-Break Using Piping Evaluation Diagram(PED)
Y.J.Yu*, K.S. Yoon, S.H.Park, K.B. Park(KAERI),

Y.J.Kim(SungKyunKwan Univ.)

2

lO:25 A Research Program for Dynamic Fracture Evaluation of Japanese
Carbon Steel Pipes

K.Kashima*(CRIEPI)
3 10:50 Effects of Dynamic Strain Aging on the Leak-Before-BreakAnalysis in SA 106 Gr.C Piping Steel

I.S .Kim*, J.W.Kim(KAIST)
4 1 1 : 1 5 Probabilistic Fracture Mechanics Analyses of Nuclear Pressure

Vessels Under PTS Events
G.Yagawa*(Univ. of Tokyo), S. Yoshimura(CRIEPI),

M. Hirano(JAERI)
5 1 1 :40 Improved Toughness of the SA 508 Class 3 Steel for Nuclear
Pressure Vessel Through the Steel-Making and Heat Treatment

J.T.Kim*, H.K.Kwon, H.S.Chang(HANJUNG)
12:05 Lunch

Session B [Meeting Room #208][Chairman, Prof. G.Yagawa]

6 1 3:30 Crack Shape Evolution of Surface Flaws under Fatigue Loading ofAustenitic Pipes

I.S. Hwang*, J.H.Kim(Seoul National Univ.)
7 1 3 :55 Study on the Effect of the Crack Length on the Jlc Value

M.Kikuchi*(Science Univ. of Tokyo)

8 14:20 Development of New Z-factors for the Evaluation of the
Circumferencial Surface Crack in Nuclear Pipings

Y.H.Choi*, Y.K.Chung, Y.W.Park, J.B.Lee(KINS),
G.Wilkowski(Battelle Memorial Institute)

9 14:45 Requirements for Pressure Boundary Integrity of Operation
Nuclear Plant-Japanese Standards
H.Kobayashi*(Tokyo Institute of Technology)

10 1 5: I O The Development Status of Mechanical Component Code for
Nuclear Power Application in Korea

N.H.Kim* , J.S .Nah(KOPEC)
15:35 Coffee BreakSession C [Meeting Room #208J[Chairman, Dr. Y.W.Park]1 1 1 5 :50 Progress of Component Aging and Structural Integrity Research

Program at JAERIK.Shibata*( JAERI)

12 16:15 Development of Expert System for Nuclear Piping Integrity
Y.J.Kim*, M.W. Suh, C.S. Seok, H.K. Jun(SungKyunKwan Univ.)
Y.W.Park, Y.H.Choi, J.B.Lee(KINS)

13 16:40 Recent Progress in Korean Nuclear PLIM Program
T.E. Jin*, H.J. Choi(KOPEC), I.S.Jeong, S.Y. Hong(KEPCO)
17:05 General Discussion on the Presented Topics
1 7:30 Discussion on Future Activities of Workshop and Conclusion
l 8 :OO Dinner hosted by Japanese Participants at KlNS' Restaurant次

目

1．計画と経緯と概要………………………一・・…………・…一・…… 1

2．日韓ワークショップ概要…一。・・・・・・…。…一一・…。・・。・…◎・・・・…。・…一・4

2．1セッションA…………… ・・・・…。・・・・・・・・・・・・・…。…6・・・・・…。・・・・・・・… 4

2．2セッションB…………… ……………・…・D……………………… 6

2．3セッションC…………… ……………・…・・……………………… 9

3．サイト見学報告………………………………………………………11

4．成果と今後の課題
5．今後の計画…………

………………一・曹・・・・・・・・・・・・…

。・・・・・…

一一・一の一一一一・一・…

D・・…

6…

。…

。・・・・・・・・・・・…

のD・・・・・・…

6・・・・…

14

15

5．1今後の研究交流について…………………………………………15

5．2講演論文の公表について…………………………………………15

6．あとがき7．付……………………………………………………一・・……16

録……………・・……

。・・・…

の6・一・・・・・・・・…

。・。・・・・・・・・・・・・・・・…

6・…

17

1

計画の経緯と概要
日韓の原子力技術の交流の計画が具体化したのは、平成7年6月に韓国原子力研究所の若い研究者Dr．H．Y．Leeが訪日したのがきっかけとなった。

川崎重工業（株）の国松委員の紹介により菊池がDr．Leeと合い、日本側の要望の伝達を依頼した。その二週間後、同研究所の主任研究員であるB．Yoo氏が来日し、東大の矢川教授室にて具体的な協議を行う運びとなった。

Yoo氏は帰国後、韓国機械学会内の原子力研究部門にて日本側の要望を検討し、積極的に進める事を決定した。韓国側の担当は成均館大学のKim教授とK

I

N

S（韓国原子力安

全研究院）のDr．Parkに決定した。

日本側もこうした事態の進展を受けて、平成7年11月の原子力研究委員会において正式に開催を決定し、国際研究連絡小委員会を母胎として、矢川委員長と菊池委員とが折衝に当たる事となった。

開催の時期は日本の連休明けの5月第二週とし、講演依頼、講演要旨の募集、原稿執筆

依頼、と準備は順調に進展し、予定通り5月8日に開催の運びとなった。
韓国では、KI

N

Sが主催する「原子力施設・機器の安全性に関するシンポジウム」が1993年以来すでに二度開催され、毎回70〜80名近い参加者を得て活発に行われてきた。
今回の企画は第3回目のシンポジウムと合同で企画されたものである。韓国国内のシンポジウムは前日の7日に同じ会場で開催された。7日の夕食会に招待された我々日本側は、予期しなかった多数の参加者に会って驚かされた。

翌8日9時30分からWorkshopは開始した。出席者は約70名であった。
冒頭にK

I

N

S副所長のK．S．Yoon博士より「韓国国内で原子力施設の安全性研究の重要性が高まっている事からこのWorkshopの成果に期待している事、またこのような技術交流の場が将来はアジア地域の共同の場となることを望んでいる」との挨拶があった。

続いて成均館大学のKim教授の挨拶ではOpen

MindedDiscussionを行おうとの呼びかけがあり、日本側を代表して矢川教授からはアジアでの原子力大国としての日一韓一台の将来の協力が重要になるとの指摘がなされた。

講演会では日本側から5件、韓国側から8件の論文が発表され、双方から活発な質疑・応答が行われた。以下にそれらの要約を、参加者の感想を含めて記す。また参加者の構成と発表テーマを韓国側でまとめているので、参考までにそれらを図1，2に示す。ただし
これは前日行われた韓国の国内シンポジウムも含めた結果である。

講演会終了後、今後の活動方針がDr．Parkより発表された。その詳細は後に記す。

その後、日本側の主催により夕食会を開催した。手配は韓国側にお願いしたが、蟹や肉一1一などの韓国料理とともにお寿司や赤飯、刺身がならび、大変なごちそうであった。
KI

N

Sのある大田市は日本のつくば市のような科学技術都市であり、多くの韓国側の出席者は車で来ているとの事で、あまりお酒が進まなかったのは少々残念であったが、たっぷりと懇親の時を持つ事ができたのは有意義であった。さらにKIN

Sの若手の研究者数

名は、夕食後に我々のホテルを訪ねてくれて、深夜まで楽しく歓談を続けた。（菊池）一2一Number◎ぜPapers329

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楓◎ε。逡織脚毒鱗
肇◎％

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図1

研究分野

C◎磯麟r竃bu舗o轟癒oP繍per＄窓29u醐鯉C◎織鱒輪y3％

鐸％1轟欝敏稔繊鯵

4㊨％u禰v馨鵬鷺y34％

図2

発表者の所属一3一

2

日韓ワークショップ概要

2．1

セッションA

セッションAでは、5件の論文（日本側3件、韓国側2件）が発表された。内訳は、配管のLB

B関係が3件、圧力容器関係が2件である。以下に各論文の発表要旨を示す。

1〉，，Application

of：Leak−Befbre−Break

Using

Piping

Evaluation

Diagram，，，by

Y．J．Yu，KS．Ybon，S。H．Park，K．B．Park（KAERI），Y。J．Kim（Sung

Kyun

Kwan大学）
韓国原子力研究所Yu氏より、配管評価線図（PED）を用いたL

B

B適用性の評価に関する発表があった。配管システムにLBBを適用するためには、配管形状、材料特性、配管の荷重など、プラント固有のデータが必要である。従って、配管設計とその手順が既知でないと、L

B

Bの評価は不可能である。本論文では、P（通常運転荷重とS

S

EDと呼ばれる配管評価線図

E荷重との関係をき裂長さをパラメータとして示した線図）をL

B適用箇所（例えば、配管一ノズル接合部）に対して作成し、設計段階でもL

B

B

B適用性

を評価できる手法を示した。本手法は、韓国次世代炉のLBB設計に活用される予定である。

2〉

A

Research

Steel

Programfbr

Pipes

by

Dynamic

Fracture

EvaluationofJapanese

CarbonK．Kashima（CRIEPl）

電力中央研究所の鹿島より、日本の炭素鋼配管に関する研究計画について発表があった。

本研究は、日本のS

T

S410炭素鋼配管の破壊に及ぼす動的荷重の効果に関するものであり、材料試験、配管試験、解析評価からなる。材料試験では、流動応力は歪み速度にあまり依存していないこと、配管試験では、破壊荷重は準静的荷重、動的荷重両者に対しほぼ正味応力基準（塑性崩壊基準〉で予測出来ることが示され、顕著な動的効果は認められないとの報告があった。さらに、繰り返し荷重下でのき裂進展からその後の不安定破壊に至る一連の挙動を予測する破壊力学評価手法を開発し、実測結果との良好な一致を示した。

3〉，，Effectof

SA106Gr．C

Dynamic

Piping韓国科学技術院のK
動的歪み時効（DSiStrainAgi皿gon

the：Leak−Befbre・Break

Steel，，byI．S．Kim

m教授より、S

A106鋼G

and

Analysis

in

J．W．Kim（KAIST）

r．C配管材のL

B

B評価における

A）の効果に関して発表があった。各種温度、荷重条件下で得られる

一4一

引張り試験データ、J−R試験データをもとに、漏洩検知可能なき裂寸法、不安定破壊時
のき裂寸法に及ぼすDSAの効果を評価して、結果を

LBB

Allowable

LoadWindowとしてまとめた。不安定破壊荷重は、プラントの運転温度域で荷重速度に依存する。このため、

L

B

Bの成立範囲は、250〜300℃付近では室温に比べて約30％減少し、また、高荷重速度下でも減少することが認められた。

4）

Probabilistic

under

PTS

Fracture

Events，，by

Mechaniucs

Analysesof

NuclearPressure

G．Yagawa，S．Yoshimura（UnivesityofVessels

Tokyo），

N．Soneda（CRIEPI），M．Himno（JAERI）
東京大学矢川教授より、圧力容器の確率論的破壊力学解析に関する発表がなされた．論文では、日本機械学会のR

C111委員会で行われた圧力容器の破壊力学解析、特に加圧

熱衝撃下における圧力容器の破損確率に対する評価手順の開発に関する研究活動が紹介された。米国ではE

P

R

I、NRCにおいてP

T

Sのベンチマーク研究が行われており、これに対し、日米で開発された4種類の解析コードを用い、破損確率に及ぼすデータの不確実性に関する各種の感度解析を実施した。

5〉，，Improved
VbsselToughnessthroughthe

of

the

Steel−Making

SA508Class3Steel

and

Heat

fbr

Treatment，，byNuclearPressure

J．T。Kim，H．K．

Kwon，H．S．Chang（HANJUNG）
韓国技術研究院のKim氏から、圧力容器鋼SA508C13の製造、熱処理における靭性改善について発表があった。真空炭素脱酸法（VCD）では、靭性値は基準を満たすものの低い値であり、改良VCDとして、最小冷却速度（15C／min）を用いることが推奨された。さらに、アルミニウムとシリコンキルド鋼を加えることにより、十分な靭性（KIC）を確保できることを示した。これは組織の細粒化によるものであり、粒の寸法は、V

CDで50ミクロンであったものが、改良VCDでは20ミクロンの値となることが報告
された。

（所感）

セッションAの韓国側の発表では、配管のLれた。韓国におけるL

B

B

B関連の成果がいくつか報告さ

B研究は比較的最近開始されたものであるが、諸外国の研究成果を急速に取り入れて国内プラントの評価を実施し、一部プラントにはLのことであり、韓国側の関心の高さがうかがわれた。L

B

B

Bを適用済みとBに対する基本的な考え方、アプローチは、NRCなど米国の基準に忠実に従っているが、動的歪み時効など米国配管材

一5一

における問題点を同様に抱えている点で、国産材料をベースとした我が国とは相違がある。

基本的な材料試験、解析評価等においては十分な力をつけているが、大型構造試験、実証試験についてはデータがまだ十分ではないとの印象を受けた。（鹿島）

2．2

セッション

6），，Crack

Shape

B
EvolutionofSurface

Pipes，，1。S．Hwang，J．H．Kim（Seoul

ソウル大学原子核工学科黄Flaws
National

under

FatigueloadingofAustenitic

U皿iv．）

一淳助教授がNi−Fe基超合金（インコロイ908）の疲労き

裂進展に伴う表面き裂形状の変化について講演した。インコロイ908は国際熱核融合実験

炉ITER（lntemationa1ThemonuclearExperimenta1R，eactor）の中央ソレノイド導管の主候補材料で、ITERの中央ソレノイドマグネットの寿命予測が研究の目的である。予測に際して、Paris則を以下で表示する。

深さ方向：da／dNニCA△Km
表面方向：dc／dN＝CB△Km
表面き裂平板の実験結果からCAとCBを決定し、アスペクト比a／cの変化と寿命を予測し、予測の精度を確認した。さらに、中央ソレノイド導管について、初期アスペクト比
をa／cニ0．1として、概念設計案（CDA）と工学設計案（EDA）に基づき、寿命予測を行い、アスペクト比の変化を考慮しないと寿命を過小評価すること、EDAの寿命は要求プラズマ
パルスサイクル（5，000回×安全係数2）を十分に上回ることを示した。
主な討論は以下のとおりである。（菊池教授）形状変化はビーチマークで検出できる。（黄助教授）ビーチマークはオーステナイト鋼に適用できない。（KAIST李順福助教授）き裂
閉口を考慮すべきである。（黄助教授）応力比Rニ0なので、き裂閉口の考慮は必要ない。

（矢川教授）ITERの設計に際しては、電磁力の動的効果を考慮する必要はないのか。（黄助教授）ひずみ速度は遅く、動的効果はない。

本研究は黄助教授の米国NBS（R．L．Tobler博士）留学中の仕事の延長のようである（韓国はITER．に参加していない）。

7〉，，Study（Science

on

the

E伍ect

Univ．of

東京理科大学

of

the

Crack

Length

on

the

Jlc

Value，，M．Kikuchi

Tokyo）

菊池正紀教授がA533B鋼とアルミニウム合金のCT試験片、3点曲げ試

験片およびCCT試験片の弾塑性破壊靭性JIcに及ぼす初期き裂長さ比a／Wの影響につい

一6一

て講演した。

見掛けのJIcはき裂先端の拘束（3軸応力）の減少に伴い増大し、き裂先端応力場の第2パラメータであるQ係数（HRR解からの偏差）とよい相関を示す。

主な討論は以下の通りである。（ソウル大学黄一淳助教授）JIcに及ぼす拘束の影響はASTMの試験方法
やASMEの基準で検討されているか。（菊池教授）検討されていないけれど、検討されるべきである。（小林教授）拘束の効果はJlcの試験方法に考慮する必要はない。構造部材の破壊評価には考慮すべきであり、ASME

Sec．XI規格委員会では破壊靭性に及ぼす拘束の影

響を検討している。

8〉，，Development

cialSur血ce

of

New

Z−factors

fbr

CrackinNuclearPipingsthe

EvahationoftheCircumfbren−

Y．H．Choi，Y。K．Chung，Y．W。RarkラJ．B。Lee（KINS），G．Wilkowski（Batte皿eMemorialInstitute）

韓国原子力安全技術院（KINS）の崔栄換博士が配管の周方向表面き裂の評価に使用する

新しいZ係数の開発について講演した。GE／EPRIの方法に基づき提案されたSC．TNP方法（薄肉管の表面き裂の解析方法）によってJ積分を解析し、J−R曲線を用いてモーメン

トー回転角関係を数値解析し、最大荷重を決定する。この方法によって、従来のASME規格の4つのZ係数（フェライト母材、フェライト溶接部、オーステナイト母材、オーステ
ナイト溶接部）に対して、新しいZ係数を開発し、48の配管破壊試験結果と比較してよく一致することを示し、ASME規格のZ係数の過大な裕度を削減できるとしている。

主な討論は以下の通りである。（小林教授）ASME規格のZ係数はRm／tの関数Aと管外径ODの積の一次式であるが、新しいZ係数はこれにAとODの二乗の積の二次項を加えている。多項式にすれば、実験結果とよく一致するのは当然で、さらに三次項、四次項が要求され、きりがない。日本の維持基準のZ係数のODの項は対数表示である。比較して欲しい（コメント）。（小林教授）破壊評価線図を用いてZ係数を決定する方が、簡単ではないか（コメント）。

本研究は崔博士の米国バッテル研究所（G．Wilkowski博士）留学中の仕事の延長のようである。

9），，New

Maintenance

Code

fbr

OperatingNuclearPower

PlantsinJapan（Dra銑）K．lida（JPEIC），H・Kobayashi（TIT），Y・Imam皿a（MHI），KHasegawa（Hitachi）

一7一

東京工業大学

小林英男教授が発電設備技術検査協会で作成された日本の維持基準原案の概要について講演した。維持基準の構成は、（1）非破壊検査（NDE）基準、（2）欠陥評
価基準、（3）補修／取替え基準であり、特に（1）と（2）についてASME規格、Sec．XI、Div．1の相違点が強調された。

主な討論は以下のとおりである。（KINS崔栄換博士）維持基準における材料劣化の扱いを説明して欲しい。（小林教授）日本の維持基準では材料劣化として圧力容器フェライト鋼の中性子照射脆化とステンレス鋳鋼の時効を取上げている。それ以外の材料劣化は日本の鋼ではない。中性子照射脆化の予測式は、米国R、．G．1．99Rev．2と特に溶接部に対して異なる。ステンレス鋳鋼の時効脆化は破壊評価線図を用いて評価するが、破壊評価線図は準備中である。（崔博士）維持基準のPDI（Performance

Demonstration

Initiative）への対応を説明して欲しい。（小林教授）日本の維持基準では、ASME規格Sec．XI

Appendix

V（企）とそこで要求される超音波探傷試験の資格試験に対応するプログラム（PDI）は採用していない。しかし、将来的には採用する見通しであり、発電設備技術検査協会に「UTによる欠陥検出性およびサイジング精度に関する確証試験」委員会が組織され、PDIへの対応を議論している。（質問者不明）維持基準におけるエロージョン／コロージョン（E／C）の扱いを説明して欲しい。（小林教授）現状の維持基準では、評価基準の中身は欠陥評価のみである。将来的には、疲労評価、E／C評価、破壊靭性評価、その他を含める予定で、現在、これらについて検討中である。

10），，ThePower

Development
Applicationin

Statusof

Mechanical

Component

Code

fbr

Nuclear

Kore紋，，N．H。Kim，J。S．Nah（KOPEC）

韓国電力技術株式会社の金南河技術士が韓国電力工業規格（KEPIC）のうちの機械的機器（Mechanical

Component）規格の現状と今後の展開について講演した。韓国は従来、原子力プラントの設計／製造規格はなく、外国規格を準用してきたが、種々の不都合が生じた。1995年11月に、機械、電気、構造、防火、品質保証の分野のKEPICが制定、出版された。KEPIC一機械はレベル1とレベルIIがあり、レベル1は原子力機器でASME規格
Sec．III

Div．1に相当し、レベルIIは非原子力機器と機器への共通要求（材料仕様書、非破壊試験、溶接認定）である。ASME規格を参照して、整備と拡大が継続されている。
圭な討論は以下のとおりである。（柴田勝之博士）配管防御規格はあるか。（金技術士）
ない。（小林）一8一

2．3

セッションC11〉，，Progress

gram

at

of

Component

JAERI，，by

K

Aging

and

Structural

Integrity

Resea■chPro．

Shibata（JAERI）

日本原子力研究所機器信頼性研究室の柴田室長より、原研における機器の経年化対策研
究の現状に関する総括的な紹介があった。主な項目は、（a）パイプの信頼性実証のための疲労試験とLBB試験、（b）経年化LWR機器の信頼性実証試験、であり、（b）の項目については、経年化の機構と予測手法の開発、経年化評価手法、安全性評価手法の各項目について現在進行中の研究テーマが説明された。また質問に答えて、研究の一部はUSNRCと共同で実施していることが報告された。

12）

DevelopingofExpert

Systemfbr

NuclearPipingIntegrity

by

Y．J．Kim，

M．W．Suh，C．S．Seok，H。KJun（SungKyuKwanUniv。），Y．W．Park，Y．H．Choi，J．B．：Lee（KINS）

成均館大学のKim教授が発表した。構造信頼性評価には多くの知識と経験が必要である
ことから、その補助のためにエキスパートシステム：NPiESを作製した。このシステムにはデータベースが蓄えられているが、古い材料のデータが不充分な場合、それを現在入手可能なデータから推論することができる。たとえば応カーひずみ関係のデータがない場合、これを降伏応力と引張り強さのデータから精度よく推論できる。今後はLBB評価へこのシステムを拡張したい。これに対し矢川教授から推論機構としてニューラルネットワーク
を利用することを推奨する意見が出された。

13〉

RecentProgressinKoreanNuclearP：LIMProgram，，byT．EJin，H：J．Choi，（KOPEC），1．S．Jeong，S．Y．Hong（KEPCO）

韓国電力会社のJin氏より発表があった。現在韓国では10基の原子炉が稼働中で、6基が建設中、さらに7基が計画中であり、PLIM（Plant

くなっている。PLIM
年までのPhase

Life

Management）の必要性が大き

Programは三期に分けて実施予定であり、現在は1993年から1996

IとしてFeasibility

Studyを実施中である。これは韓国の最初の商業炉であるKoriUIlit1を主な対象としている。今後詳細な評価を行うPhaseII（1997−2001），補修・交換を行うPhase

III（2001−2008）と、順次実施する予定である。

個人的な感想としては、Dr．Parkの言っていた、「半導体や車の分野では日韓は競争関係にあるが、原子力施設の安全性確保という分野では両者は協力関係にある」という言葉

一9一

が印象的であった。またKim教授のOpen

MindedDiscussion

率直な姿勢に感銘を受けた。（菊池）一10一の呼びかけなど、韓国側の

3

サイト見学報告
ワークショップの一環として5月9日（木）、月城（Walsong）原子力発電所の見学を行った。参加者と面会者は下記の通りである。
参加者：矢川（東大），小林（東工大〉，菊池（理科大），鹿島（電中研），柴田（原研），
馬郡（事務局）

面会者：機械部課長、ベイ（BaeYoung−Song）氏，機械部長，リー（LeeMyungBok）氏

ワークショップの翌日早朝，KINSが差回したマイクロバスで大田市を出発、見学先の月城（Walsong）原子力発電所に向かった。

天気は快晴で最高の見学日和であった。KINSのチャン（Chung

Yeon−Ki）氏が付添いとして同行した。

韓国には、古里（K・ri），月城（Walsong），ウルチン（Ulchin），霊光（Yonggwang）の4か所に原子力発電所があり、すべて韓国電力公社（KEPCO）が所有している。
月城発電所はその一つである。前日ワークショップが開催された大田市から約200km南
東の慶州市にあり、日本海に面している。大田市から約4時間の行程である。途中、慶州市内で昼食を取った後、さらに市内から1時間東に位置するサイトに向い、午後1時に到着した。

到着後、機械部課長のベイ（BaeYoung−Song）氏の案内を受けた。

最初に展示館を案内され、その後1〜4号機を外部から見学し、さらに機械部を訪問し部長のリー（Lee

Myung−Bok）氏と面談した。最後に、丘の上にある使用済み燃料の乾式貯蔵施設を見学した。

このサイトの従業員は約6，000人である。

月城のサイトでは、1号機が1983年から稼働している。現在、2，3，4号機が建設中である。1〜4号機ともAEC：L（カナダ）製のキャンドウ6型炉を採用していることが特徴である。電気出力は、1号機が68万kW，2，3，4号機が70万kWで，仕様もほとんど同じになっている。1号基は2ループタイプで、各ループには蒸気発生器が2基設置されている。

周知のように、キャンドウ炉は重水減速、重水冷却炉で，横型のカランドリア構造が特徴である。稼働中でも燃料交換できる特長がある。燃料交換は自動，半自動，手動等の方法で行われる等について燃料交換機器の説明を受けた。燃料交換は常時行われており、平均16バンドル／1日程度（20バンドル／チャンネル）である。
カランドリアを貫通するカランドリア管の数は380本である。カランドリア管の内側には、軽水炉の圧力容器に相当するZr合金製の圧力管が挿入されている。圧力管のISIはサンプリングで実施しており、超音波探傷と渦流探傷の他、肉厚測定も行われる。超音波探一11一

傷と渦流探傷の結果、インディケーションがある場合には次の年にも検査を実施することになっているそうである。定検期間は30〜40日程度でかなり効率的に行われているようである。

1号機以外については、2号機は今年の10月に燃料装荷が始まり、来年6月から営業運転を開始する予定。3号機は1998年、4号機は1999年と1年毎に運転開始する予定になっている。4号機は、5月にカランドリアが原子炉建屋内に搬入される予定になっている。当初3号機と4号機は計画されていなかったが立地が困難なため、3，4号機もこのサイトに建設されることになったそうである。これまで、運転員はAECLに派遣して教育訓練していたが、現在キャンドウ6型炉のシミュレータを建設しており、5月に完成するので、それ以後はサイト内での訓練が可能となる。

さらに、我が国と同様に韓国でも立地問題が深刻で、将来このサイトに5〜8号機の建設も検討しているとのことである。

最後に、乾式使用済み燃料貯蔵施設を見学した。乾式貯蔵施設は、サイトの裏山を整地した小高い場所に設置されている。この場所からサイ

ト全体とその後ろに日本海が見渡せる。左から1〜4号機が整然と設置されていた。

貯蔵施設には、直径3m×高さ5．5mのサイロが5体X12列の合計60体設置されている。1体のサイロで、60個X9段の使用済み燃料バスケットが保管できる。外側のコンク
リート容器の中に本体のステンレス鋼製容器が入る構造になっている。崩壊熱は、自然放熱により冷却され、ステンレス容器の設計温度は150℃である。
各ユニットにある使用済み燃料プールは10年分の貯蔵能力があるが、いずれ不足するので乾式貯蔵することにしたとのことである。6年間燃料プールで冷却後、この場所で乾式貯蔵されることになっている。乾式貯蔵を、いつまで続けるのかとの問いに対して、まだ見通しがなく永久保管の可能性もあるとのことであった。

日本では見られない、キャンドウ炉および乾式貯蔵施設を見学でき見聞を広めることができた。なお、時間の制限により原子炉の内部に立入れなかったのは残念であった。

韓国では現在、11基の原子力プラントが稼働し、その他7基が建設中、計画中が2基である。炉型はPWRとPHWR（加圧重水炉）であるが、輸入相手国は、米国，カナダフ
ランスで製造会社もWH：，CE，フラマトム，AECLと多様であり、これは国策として進められているようである。発電設備容量は、現在、950万kWとスエーデンについで世界第10位で、2002年までに建設中の7基が稼働するので、近い将来第9位となる。また、1995年の実績では原子力発電で電力全体の約50％を賄っており原子力への依存度はフ

ランスについで世界第2位で、我が国の30％を遥かに上回っている。すでに原子力発電大国になっている。

原子力発電の原価は、石油が38ウオン／kWh，原子力が20ウオン／kWhとのことで、我

一12一

が国の約1／3程度である。電力は国策会社の韓国電力公社（KEPCO：Korea

Electric

P・w−

erCorporation）が一社独占で供給しており、KAERI，KINS等の機関を含め、官民一体で原子力開発が進められているようである。我が国と同様に資源に乏しい韓国としては、コストの低い原子力発電への期待が極めて高いことが以上からうかがえる。

霊光発電所は黄海に面しているが、その他はすべて日本海側に位置している。霊光発電所を含め山陰・北九州地区からわずか200〜300km程の範囲に、我が国のプラント以外に建設中も含めると18基のプラントが在ることになる。原子力発電の安全性の観点からは、我が国だけでなく韓国の発電所発も含めて検討していく必要があることを強く感じた。

最後に、この見学の手配をして頂いた、KINSとKEPKOの関係者に御礼申しあげます。
（柴田）

一13一

4

成果と今後の課題
SMiRT会議などで日韓両国の機器や材料の健全性の研究者・技術者が顔を合わせることはこれまでにも数回あったが、今回のワークショップのような形で深く話し合ったり情報交換を行うといった機会はほとんどなかったものと思われる。韓国の技術レベルはかなり高く、両国の発表に対する討論も十分にかみ合った。このような意味で今回のワークショップは大成功であったと思われる。後で韓国側の代表者の一人に今回の感想をたずねたが、韓国側の印象としても今回のワークショップの意義は極めて高かったと評価していた。

今回は日韓の2国間のみでの会合であったが次回は台湾も含めてより強化した形での開催が予定されている。

参加者については今回は我が国からは中立研究者のみであったが次回は我が国で開催予定であり産業界、官界からの参加も期待したい。

このような試みは必ずしも短期的視点ではすぐに果実が得られるものではかもしれない
が、長期的観点から育てていくことが重要であると思われる。（矢川）

一14一

5

今後の計画
韓国側からDr．ParkとKim教授、日本側から矢川教授と菊池が出席して、今後の活動方

針に付いて協議し、以下の案を作成した。

5．1

今後の研究交流について

今後の研究交流のための機関として

ponents

EastemAssociationforIntegrityofNllclear

Com−（仮題）を結成する。当面は日本と韓国の二国からの代表で構成し将来的には台湾を初めとしてアジアの原子力発電所所有国に対象を広げる。各国は代表委員一名、幹事
一名、委員若干名を選ぶ。

Workshopを2年に一度開催する。またその際適当な施設の見学会も行う。次回は1998
年に日本で開催する。そのときは台湾にも参加を呼びかける。

5．2

講演論文の公表について

NuclearEnginering

design（NED）のSpecialVolumeとして印刷する方向で検討する。

今回のWorkshopの論文については日本、韓国それぞれで集めて、NEDの方式にしたがって校閲を行う。出版予定は1997年4月を目指す。

以上の案をWorkshOP終了後参加者全員に図り賛同を得た。

一15一

6

あとがき
今回のワークショップは、まさにアジア地区の関連研究者・技術者の対話のスタートであり、21世紀に向かってこの輪が大きく拡がっていくことが期待される。内外の関係諸機関ならびに各位の御理解と御援助を賜れれば幸いである。（矢川）

一16一

7

付録
これらはK

I

N

Sから提供された資料である。韓国における基準作りの現状、CAND

U型原子炉の詳細など、興味深いと思われるので全文を付録として木報告集に掲載する事
とした。

一17一

~!:~~:~

71 71~1 7~q~~:~)~~Part II: Intemational Workshop on the Integrity of Nuclear Components~'*=}~~!:~}~10*.}~~*~ 71 A O1

~~l~

4t-1

C ONTENT S
Session A

1

Application of Leak-Before-Break Using Piping Evaluation Diagram(PED)
Y.J.Yu*, K.S.Yoon, S.H.Par~ K.B.Park(~au),
Y.J.Kim(SungKyunKwan Univ.)
2

A Research Program for Dynamic Fracture Evaluation of Japanese
Carbon Steel Pipes
K.Kashima* (CR JEPI)

3

Effects of DynaJrnic Strain Aging on the Leak-Before-Break Analysis in

SA I 06 Gf.C Piping SteelI. S.Kim*, J.W.Kim(KAIST)
4

Probabilistic Fracture Mechanics Analvses ofNuclear Pressure Vessels

Under PTS Events
G.Yagawa* (Univ. ofTokyo), S.Yoshimura(CRIEPI),
M . Hirano(JAERI)

5

Improved Toughness of the SA 508 Class 3 Steel for Nuclear
Pressure Vessel Through the Steel-Making and Heat Treatment

J.T.Kim*, H.K.Kwor~ H.S. Chang (HANJUNG)

Session B

6

Crack Shape Evolution of Surface Flaws under Fatigue Loading ofAustenitic Pipes
I. S.Hwang*, J.H.Kim(Seoul National Univ.)

7

Study on the Effect of the Crack Length on the Jlc Value

M. Kikuchi*(Science Univ. ofTokyo)

8

Development ofNew Z-factors for the Evaluation of the Circunferencial
Surface Crack in Nuclear Pipings

Y.H.Choi*, Y.K.Chun~ Y.W.Park. J.B.Lee (KlNS),
G.Wilkowski(Battelle Memorial Institute)

9

Requirements for Pressure Boundary Integrity of Operating Nuclear
Plant-Japanese Standards
H. Kobayashi* (Tokyo Institute ofTechnology)

10

The Development Status of Mechanical Component Code for Nuclear
Power Application in Korea

N.H. Kinf, J.S.Nah(KOPEC)

Session C

11

Progress of Component Aging and Structural Integrity Research

Program at JAERl
K.Shibata* (JAERI)
12

Development ofExpert System for Nuclear Piping Integrity

Y.J.Kim*, M.W. Su~ C.S. Seo~ H.K.Jun(SungKyunKwan Univ.)
Y.W.Par~ Y.H. Choi, J.B. Lee(KlNS)
13

Recent Progress in Korean Nuclear PLIM Program
T.E. Jin*, H.J. Choi(KOPEC), I.S.Jeong, S.Y,Hong(KEPCO)

ll

SESSION A

1 Application of Leak-Before-Break Using Piping Evaluation

Diagram (PED)
Y.J. Yu*, K.S. Yoon, S.H. Park, K.B. Park (KAERI),

Y.J. Kim (SungKyunKwan Univ.)

Application of Leak-Beforc~Break Using Piping Evaluation Diagram(PED)

Y.J.Yu't, K.S.Yoon, S.H.Park, K.B.Park(Korea Atomic Energy Research Institute)

Y. J.Kim(SungKyunKwan University)

ABSTRACT
Plant specific data, such as pipe geometry, material properties and pipe loads, are required toapply LBB to a piping system. Thus, LBB evaluation can not be done until piping design androuting is completed. A simple method for evaluating LBB for piping systems during design
process is presented in this paper. This method produces a piping evaluation diagram, called
PED, for intennediate pipe locations and pipe-nozzle interface locations which defines the LBB
requirements to the piping designer for use during the design process and is independent ofpipe routing. This methodology will be used for the LBB evaluation of a new plant design suchas KNGR(Korean Next Generation Reactor).

INTRODUCTION
It was previously required that structures, systems, and components important to safety beappropriately protected against the dynamic effects of missiles, pipe whip, and discharging

fluids that may result from a postulated pipe break. Although the assumption of an
instantaneous double ended pipe break in large high energy lines provided a convenient way toenvelope the loads that might result from pipe rupture, it provided little or no relationship to theway which such pipes actually behave. This approach led to the need for substantial protective

measures to guard against the consequence of such postulated breaks. These protective
measures are expensive to build and maintain, and lead to a potential degradation of plantsafety. The placement of pipe whip restraints degrades plant safety if thermal expansion isrestricted and when the accessibility for and effectiveness of inservice inspection is reduced.

The research of the last decade on elastic-plastic fracture mechanics has led to a means tojustify a more reasonable altemative: Leak-Before-Break(LBB). The fundamental premise of
LBB is that the materials used in nuclear power plant piping are sufficiently tough that even alarge throughwall crack, which could result in coolant leak rates well in excess of those
detectable by present leak detection systems, would remain stable and would not result in a
double-ended guillotine break under maximum loading conditions. Thus the dynamic effects ofpostulated ruptures in the piping system are eliminated from the design basis when the piping
system is shown to meet all the criteria for the application of leak before break presented in

NUREG 1061, Volume 3[1] and Standard Review Plan 3.6.3[2].
Since the mid-1980's, the LBB technology has been applied extensively to a high energy
piping system[3] in existing plants. However, there are differences between the application of
LBB to an existing plant and a new plant design. In this paper a simple approach is introduced

which is intended to use for application of LBB to a new plant design such as KNGR. Thisapproach is based on LBB PED developed by Fabi[4] and extended to the pipe-nozzle interface
location in this paper.

1-1

CURRENT REGULATORY CRITERIANUREG-l061 sets forth stringent criteria to be applied to each piping system to determine if

LBB is a viable altemative to postulating a DEGB (Double-Ended Guillotine Break). "TheLBB approach should not be considered applicable to high energy fluid system piping, or
portions thereof, that operating experience has indicated particular susceptibility to failure fromthe effects of corrosion(e.g., intergranular stress corrosion cracking), water hammer or low andhigh cycle (i.e., thermal, mechanical) fatigue". To the extent that fatigue may exist in aparticular piping system, it must be considered when performing a LBB evaluation. For thosepiping systems to which LBB is applied, the following criteria must be met:
(1) A Ieak detection system is required that is capable of detecting a leakage rate (less than

1.0 gallon per minute for the primary system as a requirement of Regulatory Guide 1.45),

to which NUREG-1061 applies a margin of 10.
(2) Throughwall cracks which are large enough to leak 10times the detection capability (i.e.
lO gallon per minute for primary system piping) must be stable for I~times the sum of

the normal operation and safe shutdown earthquake(SSE) Ioads.
(3) Throughwall cracks twice as long as a crack which leaks 10 times the detection
capability must be stable under the sum of normal operation and SSE Ioads.

In developing PED, the appropriate margins required in the current regulatory criteria are
included.

REQUIRED INPUT FOR ANALYSIS
PED defines the LBB requirement to the piping designer for use during the design process. In

order to define the LBB requirements, several sets of LBB analyses are performed for eachdifferent pipe size and material considered in the LBB application. The following are the input
for the analyses to develop the PED.

Material Pro erties
Previous work[5] by ABB-CE has shown that a conservative bounding analysis results whenthe material stress-strain properties of the base metal (10wer yield) and the fracture propertiesof the weld metal (lower toughness) are used for the entire structure. Industry data [6,7,8] arereviewed to establish the lower bound stress-strain and J-R curves for each material. For boththe final design and as-built configurations, the actual material properties used for piping
systems subject to LBB shall be reviewed to confirm the application of LBB.

A Iied Loadin
There are two aspects to the loading for Leak-Before-Break. First, the normal operating(NOP) Ioad is used in determining the detectable leakage crack length. The NOP Ioading
includes the system pressure and the thermal expansion piping moment. The NOP Ioad is
generally small enough so there is no concern for stability with this loading alone. Two NOP
loads which cover the entire range of possiblc loading for the line under consideration isassumed. The second aspect of the loading is the LBB Ioadings, which is the NOP plus
additional loading(s) such as safe shutdown earthquake(SSE) Ioading and stratified flow(SF)loading. The LBB Ioading provides an additional piping design requirement which has to be
met.

1-2

Leak Detection Ca abilit
The basic premise of the Leak-Before-Break concept in piping is that a flaw will be detected
via loss of fluid prior to the failure of the pipe. This requires a detectable leak rate inconjunction with stress-strain curve for pipe material. It is assumed that the plant Is capable ofdetecting a leakage of one gallon per minute with an hour based on the requirements ofRegulatory Guide 1.45[9]-LBB ANALYSIS FOR PED
Crack stabiiit evaluation
The methodology to evaluate the stability of through-wall cracks requires knowledge of theapplied loads, a leakage crack size, and the material properties. Finite element models of threedifferent crack lengths were developed for the leakage crack and twice the leakage crack to
consider the crack extension. One set of model having crack sizes of a, a - 5 and a + 8 at

normal operation loads was used to demonstrate safety margin on the loads. "a" rs thedetectable leakage crack length, and " 8 " is a small increase or decrease in "a". Another set ofmodels, having crack lengths 2a, 2a+6, 2a-6 were used to demonstrate the margins on
crack size. The crack lengths are input to the detailed stability analysis of the through wall
cracks in piping systems. The FEA is carried out for the estimated leakage crack size and twicethat length.

Two planes of symmetry are used for pipe model to minimize the size of FE models.
Therefore, each model represents one quarter of the pipe as shown in Figure 1. The model for
location near the nozzle is shown in Figure 2. Here one plane of symmetry is used to minimize
the model size, meaning that one half of the pipe-nozzle is modeled.

/

,*

/

'

vL

Figure 1. Finite Element Mesh for Pipe

Figure 2. Finite Element Mesh for Pipe-Nozzle

Model

Model
J-Inte ral

To evaluate the intensity of the stress field near a crack tip for the elastic-plastic problem, theJ-jntegral parameter is used. The J-integral parameter is related to the energy release rate at thecrack tip. The method used to evaluate the J-integral is virtual crack extension method, where;)( dE/da) dE is the
J=(l/tandchange m the stram energy release rate tor a vrrtual crack
extension, da. To calculate dE , small virtual displacements of the nodes in elements near the

1-3crack tip are applied in the direction of the crack extension.

The J-jntegral is determined by FEA for pressure, normal operation, and SSE Ioading for threedifferent crack length for each geometric model. To evaluate the margin on crack length, Jintegral is evaluated for the applied loads for 2 times the detectable leakage crack size. For themargin on loads evaluation the J-integral is also evaluated \~ times the applied loads for thedetectable leakage crack size.

Stabilit Evaluation
The stability of the cracked pipe is assessed by comparing the J-integral value due to theapplied loads on the pipe to the material crack resistance. The stability criteria employed for
ductile crack extension is:
Crack Stability is assured if:

Japplied material

(1)

and

dJ

material

da

( 2)

da

The slope of the J-integral vs. a curve for each location is obtained by fitting a polynomial
curve through the J-integral values. By plotting J-integral vs. dj/da which is called a J-T
diagram, both parts of the slability criterion above can be evaluated simultaneously.

Develo ment of J-T dia ram
The J-jntegral vs. dJ/daboth crack length a and 2a .

curve is called a J-T diagram.
This is done as follows;Th e J-T diagram is developed forl. The J-T diagram due to applied load:a.

From FEA, determine applied J for cracks a, a + 6 and a - 5 as well as 2a,b.

This gives J as a function of crack length a. Thus dJ/ daand 2a can be determined.

c.

Polynomial is fit to J where:

2a+6, and 2a-6 for defined loading.
in the vicinity of aJ(a) = cla2 + c2a + c3

(3)

dJ

(4)

thus

da = 2cla + c2

The J-T diagram due to material properties

2.

b.

Material J-R curve defines J for different crack lengths
Power law curve is fit to J-R data curvel-4

J = Cl (Aa)c2 (5)thus,= CIC2 (~La) C2-ldaThis process is outlined in Figure 3. Once the J-T diagram is developed, the load that causes
instability is defined, for each of the highest stressed location in the pipe. At the point ofintersection of the loading curve and material curve, the following holds true:

J = J (7)applied mate'ia/

anddJ"pplied dJ_ ~ate'ial (8)dadaThis is the point above which unstable crack growth will occur for the given load. Thisinstability load is compared to the actual applied load at that point. If the LBB applied load isless than the instability load, then the point passes the LBB criteria. If LBB applied load isgreater than the instability load then the point fails the LBB criteria.
Instability Load+

Typieal Loading Typieal Materlal
Figure 3. Stability Evaluation

CONSTRUCTING A PED
PED for intermediate location
The procedure for constructing a PED for Intermediate pipe location was developed by Fabi[4]. The procedure is as follows;

For detectable leakage length, the LBB procedure requires a margin on loads. Thus the

maximum load equals to NOP+SSE.
Mmax(i) = I~( Mssc(i) + M,vop(i)) (9)This can be solved for the allowable SSE Ioading ;MssE(1) Mmax(i) / J~ - MNop(j) (10)where, i = each NOP condition for analysis of leakage crack length 'a'
For tvvice the detectable leakage length, the LBB procedure requires an additional margin on

l-5crack length. Again setting the maximum load equals NOP+SSE.
Mmax(i) = MssE(i) + MNOP(,)(11)

Solving for the allowable SSE Ioading ;MSSE(1) Mmax(i) ~ MNOP(i)( 1 2)

where, i = each NOP condition for analysis of twice leakage crack length '2a'A typical PED for a GTAW weld in a SA312 Type 347 stainless steel 12 inch schedule 160
class I pipe is shown in Figure 4.

400
3 50

300

A

e

Z 2 50
~o

J 200
w
o'
~
cJ'

:

ISO

2A

~
,OO

so

o
o

20

40NOP
CO
80 t OO , 20Lo•dt ,KN-m'

Figure 4. PED for 304 mm diameter stainless steel pipe (SA312 Type 347)

PED for i e-nozzle interface location terminal ends
Typical LBB (Leak-Before-Break) analysis is performed for the highest stress location foreach different type of material in the high energy pipe line. In most cases, the highest stressoccurs at the nozzle and pipe interface location at the terminal end. The standard finite element
analysis approaches to calculate J-integral values at the crack tip utilizes symmetry conditions

when modeling near the nozzle as well as away from the nozzle region to minimize the model
size and simplify the calculation of J-integral values at the crack tip. A factor of two is typicallyapplied to the J-integral value to account for symmetric conditions. The stiffness of the residual

piping system and non-symmetries of geometry along with different material for the nozzle,
safe end and pipe are usually omitted in current LBB methodology. A study, done by Yu et.al[11], shows that this simplified analysis can lead to conservative results especially for small

diameter pipes where the asymmetry of the pipe-nozzle interface is ignored.
In this papcr, a PED for pipe-nozzle interface location is developed to consider the effects of

non-symmetries due to geometry and matcrial at the pipe-nozzle interface. The procedure toconstruct the PED for the pipe-nozzle interface location is basically the same as that for anintermediate pipe location. However, the calculation of a detectable crack length for the pipe-l-6

nozzle interface location is somewhat different with that for the intermediate pipe location.
Consideration of the nozzle requires an iterative procedure to find an appropriate crack length
which leaks at 10 gpm and employs both the finite element model used for the crack stability
analysis and the PICEP code[10]. Since the stiffness of the nozzle is included in the stability
analysis it must also be included in the leakage calculation. The procedure uses the PICEP
program as a flow calculator for a given assumed crack length and calculated area from thepipe-nozzle finite element model. This procedure is as follows[1 1]:
Step 1: Assume a flaw length in FE model.
Step 2: Apply normal operating load to the FE model and calculate the crack opening area.

Using PICEP with the same length flaw vary the applied moment until the area isthe same area as calculated with the FE model.
Step 4: If the PICEP flow is greater than 10 gpm the crack length is decreased - go to step

Step 3:1.

If the PICEP flow is less than 10 gpm the crack length is increased - go to step 1.

If PICEP flow is 10 gpm - STOP.
The final step establishes the pipe-nozzle interface crack length which leaks at the detectability

limit(with margin) of 10 gpm.

Figure 5 presents a PED for the nozzle-pipe interface location and the intermediate pipe
locatron where "A" and "2A" are the detectable leakage crack length and twice the leakage
crack length, respectively.
400. OOl

350.00NOZZLE 'A'
300.00PIPE 'A'

E,

Z!
,,

,
3
tU
CO
C"

~
l)

A-

' IL~ '

250.00INOZZLE '2A'

200, ooJbA.

1 50.00

P,PE '2A'~
1 OO. OO

L'

50.00
0.00

o

20 60
1Lo•d•OOtKNlnj
1 20
NOP
80

40

Figure 5. Comparison of pipe and nozzle-pipe PED

LBB EVALUATIONUSING PED
For the highest stressed location for each different type of material in the line, the NOP and
SSE Ioading are plotted on Figure 6. If this point falls below both curves in Figure 6 (loading
represented by point P), the line passed LBB with appropriate safety margin already included in
the plot. If the highest stressed points fall above either curve in Figure 6(10ading represented bypoints X,Y and Z), the line fails LBB. Then the piping designer can modify the piping design in

ordcr to pass LBB.

1-7

~,ARaN a, APPLYtro LOAO,A)

,

/

eX
UARaN ON CRACK stzE'2A)e Y
SSE LOAOeZ\

N~ LOAD

Figure 6. Typical piping evaluation diagram (PED)DISCUSSIONS AND CONCLUSIONS
An approach is presented which is suitable for application of LBB for a new plant design suchas KNGR. The advantage of this approach is that the piping designer can quickly iterate to a
design which satisfies both the ASME code and LBB requirement.

However, LBB evaluation using PED can be very cbnservative for some cases because PED
uses industrial lower bound material data and loads calculated from a elastic analysis based on
the piping system being uncracked. A modified PED which can consider the effects of applying
load at the cracked pipe section due to the residual compliance of the piping system are now

under developing at KAERI and SungKyunKwan University.

REFERENCES
NUREG-1061,VOI 3 "Evaluatron of Potentral for Prpe Break "USNRC Draft Standard Revrew Plan 3 6.3, "Leak-Before-Break Evaluation
Procedure," March, 1987.
[3] K. Wichman, J.Tsao, M. Mayfield, 1996, "LBB Application in the U.S.A. Operating

[1]

[2]

and Advanced Reactors " NEJVCSNI/R(95)18 Volume I ppl3-19.
[4] R.J. Fabi. D.A. Peck, 1994, "Leak Before Break Piping Evaluation Diagram, "Presented at Thel994 PVP Conference, PVP-Vol.283, pp. 111-115.
[5]
EPRI Report NP-5057, "Analysis of Cracked Pipe Weldments", February 1987.
[6]
A.L.Hiser and G.M.Callahan, "A User's Guide to the NRC's Piping Fracture

Mechanics Data Base (PIFRAC)", NUREG/CR-4894 (MEA-2210), Materials
Engineering Associates, Inc.. Lanham, MD, May 1987.
[7] Y.J.Kim et. al, 1993, " Fracture Properties Evaluation of BOP Piping Systems for the
Yonggwang 3/4 Nuclear Generating Station," Trans. of the 12th SMiRT Conference,
Vol G, pp. 435-440.

[8] SungKyunKwan Univ.,1995, "LBB Test Evaluation of Primary Piping Weld Materials
for Ulchin 3/4 Nuclear Power Plants(RCIJPZR Surge Line Field Fabrication),"

SKKU/SAL-9502.
[9] USNRC Regulatory Guide 1.45, "Reactor Coolant Boundary Leakage Detection
Systems", May 1973.
[lO] EPRI Report NP-3596-SR Rev.1, "PICEP: Pipe Crack Evaluation Computer
Program", Electric Power Research Institute, July 1987

[11] Y.J.Yu, S.H.Park, G.H.Sohn, Y.J.Kim and W. Urko, 1995, " Application of LBB to a
Nozzle-Pipe Interface", NENCSNI/R(95)18, Volume l, pp83-91.

1-8

2

A Research Program for Dynamic Fracture Evaluationof Ja panese Carbon Steel Pi pes

K. Kashima* (CRIEPI)Presented at International Workshop on the Integrity of Nuclear Components

May 8, 1996, Kl:NS, Korea

A Research Program for
Dynamic Fracture Evaluation of Japanese Carbon Steel Pipes
K.Kashim~ N.Miura
Central Research Institute of Electric Power Industry (CRIEPI), Tokyo, JapanS.Kanno. K.Miyazakj4 M.Ishiwata and N.Gotoh
Hitachi Ltd. , lbaraki, Japan

Abstract
The research program was developed to investigate the dynarnic load

effect on Japanese carbon steel STS410 pipe. The program comprises
material tests, pipe fracture tests and development of estimation scheme.

Material property tests showed that the flow stress was nearly constantor slightly increased with strain rate. Pipe tests showed that fracture load
was nearly predicted by the net-section collapse criterion for quasi-static anddynanaic loading. A significant dynamic effect was not observed for
STS410 carbon steel piping. Crack growth was well formulated by using
J-integral parameter for low cycle fatigue. Combining the crack growth
behavior and unstable fracture criterion, an estimation scheme was newly
developed and validated for constant cyclic loading conditions.

2-ll. Intrrduction
Leak-before-break (LBB) is one of the potential concepts to evaluate the piping integrity

for light water reactors. Many LBB researches in the govemment, research institutes,

utilities and manufacturers have been conducted to demonstrate the validity of thisconcept and apply it to the piping designs in nuclear power plants.

In Japan, the LBB research programs were initiated at NUPEC (Nuclear PowerEngineering Corporation) for stainless steel piping in 1977 and carbon steel piping in

1 985 under the sponsorship of MITI [ I .2]. In these programs, fundamental material

tests, full-scale pipe fracture tests and LBB assessment were conducted. Many
researchers ~)ined the programs and CRIEPI was responsible to evaluate unstable pipe

fracture using fracture mechanics approach. JAERI also developed another LBB
program including the evaluations of fatigue crack growth, unstable pipe fracture andpipe whip restraints [3]• In 1987. USNRC organized an int
group, "IPIRG" under the funding from 9 countries including USA (USNRC, EPRI), UK(Nuciear Electric), France (CEA), Italy (ENEA), Sweden (SPI). Switzerland (HSK),

Canada (AECB), Taiwan (TPC), Japan (CRIEPI). The main ob~ctive of the IPIRG
program was to study the effect of seismic (dynamic/cyclic) Ioading on pipe fracture.
The program has bcen in Phase-2 since 1 992 and wm be terminated in April, 1996.

Tue IPlRG results showed the dynamic and cyclic loading effects on pipe fracture forsome materials. This trend was signlfrcantly observed in AI06 Grade B, typical carbon
steel in USA. In 1 991,CRIEPI and Hitachi initiated a collaborative research program toinvestigate the effects of dynamic and cyclic loading on the fracture of Japanese carbon

steel piping [l~6]. This paper presents the summary of this program and the major
fmdings.

2. Program Milestones
The 6-years research program was started in 1 99 1 as the joint study of CRIEPI with
Hitachi and will be terminated in 1 996. The program includes the following milestones:(1)

Conduct material tests under quasi-static and dynamic loading conditions.

(2)

Conduct pipe fracture tests under monotonic and cyclic loading conditions,(3)

Deveiop an effective evaluation scheme for highly low cycle fatigue crack growth and( 4)

Evaluate the dynamic and cyclic loading effects on Japanese carbon steel piping.

2-2

The schedule of these mnestones is shown in Table I .

3. Material Property Tests
Material tests were conducted to obtain the tensile properties f'or base metal andsubmerged-arc weld metal of STS410, one of the typical carbon steels used in Japanese

nuclear piping. The che~tical compositions of the materials are tabulatcd in Tabie 2.

Tensile tests were conducted at room temperature and high temperature (288C or 300O
under quasi-static strain ratc (1O~4 Iscc) to dynarnic strain rates ( 101 /scc). Test
specimens were taken f'rom the pipe in the axial direction. Figure I shows thc gcomctry

and dimensions of the round-bar specimens.

Figures 2 (a) and (b) show the rcsults of tensue tests tbr base mctai and wcld metai. For

both materials, yicld stress (0.2% proof stress) and ultirnate strcss increascd with thc
incrcase of strain ratc at room tcmpcraturc. At high temperature, yield stress increascd

with strain ratc and ultimatc strcss decreased with strain rate. These results clearly

showed the dependency of the tensile properties on strain rate. However, it should be
noted that the flow stress, a major dominating parameter for plastic couapse of ductile
material, did not show a signifllcant dynamic effect on strain rate at high temperature
because it was nearly constant or slightly increased with strain rate when it was dct~incd as
the mean value of yieid stress and ultimate stress.

4. Pipe Fracture Tests
Specimens fbr pipe fracture tcsts werc circumferentially through-wall or surf'acc-crackcd

pipcs of STS410 basc metal and wcldcd joint with 6-inch diamcter. Figurc 3 shows thc
gcometry and ditnensions of thc spccimen. Initial notch with the total angle of 3O or 60

degree was introduccd by clcctrical discharge machining afier mechanical machining.
Figure 4 iliustrates the apparatus used f'or the pipe fracture tests at room temperaturc andhigh temperatures. In high temperaturc tests, the specimens werc heated upto 265C to285C by forccd circulation of hot air.

Thc pipc spccimcns wcrc sub~ctcd to tbur-point bcnding undcr thc tbllowing loading
types:

Type ( I ): Monotonic loading tests

2-3

Type (2): Constant amphtude cyclic loading tests
Type (3): Incremental amputude cyclic loading tests

Type (4): Random cyclic loading tests

Figure 5 illustrates these loading types. Strain rates wcre quasi-static except for Type( I ), where both quasi-static (0.05 mm/sec) and dynamic (5 mm/sec) displacement rate
were applied. The effect of loading rate on monotonic pipe fracture was investigated in

Type ( l) test. In Type (2) tests, the loading wave was an altemating triangle of theconstant load amplitude with the frequency of 0.1 Hz. The numbcrs of cycles at pipe
f'ailure were measured for pipe specimens sub~ctcd to thc constant load amputudes equal

to 5O-9O% ofthe plastic col]apse load. Typc (3) and (4) dcal with the complex loading
types of increasing load amplitude and random load.

More than 40 pipe testswere conducted in Type( I ) to (4). Figure 6 shows the resu]ts ofType (1) tests as the relationship between normalized maximum load and displacement

rate. The normalized maximum load was defmed as the ratio ofthe measured maximum
load to net-section collapse load and the displacement rate was measured at the loading

point on pipe specimen. These results indicates that the normalized maximum load [snearly constant or slightly increasing with strain rate. They mean that fracture load isneariy predicted by the net-section collapse criterion for quasi-static and dynamic loading

and STS410 does not show the signiflcant dynamic effect as was observed for AI06
carbon steel [7].

5. Estimation Scheme to Predict Failure LifeAn estimation scheme was developed to evaluate a series of fracture behavior from thecyclic cTack growth to unstable fracture for clrcumferentiauy cracked pipe'sub~cted toconstant amplitude cyclic bending.

For crack growth under very low cycle fatigue condition, the elastic-plastic stressconditions near the crack tip should be taken into account. Therefore, the elastic-plastic

fracture mechanics parameter, ~L J, was introduced to the crack growth rate equation.

instead of thc prevailing elastic parameter, A K. A new equation was developed to
calculatc ~ J for circumferential crack using thc Jmax (monotonic J-jntegral) and crack

length [4]. Jmax can be easlly calculated using thc GE-EPRI handbook [8]. The
calculation steps are as fouows:2-4( I ) Calculate Jmax for initial crack length (ao) in the t~Irst loading cycle.

(2) Translate Jmax into A J.
(3) Calculate cycbc crack growth rate (d~dN) from "(da/dN)- ~L J" relation.
(4) Update the crack length (a) by adding the crack growth (da).
(5) Decide whether pipe will fracture or not by net-section couapse criterion.
(6) If no fracture wilJ result, repeat the above steps (1) to (5) in the next loading cycle.

Figure 7 shows the comparison of the experimental and predicted relationship between

the crack extension and number of cycles for 6 cases at room temperature and hightemperature. The predicted behavior is in good agreement with the test results.

6. Conclusions
The research program was developed to investigate the effect of dynamic and cyclicload on Japanese carbon steel STS410 pipe. The program comprises material tests, pipe
fracture tests and development of estimation scheme.

l ) Material property tests showed that tensile stress and yield stress were dependent on
strain rate but the flow stress was near]y constant or slightly increased with strain

2) Pipe tests showed that fracture load was nearly predicted by the net-section collapse

criterion for both quasi-static and dynamic loading. A signlflcant dynamic effectwas not observed for STS410 carbon steel.

3) Cyclic crack growth was formulated by introducing J
Combining the cycuc crack growth equation and unstable fracture criterion, a new

estirnation scheme was developed and vaudated for constant cyclic loading
condit ions.

References[ 1 1 NUPEC. "Proving Test on Stress Corrosion Cracking in BWR",NUPEC.(1984).
[2] NUPEC. "Provhrg Test on the Integrity of Carbon Steel Piping in LWRs", NUPEC,( 1 989).

[3] K.Shibata, et al.. "Results of Reliability Test Program on Light Water Reactor Piping",
Nuclear Engineering and Design, 153, ( 1994), 7 1 .

2-5

[4] N.Miura, et al., "Low Cycle Fatigue and Ductile Fracture for Japanese Carbon Steel

Plplng under Dynarruc Loadmg" Nuclear Engmeermg and Design, 153, ( 1 994), 57.

[51 N.Miura, et al., "Low Cyclc Fatiguc Crack Growth and Ductile Fracture under
DynamiclCyclic Loading f'or Japancse Carbon Steel Piping ,Part-2:Analytical Study",

ASME PVP 266,( 1983), 175.
[6] N.Miura, et al., "Application of Cyclic J-integral to Low Cycle Fatigue Crack Growthof Japanese Carbon Steel Pipc", Proceedings of LBB95, Lyon, France, October, 1 995.

[7] R.A.Schrnidt, G.M.Wilkowski. "The Intemational Piping Integrity Research Group(lPIRG) Program: An Overvicw", SMIRT- I I , Tokyo, Vol.G, ( 1 991 ), 177.

[8j V.Kumar, et al. Purther Developments m Elastrc Plasuc Analysls" EPRI NP 5596
( 1 988).

Table lITEM(1 ) Material Tests(2) Pipe Tests(3) Estimation Schemc

Program milestones

l 994 1 9%1991

l 992 1993e BasceBase eweld( RT)(rrr) (RTIHT)

e Base(RT)

OBasc eweld eweld(HTD( H~ (RT)e BasceBasc OBase eweld ewcad( RT)(RT) (HT)l 995(RT) (Pnr )o(4) Evaluaticn of Dynamic
Cyclic Errect fer STS410

Table 2

Chemical compositions of materials(wt~,)

2-6

,r,

,1

dmj~
Figure lGeometry and dirnensions of tensile specimen

700
600
,~

~ 500

~400
~5

300

200
l O 'S

l 0-4

l 0'3 t 0'2 - l

l OO

tO

lOl

(a) Base metal

St!ain Rate ~ Is

700
600

~

~ 500

~400
~;

300
200
l O 'St 0~ t 0'3 1 0-2 ~ I O t
Suain Rate ~ /s (b) Weld metal
lO '

I Oo

Figure 2 Results of tensile tests for vanous stram rates

2-7

wcld
~~

' ~ d~fcct ~ l~

s
=-

weld
E
=

2

outcr diarnetcr :l 14.3 mm

inncr diametcr :pipe thiCkncss :97.1 mm
8.6 mmo32.7c•~•- •t••••~l~~ '4

undcr 0.3
i

r

f

h-w l

Ceometry and dimensions of notched pipe specimen

Figure 3

supportino roiler

Figure 4Schernatic view of plpe fracture tests apparatus

'1:,

'~

'~

=
O
C_
O

~

U
CQ
'l'

1~

une

Ime

m n t m I in
Type( I )m n l.

Type(3)

Type(2)

Figure 5li lLoading types for pipe fracture tests

2-8

a in

Imc

,1

n mc li lType(4)

in

I .4

~
Qb

~
,y,:

I .2

G:

lo

'1

S

l .O

E

.~

~ 0.8
='

o
::!

-

0.6

,g
EL

o

Z

0.4
O. l

l

Displacement Rate A. mmls

Relationship between normallzed maximum load and displacement rate

Figure 6

50
4S

40

E
E 35

~ 30

=
o 25
o 20
u,=

~
~e

V

15

(J

lO

csL

5

O
No. of Cycles N

Figure 7

Relationship between crack extension and number of loading cycles

2-9

3

Effects of Dynamic Strain Aging on the
Analysis in SA 106 Gr.C Piping Steel

I.S. Kim*, J.W. Kim (KAIST)

Leak - Brefore - BreakEffects of Dynamic Strain Aging on the Leak-Before-Break Analysis
in SAI06 Gr.C Piping SteelJin- Weon Kim and In-Sup Kim

Department ofNuclear Engineering
Korea Advanced Institute ofScience and Technolo~l

Abstract

The characteristics of dynamic strain aging (DSA) on material properties used in leak-before-

break (LBB) analysis were discussed. Using these material data, the effect of DSA on the LBB

analysis was estimated through the evaluation of leakage-size-crack and flaw stability in
SAI06 Gr.C piping steel. Also, the results were represented as a form of "LBB allowable load

wmdow" In the DSA temperature region, the leakage-size-crack length was smaller than thatat other temperatures and it increased with increasing tensile strain rate. In the results of flaw

stability analysis, the lowest instability load appeared at the temperature corresponding tominimum J-R curve which was caused by DSA. The instability load near the plant operating
temperature depended on the loading rate of J-R data, and decreased with increasing tensile
strain rate. These are due to the strain hardening characteristic and strain rate sensitivity of

DSA. In the "LBB allowable load window", LBB allowable region was the narrowest at thetemperature and loading conditions where DSA occurs.

1. Introduction

The leak-before-break (LBB) concept has been applied to design for high energy
piping in nuclear power plant. The stress-strain curve and J-R curve of material are used for

LBB analysis, and they have direct influence on the evaluation of LBB conditions [l].
Accordingly, it can be expected that the characteristics of dynamic strain aging (DSA) inmaterial may influence the results of LBB analysis, since the occurrence of DSA depends on
temperature and deformation rate and changes mechanical properties.

3-lAlthough a number of studies on DSA related to LBB were reported [2,3,4], most ofthem focused on the crack jump at quasi-static loading rate and on the loss of fracture
toughness under dynamic loading at the normal plant operating temperature. The effect of
DSA on the results of LBB assessment has not been quantitatively estimated. In this study,

therefore, the effect of DSA on the LBB analysis was estimated through the evaluations ofleakage-size-crack and flaw stability for simple case using material properties obtained fromsystematic tensile and J-R tests.

2. Matenal Propertres for LBB Analysrs

The calculations of leakage-size-crack length and flaw stability in LBB analysis
require the stress-strain curve and J-R curve [5]. The tensile and J-R tests were carried out

under various temperature and loading conditions to illustrate the characteristics of DSA in

SAI06 Gr. C steel which has been used for main steam line piping in nuclear power plant. In
the results, the effect of DSA on the material properties for LBB analysis was discussed.

2.1 Yield Stress

As shown in Fig. I , the 0.20/0 c~ys generally decreased with increase in temperature

although there was a plateau at a certain range of temperature. The temperature at which aplateau appeared, shifted to higher temperatures with increasing strain rate and coincided withserrated flow region. The oly, at temperatures lower than 2000C showed the usual positive

strain rate dependence. At temperature range of plateau associated with serrated flow,
however, positive strain rate dependence of ay, disappeared and reappeared at 4000C.

It is noted that ay, is affected by static strain aging (SSA) rather than DSA [6].
However, many studies observed a plateau or a small peak of c~y, [7], and a slight strain ratesensitivity [8] in the temperature range of DSA. Therefore, the decrease in c:y, is due to hightemperature softening, whereas the occurrence of plateau and disappearance of usual positive
strain rate sensitivity reflect the effect of DSA over the high temperature soitening.

3-2

2.2 Ramberg-Osgood Parameters
Ramberg-Osgood relation is usually used as power-law hardening relationship and it's

parameters n and oc are important input data in the LBB assessment. Generally, RambergOsgood relation is represented as following equation.

a

[n

+a
ao

(1)

Co (TO
where a and 6 are true stress and true strain, o:o is reference stress that is usually equal to yieldstress, g0=a/E, n is strain hardening exponent, and cc is dimensionless constant. The
Ramberg-Osgood parameters are strongly dependent on strain range of data fitted into Eq. ( I )[1,9]. In the present study, stress-strain data between I .050:o and 100/0 strain were selected toscreen out points on the yield plateau.

The parameter n is plotted against testing temperatures at various strain rates in Fig. 2.

In comparing dependence of n on temperature and strain rate with that of ol~ts' the temperature
region where n is smaller than that at room temperature (RT), is consistent with ol~ts hardening

region where DSA operates. In this temperature region, also, strain rate dependence of n
showed inverse trend in comparison with other temperatures, in the same manner as the strain
rate dependence of a;~ts' These suggest that the parameter n is directly affected by DSA, andthe decrease in the value of n and positive strain rate dependence at the normal plant operating

temperature range are caused by DSA.
The parameter cc varied with temperature in the similar manner as the ays as shown in
Fig. 3 , although the dependence of temperature and strain rate is less clear than that in ays' Tuevalue of cc decreased gradually with increase in temperature and showed positive strain ratedependence except the region of temperature in which the plateau appears. The dimensionlessconstant cc is related to both o~ys and strain hardening of material. In particular, cc is sensitive tothe value of c~ys [ I O]. This is consistent with the present observation. From the characteristics

of oc, it can be concluded that the appearance of plateau and slight strain rate sensitivity near3-3

300"C are attributable to DSA as discussed in c~y.. However, the influence of DSA on thisparameter is less obvious than that on the parameter n.

2.3 J-R Curves

Fig. 4 exhibits the J-R curves of material at different temperatures for load-line

displacement rates of 0.4 and 4.0 mm/min. The dependence of the fracture resistance on
temperature and loading rate was apparent. As shown in the Fig. 4, it was noted that as thetemperature increased from ambient, the J-R curves exhibited lower values until a criticaltemperature reached and exhibited higher values at higher temperatures. The criticaltemperatures at which J-R curve attained a minimum, were shifted to higher temperature withfaster loading rate and were observed at the temperature ranges 2500C and 2960C for load-line

displacement rates of 0.4 and 4.0 mm/min, respectively. The J-R curves at load-line
displacement rate of 0.4 mm/min were higher than those at 4.0 mm/min between RT and
2500C, whereas the J-R curves above 2960C sho\ved negative loading rate dependence.
The variation of Ji With temperature for each load-line displacement rate is given in Fig.

5. The cl~*k miuation toughness, J,, was defmed as value of J-integral at the crack initiation

point obtained from DCPD method. The trend of variation in Ji with temperature and loading
rate was similar to that of J-R curves. The minimum points of Ji appeared at temperatures of200-2500C and 250-2960C for load-line displacement rates of 0.4 and 4.0 mm/min,
respectively. At these temperatures the values of Ji were smaller, about 300/0, than those at RT
for each load-line displacement rate. Up to 2000C the Ji increased with increasing loading rate,
but the trend was inversed, that is negative strain rate sensitivity, at the temperature of 2960C.

It was observed that reduction of fracture toughness and negative loading rate dependence

occur at the temperature region of DSA. Also, the lower fracture resistance region withioading rate is a important characteristic of DSA in fracture behavior.

3. Effect of DSA on LBB Analysis

3-43.1 LBB Analysis
The evaluation of leakage-size-crack and flaw stability was performed for simple
piping system using material data discussed in the previous section. The sets of material
properties to identify the characteristics of DSA were employed in the analysis as input data.

The J-R curves were extrapolated using a following power-law relationship [ I I I in order to
get the large amounts of crack growth needed for the analysis.

J = J, + CAam(2)

where C and m are fitting parameters. The fitting parameters obtained from the regression ofdata points between 0.5 and 2.5 mm in crack extension, and listed in Table I . It was assumed

that pipe has a circumferencial through wall crack and a dimension of 670 mm outer diameterand 32 mm thickness. The piping system was operated under 7.3 MPa and 289'C steam, andwas subjected to remote axial tension and bending moment as shown in Fig. 6.

The calculation of crack length for given leak rate was performed using PICEP code[12]. All calculations were based on the assumptions of a flaw suface roughness of 0.05 mm

and an elliptical circumferential crack which plastic-zone was corrected. For a given load

level, the crack length that produces the desired 37.85 l/min (lOgpm) Ieak rate wasdetermined.

J/T diagram method was employed for evaluation of flaw stability [l]. The elastic-

plastic J-integral estimation formula from EPRI NP-6301-D [13], known as the EPRJ/GE
estimation method, was used in the J and T calculations. This method superposes solutions
corresponding to elastic and fully plastic conditions to obtain the elastic-plastic results forthrough-wall crack in a pipe. For the case of a through-wall crack in a pipe under remote axial

tension and bending moment loading, the following J-integral estimation equation is used forthe elastic-plastic solution.

J = Je + Jp[ ) [ n+it' )-ff '4Rt2E+fb' R3t2E+aobcoR(lT-e) - ' po
jT

3-5

(3)

The individual terms are as described in detail in Reference [ 1 3]. The instability point in the'ppli'd and Jm't.,ial Versus Tmat*,i*1 on a single

J/T method is found by plotting Jappli'd against T

figure as illustrated by a plot showing J versus T in Fig.7. The intersection of the two curves is

the instability point and the corresponding J value is Jj~st, from which the instability load canbe determined.

3.2 Effect of DSA on the Leakage-Size-Crack
Fig. 8 shows the dependence of leakage-size-crack length on temperature and strain
rate of material properties. The crack length at RT was the largest. The smallest crack lengthfor same load level was observed at 296"C rather than at 4000C. The variation in crack length

was not linear with temperature, particularly it was clear at I .39xIOA/s. The crack length
increased with increasing cRy, and n, and with decreasing cc. The role of c~y* and n is dominant.

Accordingly, the largest value of ay* and n at RT is a cause of maximum crack length.
Although the value of c~y* at 296'C is larger than that at 4000C, the crack length at 296~C issmaller. This is owing to small n at 296"C associated with DSA in tensile properties.

In comparing the leakage-size-crack lengths at each temperature, the crack length for6.95x I 0~2/s at 296'C was larger than that for I .39x I 04/s, whereas crack length at I .39x I 04ls

was larger than that at 6 95xlO /s for other temperatures The vanation m cly, and ot withstrain rate at same temperature is small compared with that in parameter n except for data at

RT. Therefore, the inverse trend of crack length with strain rate at 296"C relates to thechanging of dependence of n on strain rate in the DSA region. Consequently, the strain
hardening characteristic of DSA results in a decreased leakage-size-crack length at 296"C, andthe negative strain rate dependence in DSA region is a cause of increasing leakage-size-crack
length with increasing strain rate.

3.3 Effect of DSA on the Flaw Stability
The effect of loading rate in material properties on the instability load for a given crack

length, 2ec/C=0.1, was represented in Fig. 9 as a function of temperature. It showed that thevariation of instability load with temperature was similar to that of J-R curve for each loadingrate. The minimum load appeared at a certain range of temperature which depends on loading
rate of J-R data. It shows that a decrease in J-R curve caused by DSA alters instability load

3-6

significantly, although the crack driving force of J-applied decreases due to the enhanced
tensile properties and balances out the decrease in J-material [ 1 4].

Also, the effect of strain rate of tensile data on the instability load was observed. In thetemperature below 200*C, the instability load at I .39x 104/s was smaller than that at 6.95x I O~

2/s. Between 200 and 296"C, however, the trend of variation with strain rate was reversed.
According to the results of parametric study, the instability load for a given crack length
increased with increasing ay., and with decreasing cc and n. In the lower temperature region,
therefore, an increase in instability load with strain rate is due to high value of (~y.. However,the variation of (~y, with strain rate is negligible between 200 and 296"C, while the value of nincreases with increasing strain rate. Accordingly, a decrease in instability load with strain rateis caused by high value of n at high strain rate in this temperature region. The disappearanceof strain rate sensitivity of ay, and positive strain rate sensitivity in n are characteristics of

DSA in the tensile properties.

Because of the characteristics of DSA in the material properties, the minimum

instability load occurs near the nuclear power plant operating temperature, and thetemperature corresponding to minimum load depends on loading rate of J-R data. In addition,

the DSA decreases instability load with increasing strain rate of tensile data in thistemperature region.

3.4 LBB Allowable Load Window
Figs. I O and 1 1 represent the effects of DSA on the LBB analysis as a form of "LBB

allowable load window" [ 1 5]. The band between the minimum moment to produce desired
leakage rate and the maximum-allowable-moment for a given crack length is LBB acceptableregion. In these figures, the safety factor for leakage-size-crack length and applied load was

not applied. Fig. I O shows the LBB allowable region as a function of temperature of material

properties at quasi-static loading condition. The LBB allowable region at 250 and 2960C weredecreased by about 300/0 compared with that at RT or 350"C. It reflects that DSA reduces LBB

allowable region significantly. Fig. I I exhibits the LBB allowable region with various
combinations of tensile data with J-R data at 2960C, near the plant operating temperature. The

variation of LBB allowable region with loading rate of J-R data was considerable, whereas theinfluence of strain rate in tensile data was negligible. LBB allowable region at loading rate of

3-7

4.0mm/min in J-R data was smaller, about 1 50/0, than that of 0.4mm/min for same strain rateof tensile data. As shown in Fig. I I , LBB allowable region depends on the loading rate ofmaterial data at plant operating temperature. However, it is very difficult to know loading rate

for practical piping system with certainty. In order to obtain the conservative results in the

LBB assessment, the effects of loading rate on the material behavior have to be investigated

and then the loading condition corresponding to lower bound material properties should be
used for material test.

4. Conclusions

Using material properties in SAI06 Gr.C piping steel obtained from various testing
conditions, the evaluations of leakage-size-crack and flaw stability for simple piping system
was conducted to estimate the effect of dynamic strain aging (DSA) in material on the resultsof leak-before-break (LBB) analysis. The main conclusions are as follows:

1 . The leakage-size-crack length in the dynamic strain aging (DSA) region was smaller thanthat at other temperatures. Also, DSA increas~d the crack length with increasing strain rate.
These are due to the strain hardening characteristic and negative strain rate sensitivity of DSA
in tensile properties.

2. The instability load reduced in the DSA region, and the temperature corresponded to aminimum load depended on loading rate of J-R data. In this region, the instability loaddecreased with increasing strain rate of tensile data. This is owing to a decrease in fracture

toughness, a disappearance of strain rate sensitivity of ay*, and an enhancement of strain

hardening caused by DSA.

3. LBB allowable region in the range of DSA was decreased by about 300/0 compared withthat in other region of temperatures. Also, it varied with loading rate of material data at plant

operating temperature.

3-8References[ I] NUREG-l061 , Vol.3, October 1 984.
[2] B. Mukherjee, Nucl. Eng. & Des. I I l, 1989, p63.

[3] C. W. Marschall. M. P. Landow and G. M. Wilkowski, NUREG/CR-6098, October 1993.
[4] C. W. Marschall, R. Mohan, P. Krishnaswamy and G. M. Wilkowski, NUREG/CR-6226,
October 1 994.
[5] B. F. Beaudoin, D. F. Quinones and T. C. Hardin, J. Pres. Ves. & Piping 43, 1990, p67.

[6] S. Lou and D. O. Northwood, J. Mater. Eng. & Perf. 3, 1994, p344.
[7] S. S. Kang and I. S. Kim, Nucl. Tech. 97, 1992, p336.
[8] C. C. Li and W. C. Leslie, Metall. Trans. 9A, 1978, pl765.
[9] M. L. Vanderglas, Int. J. Pres. Ves. & Piping 43, 1990, p241 .

[lO] NRCPIPES-Computer Code and User 's Guide, Mar. 1995.
[1 I] G.M. Wilkowski, G. Karmer, P.Vieth, R. Francini, and P. Scott, PVP-Vol.280. ASME,
l 994, p22 1 .

[12] D.M. Norris and V. Chexal, PICEP, EPRINP-3596-SR, Dec. 1987.
[13] A. Zahoor, EPRI NP-6301-D, Vol. I , June, 1989.
[14] O. E. Lepik, B. Mukhajee, Int. J. Pres. Ves. & Piping 43, 1990, p285.

[1 5] A.D. Nana and K.K.Yoon. PVP-Vol.287/MD-Vol. 47, ASME, 1994, p35.

Table I Parameters of J-R curves obtained from power-law fitting(Aa = 0.5-2.5 mm).

3-9

5.0

380

340
320

&
:E
ol

-1- "I '39X1 Od"'1-•- i:'1.39XIO~ s~1-•-- ill.3gl~IOJ •'t-,~- i•1.39XIO~ •'1-v- i•e.9GD(1O~ ,r's60-c- ill '39XIO:8 l't-~- i'l'39xIOle t"

4.5-1P- '=6 99(1O" l~\~~\\ ',

3co*~~\'^\~:

200~\\a~~(~~~\~~\~\ab~\\:~~~:~\\L~~ Iv

,

c 4.0

200

=:==

3.5

240

." ~"

220

O~ hafdeoiog

lea8ted f•gbo~~'~ l'-\.leoh•rderdrl,egial

regba
3.0

200

o co loo leo 200 2co 300 3so 40o 450

o co loo leo 200 2eo 300 350 400 4eo
TempentJre Co

Temperature Cc)

Pig. I Dependence ofyield stress on temperature

Fig. 2 Dependence of n parameter in Ramberg-

and strain rate.

Osgood relation on temperature and strain rate.

250

8-1- i•1.39XIOJ •'1-•- i=1.39XIO'8 s~1

7

4.a mmltak,

200-,h e=1.39XIO~ ,it6t's-v- ~=6.95X1 O~ t~1' \:'\~~X
C:=] '\L-~~

(~

~ Ieo

~
~:

~

4

~:

*-t~l 'ellb~

3

au' ha,dentr,gv

reglolll

2

GL4 maVmu,-c*-'h-•ernted fogtoa

lOO'o ha,~,edng.

regky,

eoo so loo leo 2co 2co 300 350 4co 4co

o 50 1 oo I co 20o 2eo 30o 350

Fig. 5 Dependence ofcrack initiation fracture
toughness, Jj, on temperature and loading rate.

Fig. 3 Dependence of n parameter in RambergOsgood relation on temperature and strain rate.

2ooo

2000
1 800

a RT

1 6co

A 2000cOo 1 50eC

L=0.4mmlmin

150ec

o 296'C* 350qC ll'

1 200~~ 1000

/

leoo

JL.

1400

V

1200

,
X

J~A

c~

lvvvv7vV'vvlPvvvvvvvvvvvvv v

2OOec/
25Oec eee'e e AAAJ~LLJLLLL
296'C eee AAA .~L
AA

~ 80o
1

6cc

ooo

400

400

2 oo

200
2

6

4

8

10

Crack extension, Aa (mm)

o

o

x

IF1'1'(F" 1)1'F c 'vvvvv:•o""cl'F"'
vv

!

800

ll

eee JbJb
3eo'C eeeeexx"x'cc xx'~dj~)'xx
xr,y7v I v 'rvvvvvvvv v~ 100o!_o

RT L=4.0mm/mul .ll~'d"'d

1800

o

v ' 2SOQC

1400

o

400

TemperatJ,e Co

Temperature Co~"'c'ocl'c'
.,p

2

4

e

8

crack extension, Aa (mm)

(b)

(a)

Fig. 4 Dependence of J-R curve on temperature and ioading rate of(a) 0.4mmlmin and (b) 4.0mmlmin.
3 - lO

10Mb (kN-m),\y

600 800

400
7000

1 ooo

o (Jrr)

t

e

6000

(F- (J/T)

'o

i

sooo

o

i
/

2000
,

~) ~ ,;,~'u Cn /

o

.

.

e

1 ooo

P. QJ2. t

/

o

~iS,

~

l'/

~1.

l,l

.

l-~1 -~1•1'

o

.

/

,pp

-'- PIJ.

/~~ 3000m"eo!

c~ 4 oooE'o1400 1600 1800 2000

1 200

/

/

..

o

coo 800

200

1000

Tearin9 Modulus

Fig. 6 Circumferential throughwall-cracked in a

Fig. 7 Calculation ofinstability load by using
typical Jfr method.

pipe and applied loads.

2.0

4.5

-1- RT J_._ 29~ t'l jgXla""leak rate=37.85 UminJ- -a- - a=6.95XIO~/• L=0.4mrrurrdll

J f r4netr,od

4.0-~tF 40OeC

1.5-1- '*1 .39XI0~h

20 Jcaj. I o-1- a*1.39Xlarl/$- -(h - N.95XIOs/$ L~LOm,rvm~lJ-- --ap-a- - RT
400GC l- -o- - 296~ i-6.g5xla""~ 3.5e;, 10z:Ez:~ '0::•--~~ 3.0

~
i

._aL

~F
0.5
2.5

0.10 0.15 0.25

0.0

2.0

0.20

0.05

Fig. 9 Effect of temperature and loading rate ofmaterial data on the instability load for 2eJC=0. I .

5.0

5.0

~11'39XIO""' L=0'4 ruTumh

4.S

4.5

4.0

-~- 25ctC

3.5
3.0

A -1~ 296~,;_::~\_:\:. :A . t_3eoec

3.0

z2~ 2.5

~~
1.S

0.00

iF

o a LBB alleWable__ reslon

J L=0'4fTvlvmtr,

J L;4,0r~'~Q--' o"' Ill8.95XIO't,s

A 8""'\~~~\ \'\ L.jB aliOWable
re9'on

0.5

Jo 'a

o.05

2.0
1 .S

1 .O

JI'o

o. 5

0.0

~

Oo

1 .o

-c'-P I 1 1 '39Xlar$/s

3.5

2. 5

2.0

T1296ec -1'-' il I _ 39x I Gr4/$-- c'-' il6.95XIO"I'-df- RT

4.0

i

300

Temperature CO

Fig.8 Minimum bending moment versus leakagesize-crack length calculated by PICEP code.

~

200

lOO

O

20 JC

0.0

0.10 0.15 0.20 o.2s 0.30 0.35

0.00 o.05 0.10 o. 15 o.:o o 25 o 30 o 35

20 JC

20 JC

Fig. lO Variation in LBB allowable region withtemperature of tensile and J-R data at quasi -static

Fig. I I Variation in LBB allowable region withloading rate of tensile and J-R data at 296'C.

loading rate.

3-11

400

4

Probabilistic Fracture Mechanics Analyses of

NuclearPressure Vessels Under PTS Events
G. Yagawa* (Univ. of Tokyo), S. Yoshimura(CRIEPI),

M. Hirano (JAERI)PROBABILISTIC FRACTURE MECHANICS ANALYSES
OF NUCLEAR PRESSURE VESSELS UNDER PTS EVENTSGenki YAGAWA and Shinobu YOSHIMU~
Depar~lent of Quantum Engineering and Systems Science,
The University of Tokyo,
7-3-1 Hongo, Bunkyo, Tokyo I 13, Japan
Tel. : +8 1-3-3812-21 1 1 Ext. 6993

Fax : +81-3-56843265
e-mail : yagawa@gen.u-tokyo.acjp

Naoki SONEDA
Central Research Institute of Electric Power Industry,

2-1 1-1 Iwado-kita, Komae, Tokyo 201, Japan

Masashi HI~NO
Japan Atornic Energy Research Institute,
2-4 Shirokata-shirane, Tokai-mur~ Naka-gun, lbaraki 3 1 9- I I , Japan

Abstract

This paper describes some recent research activities on probabilistic fracture mechanics (PFM)
for nuclear reactor pressure vessels (RPVs) performed by the RCI 1 1 research cornmittee of theJapan Society of Mechanical Engineers (JSME) under a subcontract of the Japan Atomic Energy
Research Institute (JAERl). To establish standard procedures for evaluating failure probabilities ofnuclear RPVs, we have performed PFM analyses for aged RPV under pressurized therrnal shock(PTS) events. The basic problems are chosen from some of US bcnchmark problems such as
EPRI (Electric Power Research Institute) and US NRC (Nuclear Regulatory Conunission) joint

PTS benchmark problems and H. B. Robinson problems. Various sensitivity analyses are
performed to quantify effects of uncertainty of data on failure probabilities. Employed in this study

are four PFM computer codes developed in Japan and in USA.

l . INTRODUCTION
Studies on efficient utilization and life extension of operating nuclear power plants have bccome
increasingly important since ages of the first-generation nuclear power piants are approaching theirdesign lives tl]. It is easy to imagine that a practical life ofplant might be usually longer than its

design life by considering conservatism embcdded in design practices. In order to predict a
remaining life of plant, it is necessary to select those critical components that strongly influence theplant life, and to evaluate their remaining lives by considering aging effects of materials and otherfactors. However, when cvaluating reliability of nuclear structural components, some problems arequite forrnidable beeause of lack of information regarding a past operating history, material
property change and unccrtainty in daJrnage models. Accordingly, if structural integrity and safety
are evaluated by deteministic fracture mechanies approaches, it is expected that the results obtained
are too conservative to perform a rational evaluation of plant life and to make judgment of life
extension because of accumulation of conservatism of au related factors.
In this regard, the Probabilistic Fracture Mechanics (PFM) has become an Important tool [2-4] .
The PFM approaches are regarded as appropriate methods to rationally evaluate plant life since it
can consider various unccrtainties such as sizes and distributions of cracks, degradation of material

strength due to aging effects, accuracy and frequency of pre- a!id in-se!vice inspections. For
example, they can be used to improve maintenance or inspection schedule of structwal: componentsof nuclear power plants [5-7]. They are also expected as tools to derivc input data for the

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probabilistic safety assessment (PSA) or the probabilistic risk assessment (PRA).

A numbcr of PFM computer programs have bcen developed and applied in praaical situations in
the last two decades [8-22]. Some of the present authors have developed the three-dimensional(3D) J-integral database based on fully plastic solutions [23, 24] and a fully automated finite
element system for calculating 3D stress intensity factors [25], and perfol~led PFM analyses ofnuclear piping using the database [26, 27]. They have also developed an efficient PFM computer
program using a parallel Monte Carlo method [28, 29].
On the other hand, the PFM approaches have some weak points such that PFM analysis results
are hardly verified through experiments. Thus it is key issues to verify the consistency and thevaiidity of PFM computer programs through the comparison of calculation results of some welldefined benchmark problems [30, 3 I].
In Japan, one research activity on PFM approaches to the integrity studies of nuclear pressure
vessels and piping (PV&P) was initiated in 1987 by the LE-PFM subcomrnittee organized in theJapan Welding Engineering Society (JWES) under a subcontract of the Japan Atomic Energy
Research Institute (JAERI), and had continued for three years [31, 32] . The activity was followedby the RCI 1 1 research comrnittee organized in the Japan Society of Mechanical Engineers (JSME)in
1991, and fmished in 1995 [33]• Many researchers from 4 univers
utilities, 5 venders and 2 software companies participated the activity. The purpose of the activity
was to establish standard procedures for evaluating failure probabilities of Japanese nuclear PV&P,

combining the state-of-the art knowledge on structural integrity of nuclear PV&P and modem
computer technology such as parallel processing. Within the activity, we have set up the following

three kinds of PFM round-robin problems on (a) primary piping under normal operating
conditions, (b) aged RPV under norrnal and upset operating conditions, and (c) aged RPV underpressurized therrnal shock (PTS) conditions. The basic part of the last PTS problems is takenfrom some of US benchmark problems [30, 34]. For these rOund-robin problems, various
sensitivity analyses were performed to quantify effects of uncertainty of data on failure
probabilities. Some of the analysis results can bc found in elsewhere [28, 31, 35-37].

This paper sumrnarizes some sensitivity studies on aged RPVs under PTS events. Here two
original Japanese PFM computer programs [17, 20] are utilized, together with two US ones suchas OCA-P [12] and VISA-II [1 5].

2. PFM ANALYSIS PROCEDURE OF RPV FAILURE UNDER PTS EVENTS
A nuclear RPV is typically modeled as a thick cylinder, which is subjected to operating intemalpressure and temperature. There are several scenarios in PTS events. For example, when a
Ioss-of-coolant accident (LOCA) is occurred, the inner surface of the vessel is suddenly cooled
down by lower temperature water provided from an emergency core cooling system (ECCS). If a

LOCA is less severe, i.e. a small-break LOCA (SBLOCA), both thermal tensile stress andmechanical tensile stress due to the intemal pressure are applied to the vessel simultaneously, andan postulated inner surface crack might start to grow under a certain condition [8, 34, 38, 39]. In
the present PFM analyses, a conditional probability of vessel failure, i.e. crack initiation and vessel

break, under one PTS event is caiculated using the Monte Carlo method. The PFM analysis
consists of the deterministic analysis part and the probabilistic analysis one. In the deterministic
analysis part, we first calculate a transient temperature distribution, a transient stress distribution,and stress concentration, considering a specific transient. In a real: situation, the beltiine portion of

the vessel is first cooled by the ECCS, and thus the vessel temperature distributes all in the
thickness, axial and circunferential directions. In the typical PFM analyses of PTS events, only atemperature disoibution in the thickness direction is taken into account for the purpose of
simplicity. The thermal stress caused due to such a temperature distribution and the mechanicaltensile stress caused due to the intemal pressure are mainly considered. Tue peak stress caused due
to mismatch between clad and base materials in thermal expansion cocfficient is taken into accountas well. If reliable data are available, residual stress could be inciuded in the analysis.

The linear elastic fracture mechanics (LEFM) is basically employed in the present PFM analyses,
so that the stress intensity factor K is calculated. Fracture toughnesses, i.e. crack initiation

4-2

toughness Ktc and crack arrest toughness Kla, are regarded as functions of neutron fluence andtemperature of material. Thelmal aging effects are not explicitly considered, but are alreadyincluded in bacture toughness values together with imdiation embrittlement effects.
Random variables to be considered aJre initial crack depth and aspect ratio, fracture toughness,
neutron fluence, RThDr, impu:ity content of vessel material, and so on. The norrnal distributions
a!1e assumed for all the random va:iables except the initial crack depth. If any leliable distributions

are available, they should be employed in the analyses. Otherwise, well known open data found inliterature are used. The numbcr of cracks existing within the vessel could also be a randomvariable. To consider such a situation, the Poisson distribution is assumed.
In the present analyses, the probability of crack initiation exceeding a Ktc value and that of vessel

break, i.e. the K value of a growing crack exceeding a Kh value are calculated. As optional
calculations, hacture at an upper shelf toughness are~ plastic collapse of uncracked ligament andwarm prestress effects are considered.

3. PFM CODES AND ANALYSIS PROBILMS
Four.PFM computer programs are employed in this study. The two of them, i.e. PROFMAC-II[17] and MHIPFM [20] aJre Japanese codes, and the others aJre US codes, i.e. OCA-P [12] and
VISA-II [15]. PROFMAC-II is developed by Central Research Institute of Elecuic Power Industry(CRIEPI), while MHIPFM is developed by Mitsubishi Heavy Industries Ltd. (MHI). All thecodes adopt the Monte Carlo method.
Two sets of round-robin problems are analyzed in the present study. The one is taken from the

PTS bcnchmark study organized by both EPRI and US NRC in 1992 [30]. The other is takenfrom US H. B. Robinson problems [34], which were solved by Oak Ridge National Laboratory(ORNL) in 1985 using OCA-P. In the latter analyses, vessel failUre probabilities were calculated
for 28 postulated transients of the reactor.

The EPRVNRC PTS benchrnark problems originally include thirteen cases and some additional
special cases, which were set up to clarify the difference of the existing US PFM codes. Mainfeatures of the basic cases named Al and B I can bc summarized as follows :(a) No cladding is assumed.
(b) Exponential decay of coolant temporature with time is assumed.
(c) Coolant-wall heat transfer coefficient and intemal pressure are assumed to bc
time-independent.
(d) Axial (case Al) and circumferential (case B l) infinitely long cracks are assumed.
(e) Original OCrAVIA crack depth distribution [8] is assumed.
(D RTNDT shift and through-wall fiuence artenuation in US Regulatory Guide I .99 Rev.2 [40]are assumed.
(g) Ktc and Kh criteria aJre employed as crack initiation and crack arrest criteria, respectively.

(h) Upper-shelf toughness of 220 MPaV~ is employed.
(i) No warm prestress (WPS) effect is considered.
Main features of the H.B. Robinson problem are as follows :(a) Cladding is considered.
(b) Temperature, coolant-wall heat transfer cocfflcient and intemal pressure aJretime-dependent, and specific to each of postulated transients.
(c) Axial inflnitely long crack is assumed.
(d) Marshall distribution of crack depth and Marshall crack detection probability [41] areempioyed.
(e) RTNDT shift of Guthrie equation [38] and through-wall fluence attenuation in US

Regulatory Guide I .99 Rev.2 [40] are assumed. . ,(O Ktc and Kh criteria aJre employed as crack initiation and crack arrest criteria, respectivcly.

(g) No WPS effect is considered.

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One of the 28 postulated transients, called case "9.22B" assuming a main stearnline break event,
was adopted in the present study.
To evaluate K values for a crack subjected to arbi~rily distributed therrnal loading, so called
influence function methods are uscful. Ref. [42] gives the influence functions for infinitely longaxial cracks and those for fully circumferential cracks. Ref. t43] presents the influence functions
for a semi~Wptical suface crack in the axial direction in cylinder. Ref. [ 1 5] prcsents the influencefunctions sirnilar to those of Ref. [42], but taking into account cladding. In the present study,
those functions are adopted.

4. COMPARISON STUDY OF FOUR PFM CODESThe EPRIINRC PTS benchmarking problems were first analyzed using the following three PFM
codes, i.e. PROFMAC-n [17], OCA-P tl2] and VISA-II [15]. The analysis results of case B1,
i.e. conditional break probability vs. neutron fluence, are shown in Figure I . The figure show
fairly good agreement.

Next the transient "9.22B" of H.B. Robinson PTS problem was calculated using OCA-P and
MHIPFM [20]. The conditional break probability of an original ORNL analysis [34] was 3.61 x
l04 , while those of the present analyses are 4.38 x 104 (OCA-P) and 4.87 x 10~ (MHIPFM).
Although the present analysis results are slightly higher than that of the ORNL analysis, they agree

well among others in an engineering sense. This slight difference might bc caused due to someerror of input data. Coolant temperature, heat transfer coefficient and pressure are not given in any
numeral forrns in the original literature [34], and then we took these data from the frgures.
These results clearly demonstrate that the Japanese and US PFM codes are reliable to perform thepresent study.

5 . SENsrTIVITY ANALYSES
Using the four PFlvi codes, various sensitivity analyses are perfomaed on the two sets of PTS
problems in order to quantitatively evaluate effects of input data on failure probabilities of RPV
under PTS events. The input data studied are initial crack shape, the probabilistic distribution of
initial crack depth, cladding, RTNDrr shift, impurity content, the through-wall distributions ofmaterial properties, preservice inspection (PSI) and wann prestressing.
5.1 Effects of Initial Crack Shape
Using VISA-II code, we exarnined effects of initial crack shape on failure probabilities, solving

the EPRVNRC PTS benchrnarking problems. Figure 2 shows calculated break probabilities foraxial crack cases. Here an infmitely long axial crack and semi-euiptical surface cracks with four
different aspect ratios, i.e. 2, 6, 10 and 100, are assumed. The figure shows that as increasing thecrack aspect ratio, break probability increases, and that assurning an infinitely long crack gives thelargest break probability. When assuming an aspect ratio of 6, the results are very close to those of
the inflnitely long crack cases. We perfo!med similar sensitivity analyses for circuuferential crack
cases, and found the same tendency. There are very few data on probabthstic distributions of crack
aspect ratio in literature. Only a norrnal distribution and a log-no!mal one proposed by LLNL tl3]
are often utilized. These distributions suggest that the averaged aspect ratio ranges 2 to 3.

5.2 Effects oflnitial Crack Depth
Several probabilistic distributions of initial crack depth have been proposed and used in PFM
analyses. Figure 3 shows six kinds oftypical distributions, i.e. (a) & (b) the upper and the lowcr

limits of the Lorence Live!~lore National Laboratory (LLNL) distribution [9], (c) Marshall
distribution [41], (d), (e) Weibull and log-normal distributions of Brueckncr-Foit model [lO] and(O OCTAVIA distribution r8]. Figure 4 shows calculated break probabilities for axial crack cases,

4-4adopting the six distributions shown in Figure 4. Analyzed here are again the EPRLrNRC PTS
bcnchmarking problems. VISA-II code is used. Figure 4 shows that the discrepancy among the six
results ranges over two orders of magnitude. The upper limit of the LLNL distribution, theMarshall distribution and the lower limit of the LLNL distribution give almost the similar andhighest results among the six distributions. Although there are some research activities to obtainplant-specific distributions of crack depth using pre-service / inservice inspection (PSI/ISI) data[44], it is still popular to use open data in literature. The Marshall distribution may be the mostpopular one. Frgure 4 suggests that using the Marshall distribution gives us reasonably
conservative results for PFM problems ofPTS events. Further detailed discussions on effects ofcrack depth distributions on failure probabilities can bc found elsewhere [35, 36] .

5.3 Cladding Effects
Using the influence functions for K values in refs. tl5, 42], we examined cladding effects on
failure probabilities. PROFMAC-II is used here. Figure 5 shows ealculated break probabilities.
Here axial cracks are again assumed. Case Al (without cladding) and case A I I (with cladding) of

the EPRVNRC benchrnarking problems are solved. Figure 5 clearly shows that considering
cladding raises break probabilities by about one order of magnitude. 'Since any influence functions
for a circunferential crack with cladding have not been developed, cladding effects are still notclealty understood in such cases. As conclusions, cladding effects have to bc taken into account.

5.4 Effects ofRTNrTr Shift
VISA-II code employs the three kinds of RTNUr shift equations, i.e. (a) the modified Gathrieequation [42], (b) Randall equation [12] and (c) the equation in US'Regulatory Guide I .99 Rev. 2[40]. (d) One Japanese equation is also proposed in JEAC4206 [45]. In the original EPRLfNRC

bcnchmarking problems, the equation of (c) was employed. In the present study, case Al(infinitely long axial crack) is solved assuming the above four RTNDT shift equations. VISA-II
code is utilized. The analysis results show that equations of (b) and (c) give almost the same
results, while (a) and (d) give almost the same results. The four results range within a few factors
difference. In other words, any of the above four equations for RTNDT shift can be used.

5.5 Effects of Impurity Content
Impurities such as Cu and Ni are known to have large influence to the estimation of neutron
irradiation embrittlement of vessel materials. According to the RTNDT shift equations mentioned
previously, Cu content affects RTNDT Shift value more directly than Ni content. Then influences of

Cu content were examined in detail, soiving the EPRIlNRC bcnchmarking problems with OCA-P
code.
Figure 6 shows calculated break probabilities plorted against wt. % of Cu content. The variation
range of Cu content in the figure corresponds to the actual data for the base metals of US RPV
materials. The figure clearly illustrates that the variation of Cu content significantly influences
break probabilities, and that the break probability for Cu content of 2.2 wt.% is more than twoorders greater than that of I .2 wt.%.

5.6 Through-wall Distribution of Material Properties
As pointed out in the report of the EPRLfNRC PTS benchrnarking study [30], one of the largest

differences among various PFM codes for PTS problems was caused due to the difference inmodeling the through-wall distribution of material properties such as Ktc, Kta and RTNDT shift.

The question arisen was whether deviations of these properties from their mean values areresimulated at every steps of crack propagation or not. When this problem was solved at the firsttime, VISA-II resimulated au of the three material properties at every steps of crack propagation(Procedure l), OCA-P resimulated only Ktc and Kla values (Procedure 2), and WPFM [ll]4-5resimulated none of them (Procedure 3), i.e. fixed these material properties during emck
propagation, Since WPFM estimated the highest failure probability, Procedure 3 was finauy
recomrnended as the standard procedure.
In the prcsent study, thc effects of thc modeling of the through-wall distribution of the tineematerial properties are again exaJrnined using OCA-P and VISA-II codes. Figure 7 shows
calculated break probabilities for the threc different procedures. The difference between the resultsof Procedures I and 3 is of more than one order of magnitude, and Procedure 3 gives the highest
break probability. Conservatism embedded in Procedure 3 is a strong motivation to adopt thisprocedure as standard. However, further investigation is still required since there is not any
physical basis for this procedure.

5.7 Effects ofPSI
Recently, a new data on probability of non-detection of crack, i.e. Arakawa equation, which isfor UT-based PSI, was proposed in Japan [46], based on a number of expedmental data. The
Marshan equation [4l] is also popularly used. These data are shown in Figure 8. Here Arakawa(inner) means the data for embcdded cracks, while Arakawa (surface) means the data for suface
crack. The EPRVNRC benchmarking problems of axial crack cases a:e solved, assuming the threedifferent PSls of Figure 8. PROFMAC-n is used here. Figure 9 shows the analysis results. For
the purpose of comparison, 'the figure also shows the results without any PSI. Figure 9 showsthat PSI with the accuracy of Arakawa equations reduces break probabilities significantly, i.e. twoto three orders of magnitude. In Japan, UT-based PSI is applied to a whole volume of vessel, and
X-ray based PSI is further applied to welded portions. Considering the fact that actual cracks are

generally embcdded, the Arakawa equation for embedded cracks seems a reasonable choice as
standard in Japan.

5.8 WPS Effects
The wann prestressing (WPS) effect has been well verified through various experiments.
Nevertheless, most of integrity assessments for PTS events are performed without considering thiseffect. Neither PFM problems employed in the present study originally take the WPS effect into
account. The effects are here evaluated, assuming that a crack does not begin to propagate when
the K value is in a decreasing process. Figure 10 shows break probability for the axial crack caseofthe EPRIlNRC benchmarking problem. Figure 1 1 shows that ofthe H. B. Robinson problem.
In both problems, failure probability reduces by considering WPS effects. In the H. B. Robinson
problem, such a reduction effect is not so significant, while failure probability reduces by nearly

two orders of magnitude in a lower neutron fluence region in the EPRVNRC bcnchmarking
problem.

6. CONCLUSIONS
Benchmarking study using the four different PFM codes, i.e. two Japanese and two US's, were
perforrned, and the consistency of the codes were verified. Using these codes, some sensitivity
analyses were conducted to quantitatively evaluate the influences of the input data, i.e. (a) initial
crack shape, (b) the probabilistic distribution of initial crack depth, (c) cladding, (d) RTNDT shift,(e) impurity content, (O the through-wall distributions of material properties, (g) pre-service
inspection (PSI) and (h) wa~n prestressing. It is clearly shown that in most cases, these data affectfailure probabilities significantly. Therefore, we should use in the PFM analyses as reliable input

data as possible. However, if any reliable data are not available, the data resulting in mostconservative results could be chosen, referring the analysis results presented in this paper. In orderto establish standard procedures to evaluate best estimates of failure probabilities of nuclear PV&P

components, we will continuously improve the round-robin problems and accumulatc sensitivity
analysis results.

4-6

ACKNOWLEDG~NTS
The authors wish to thank all the membcrs of the RClll cornmittee in the Japan Society of
Mechanical Engincers (JSME) for their efforts to porfonn this work, and also thank the Japan
Atomic Energy Research Institute (JAERI) for financial support.

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Fig.4 Break probabilities of axial crack cases
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4 - lO4

5

Im proved Toughness of the SA 508 Class 3 Steel for
Pressure Vessel Through the Steel-Making and Heat

J.T. Kim *, H.K. Kwon,

H.S. Chang (HANJUNG)

NuclearTreatmentImproved Toughness of the SA 508 Cl. 3 Steel for Nuclear
Pressure Vessel Through the Steel-Making and Heat Treatment
J. T. Kim, H. K_ Kwon, and H. S. Chang

R & D Center, HANJUNG, Changwon, Korea
Ab stract
The compartive investigations of forgings by three steel-making procedures of SA

508 class 3 have proven that vacuum carbon deoxidation(VCD) offered low toghness

properties, even though the toughness met the required values. The toghness
properties of the steel by VCD were closely related with the cooling rate from thesolution treating temperature. To obtain the secure toughness for the VCD steel, therecommendable minimurn cooling rate from the austenitizing temperature is 151C/min.

The modified VCD steel by adding aluminium and silicon-killed steel were thoughobtained the secure toughness properties. The fracture toughness(Klc) was
significantly improved by the silicon-killing and the modified VCD. These wcre
resultcd from the fine austenitic grain size. It is observed that the grain size wasbclow 20~m when using the modified VCD and silicon-killing, and that was 50~m

whcn using VCD.

Introduction
The Reactor Pressure Vessel(RPV) of the pressurized water reactor nuclear power

plants is manufactured using forgings and plates in manganese-nickel-molybdenum
low alloy steel designated according to the ASME/ASTM standards SA 508 class 3
and SA 533 Type B class 3. This steel is manufactured by silicon-killing or VCD.

However, since the influence of the steel-making practices on the toughness
properties of the RPV steel has not been systematically investigated[1, 2], we havestudied to improve the toughness properties of the forged steel for RPV through thesteel-making procedures. It has been investigated the effects of cooling rate at
austenitizing temperature on the Charpy irnpact resistance of the 2 l/4Cr-1Mo steel[3,41, however, few report for the SA 508 class 3 steel has been found[5].

Thcrefore, the effects of the steel-making procedures on the toughness properticsof the SA 508 class 3 steel were investigated by using several shcll forgings, whichwere manufactured by three steel-making methods. The rrrst is the VCD, the sccondis the modified VCD by adding aluminium, and the third is the silicon-killing mcthod.
Thc effccts of cooling rate from the solution treating temperaturc on the toughncss

5-1

properties were also studied because the secure toughness properties in the VCD
steel had not been obtained.

Experimental Procedures
Several shell forgings for RPV of ICOO MWe nuclear power plant were
manufactured by using conventional hot pierced ingots. The applied steel-making
procedures are VCD, modified VCD by adding aluminium, and silicon-killing. The
chemical compositions of the exarnined steel are indicated in the Table l. The forgedshells were norrnalized at 900 ~C , quenched in water at 880 -9001C and then tempered
at 650~. The simuiated post weld heat treatrnent for the test blocks was executedfor 32h at 620-6301C. The cooling rates were measured at various locations of theshell with attached thermocouples during quenching. In order to exarnine the cffectsof cooling rate on the Charpy V-notch impact properties for the VCD steel, thecooling rates of the specimens were varied from I ~~/min to 1501C/min. The heat
treating conditions of the specimen manufactured by VCD method are described inTable 2.

The influence of steel-making procedures on the Charphy V-notch impact
properties of the SA 508 class 3 steel has been exarnined at through-thickness of theshell. The nil-ductile transition temperature(TNDT) by the drop weight test andfracture toughness were examined at If4-thickness of the shell thickness. The

fracture toughness(Klc) values were converted from the Jlc tested at room

tempern*ure, that is, Klc = [ E • Jlc/(1 - v2) Jl/2, where E is Y
is Poisson's ratio.

Results and Discussion
The tensile properties of the SA 508 class 3 steel manufactured by threesteel-making methods were measured at l/4-thickness of the shell. These results aredescribed in Fig. l. From the these results it seems that the tensile properties are
not related with the steel-making procedures for this steel. However, the elongation
was slightly improved by the silicon-killing, while the tensile and yield strength werealrnost same regardless of the steel-making methods.

The Charpy V-notch impact transition curves of the SA 508 class 3 stcel withsteel-making methods in the tangential and axial direction are shown in Fig. 2 andFlg. 3, respectively. These are the typical results obtained at l/4-thickncss of theshell for RPV. From these results the 68Joule(50ft-lb) energy transition
temperature(vTr68) and 50"/o shear fracture appearance transition temperature(FATTso)

are summarized in Table 3. The vTr68 and the FATT50 Of the VCD steel werehighcr than those of the modified VCD and silicon-killed steels. Furthermore, thedcviations of vTr68 and FATT50 Of the VCD steel were larger than those of thc
modiflcd VCD and silicon-killed steels.

5-2

The life-time of the RPV is limited because of the core area embrittlement, so theresistance against fast fracture of the SA 508 ciass 3 steel shell is indexed by themeans of the reference nil-ductile transition temperature(RTNDT). Fig. 4 shows thevariations of various transition temperatures of the SA 508 class 3 steel for RPV
with steel-maldng processes. From this result silicon-killed steel has the best values.

The fracture toughness(Klc) of the SA 508 class 3 steel was significantly
improved by the silicon-killing and modified VCD. The Klc values of forgins by the

VCD, modified VCD and silicon-killing at room temperature are 426, 574 and665MPa~ m, respectively.

It is suggested that these were resulted from the reduciug of sulphur content andthe fining of the austenitic grain size. It was analyzed that the sulphur contents of

the VCD, modifiod VCD and silicon-killed steel were 0.004, 0.0035 and 0.002.0/*
respectively. Therefore, the contents of deliterious sulphidic non-metallic inclusion for

the fracture toughness and Charpy impact resistance might be reduced by decreasing

the contents of sulphur. It was observed that the grajn size was 20~m when using
the modified VCD and silicon-killing, and that was 50~m when using VCD.
Many investigators[1, 2] have been studied the effects of the steel-making
methods on the mechanical and metailurgical properties of a steel. P. Bemabei et al[1]are recomrnended VCD process for the SA 508 class 3 steel, regardless of thedeoxidation practice, for a given grain size, Charpy V-notch impact energy at 4.4~

aftcr welding is always higher than in quenched and tempered condition. Comparing

the steel manufactured by the three steel-making methods, they observed that the
VCD heat showed the best FATTSO and absorbed energy values. However, when theheavy wall thick steel is manufactured by VCD, the grain size is coarsened and the
secure toughness properties are not obtainable. Therefore, many a forgemasters havebeen preferred to add aluminium or deoxidize by silicon to manufacture the better
forging shell for RPV.
Fig. 5 shows the variations of vTr68 and FATT5O with cooling rate of the SA 508

class 3 steel by VCD. The cooling rates were measured between 800~C and 500t,
which is the range of bainite trasformation temperature in this steel. From theresults the transition temperatures of the vTr~ and FATT50 were lowered as thecooling rate increased. The Charpy V-notch impact energy at 4.4 ~C (E4.4) were also
improved as the cooling rate increased. The variations of E44 with the cooling rates

are shown in Fig. 6. When cooling rate at 880t was I ~/min, the ferrite content of
about 20% was observed. However, when the cooling rate is above the l0.11C/min, it

was not observed any ferrite phase as shown in Fig. 7. In case of l0.l~:/min incooling rate, the Charpy impact resistance was though satisfied the requiredconditions, but its values were low margin states. From these results, in order toobtain the secure Charpy V-notch impact properties in the SA 508 ciass 3 steel for

RPV, it is suggested that the recornmendable minimurn cooling rate fromaustenitizing temperature to bainite finish temperature is lorlC/min.

Moreover, the influences of the double heat treatment on the Charpy V-notch

5-3

impact properties of this steel were also exarnined, and these effects were shown in

Fig. 5 by open indicators(O, A ). The double heat treatment means the repeatedprocess of quenched and tempered treatment When double heat treatrnent for thissteel was performed, the transition temperatures and impact energy were improved.
Fig. 8 shows the grain sizes of the single heat treated and double heat treated stcel.

The grain size of the double heat treated steel is fmer than that of the single heat

treated one. It may be resulted from the new grains formed at prior austenitic andbainitic lath colony boundaries via dissolution of the two phase(a + Y) duplex
stnlcture which had formed on heating.

Concluding Remarks
The tensile properties were almost same regardless of the steel-making for the

SA 508 class 3 steel, though the elongation was slightly improvod by thesilicon-killing. Comparing the 68 Joule energy, 50~6 shear fracture appearance, andreference nil-ductile transition temperature(vTr68, FATTSO and RTNDT) of the steels

manufactured by the VCD, modified VCD and silicon-killing methods, the sihcon-kmedsteel has the best properties. The fracture toughness(Klc) was significantiy improvedby the modified VCD and silicon-killing. The are resulted from the fining of austenitic

grain size. The toughness properties of the steel by VCD
the cooling rate from the solution treating temperature. To obtain the secure
toughness for the VCD steel, the reconunendable rrtinimum cooling rate from theaustenitizing temperature is 151C/min.

Ref erencesl. P. Bemabei, L. Callegari, M. Scepi, and T. Salinetti, Steel Forgings, ASTM STP

903, E. G. Nisbett and A. S. Melilli, Eds., Arnerican Society for Testing andMaterials, PA, 1986, p. 275 ' '2. M. Erve, F. Papouschek, K. Fischer, and Ch. Maidom, Nuclear Engineering and
Design, vol. 108, (1988) p. 487

3. W. E. Erwin and J. G. Kerr. WRC 275, Feb., (1982)

4. Y. Murakami, T. Nomura, and J. Watanabe, Application of 2 l/4Cr-lMo Steel for

Thick-Wall Pressure Vessels, ASTM STP 755, G. S. Sangdahl and M.
Semchyshen, Eds., Arnerican Society for Testing and Materials, PA, 1982, p. 383
5. T. Kodaira, Tetsu to Hagane, vol. 73, No. 14, ( 1987) p. 1656

5-4

CJ'
I

C!1

used in study.

Table I .

Chemical compositions of the SA 508 class 3 steel

Table 2.

Heat treatment of the SA 508 class 3 steel used in study_temperature(vTr68) withTable
the steel-making practices of the SA 508 class 3 steel.

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Steel-making Mehhod

Steel-making Method

Fig. 1 . Comparison of tensile properties of the forged SA 508 class 3 steel for the reactor

pressure vessl with steel-making method.

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5-6

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3. Charpy V-notch impact tranSition curVeS Of forged SA 508 claSS 3 Steel for reactOr preSSurevessel with steel-making method. The specimenS Were taken in the axial direction.

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Comparison of varlous transition temperatures of the forged SA 508 class 3
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5-7

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5-8

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Fig. 8. Grain sizes of the SA 508 c~ass 3 steel

Fig. 7. Microstructures of the SA 508 c!ass 3 steel

with cooling rate at 880C. a~ IC/min,with the coolino rate at 880T~_ (1) I r]/min,b) 10.It/min, c} 10.1 + 1 1.41C/min(doubleb) 10.1r)/min, c) IO.{ ~ I i.4r]/mln(a40uble

quenched and tempered treatment)

Quenched and tempered trea[me)nt)

5~9

SESSION B

6 Crack Shape Evolution of Surface Flaws under Fatigue Loadingof Austenitic

Pi pesI.S. Hwang*,

J.H. Kim (Seoul National

Univ. )

Crack Shape Evolution of Surface Flaws
under Fatigue Loading of Austenitic Pipes

I.S.

Hwang and J.H. Kim

Department ofNuclear Engineering, Seoul National University
Shinlim-dong, Gwanak-ku, Seoul 1 5 1 -742, Korea

From the results of fatigue crack growth and fracture tests of a Ni-Fe base superalloy (Incoloy908), an improved fatigue life analysis model has been derived from the framework of
Newrnan and Raju. For a plate geometry with an initial semi-elliptical surface crack in its
thickness direction, the new model can predict the evolution of crack aspect ratio for a wide

range of initial crack geometry for Incoloy 908 that has been considered as a primary

candidate for the ITER central solenoid conduit. The improved model is applied to the
ITER central solenoid magnet life prediction. The model predicted the conduit fatigue lives
that are about two times the values obtained by assuming the constant aspect ratio during thecrack growth, for both free-standing and bucked against central solenoid designs. Therefore
the realistic modeling is recommended for the future magnet conduit designs.

l. Introduction

The fatigue crack growih behavior and fracture mechanics properties of a nickei-iron basesuperalloy (Incoloy 908) have been presented for a temperature range from 298 K to 4 K [1 J.

6- l

The material used for property measurement had a processing history that simulates the
anticipated manufacturing schedule of a central solenoid magnet conduit for the International

Thermonuclear Experimental Reactor (ITER) [2]. The Conceptual Design Activity(CDA) of the ITER has shown that the plasma pulse-cycle life was limited due to fatigue

crack growth phenomenon in the superconductor conduit when a free-standing coil design

was used for the central soienoid [3]• During the Engineeri
BACS (Bucked Against Central Solenoid) design has been adopted in order to alleviatetensile load and hence the fatigue impact on the conduit. At the bottom and top ends of thecentral solenoid, however, the bucking is not accomplished due to the configuration of thetoroidal field coils. Hence, the EDA design analysis must ensure that the fatigue crack
growth in the conduit will not limit the ITER system life.

Fatigue analysis during CDA was based on the assumption that a semi-elliptical surface crack
will grow in the conduit thickness direction with a constant aspect ratio (crack depth/crack

half length) until the rmal fracture. The final fracture was predicted based on the linear

elastic fracture mechnics (LEFM). The fracture mechanics measurements for Incoloy 908

surface crack changes systematically with crack growih and 2) the plasticity involved in thefmal fracture is too extensive to apply the LEFM approach [1 J.

Since the accuracy In fatigue and fracture analysis of superconducting magnet structure
impact directly on both reliability and economy of the future nuclear fusion system, thedevelopment of refined design models are desirable. In this paper, we present an improved
design mode[ for fatigue analysis of Incoloy 908 conduit based on the property measurementand data analysis. The model is then applied to the current central solenoid design and theimplication of the improved prediction results on the ITER magnet R&D has been discussed.

6-2

2. Fatrgue Crack Growth Model Development

The fastest fatigue crack growih is expected to take place in the conduit thickness direction

from the initial suface flaw with a semi-elliptical geometry, as shown in Figure I .
Although the initial flaw is likely to be associated with welds, the model development isbased on the currently available data of Incoloy 908 base metal. Newman and Raju showedthat surface crack growth can be predicted by applying Paris Law to each of the transverse
and thickness direction, as follows[4];dcdawhere dc/dN and da/dN are the crack growth per stress cycle expressed in mm/cycle, andconstant CA, CB and m are the material constants, and AK is the amplitude of stress intensity

factor during a fatigue cycle. They also developed the stress intensity factor for the semielliptical surface crack in Ref. [4]-In the surface cracked tension test of Incoloy 908, crack length at surface was measured using

an optical microscope. Table 2 shows the test data for two tension fatigue tests (#203 and#204) started from a semi-circular EDM notch which have more measurement data points
than other tests. Using these data, dc/dN is computed after Eq. (1) where the exponent m is
set to 3.0 for the temperature range from 298 K to 4 K, based on our earlier fatigue test resultswith compact tension specimens [1]. The calculation of stress intensity factor, K, after

Newman and Raju equation [4] at c = 90' (A) and = O'( (B) requires the crack aspect ratio
value. Since these tests showed fmal aspect ratio values different from the initial of I .O, theestimation of CB is possible by assuming two diffrent values, i.e., the initial and the fmalvalues. Two different CB-values are then geometrically averaged to determine the constant.
as fol]ows;

6-3

CB ~d•~~ = 4 . 06 x 10 9 ,( mm / Cycle
Thus we obtained the Paris Law model for a surface crack growing in the transverse directionin Incoloy 908 plate, as follows;dc,(mm )3
I Cycle)dN = 4 .06 x 10 9 x ( AKo (3)Figure I compares the suface crack data with Eq. (4) as well as Paris Law correlation ofcompact tension specimens obtained for Incoloy 908 at room temperature. The correlations
for two specimen designs agree to each other within a factor of two.

For the development of da/dN. Newman-Raju suggested to apply a correction factor of 0.9 on
the stress intensity factor calculated for the point B of Fig. I after Ref. [4], in order to better

fit to experimental data such that the coefficient CA can be determined from CB by;

CA = I CB
O. 9~Table I summarizes the surafce cracked tension test results for a wide range of aspect ratio.

Low aspect ratio data were obtained by bending fatigue loading[l]. To derive theappropriate correction factor for Incoloy 908 from experimental data, the coefficient CA isassumed to be;

CA = I CB
oc~

By the least square frtting ofthe data set, cc is deterrnined to be 0.904. This value is almost

the same as 0.9 used by Newman and Raju[4]• Finally Eqda

5 5 x 10~9 x ( AK )3.0 ,( mm / Cycle)

Using the final fatigue crack growih analysis model given by Eqs.(3) and (4), the entire
surface cracked tension data set of Incoloy 908 that inciuded both testion and bending tests
are predicted reasonably well, as shown in Fig 3.

Using the model for Incoloy 908, the fatigue life of test specimens in Table 2 was calculated.

6-4

The measured and calculated fatigue life to reach the final measured crack depth is shown toagree to each other with the maximum error of about 500/0, as shown in Figure 4. The factthat Incoloy 908 behavior agrees well with the model in Ref.[4] suggests that the base metal
has fairly uniform mechanical properties in both a- and c- directions.

3. Application to Fusion Magnet Design Analysis

ITE~_CpA_C~Ptral__S_9,e_noi_d_pe_s~ig ~In the ITER CDA, the fatigue life in the central solenoid conductor conduit was calculated
by assuming an initial surface flaw with an aspect ratio (a/c) of 0.2. The design requires a

total of 40,000 plasma pulse cycles. During each plasma pulse cycle, two tensile loading

cycles are involved ; one associated with the pre-bias and the other with the end ofbum(EOB). The stress during each loading cycle ranges between a peak value and aminimum of zero. The geometric and material properties of central solenoid conductor
conduit used for the CDA analysis are listed in Table 3.

In order to examine the effect of evolving aspect ratio during crack growih on fatigue life,

the calculation is made by applying the new correlations of Eqs.(3) and (4) while assuming

Incoloy 908 as the conduit material. Although Eqs. (3) and (4) were obtained from thefatigue test at room temperature, these are applied to the magnet operation temperature of 4
K. As the fatigue crack growih rate of Incoloy 908 is iowered by a factor of two at 4 K, theuse of room temperature model would result in pessimistic predictionstl]. The EOB stressis assumed to be 0.8 times the pre-bias stress. The results of calculation are compared forthe case of a fixed and changing aspect ratio in Table 5. From these results, it is shown thatthe new model allowing the evolution of aspect ratio with crack growth predicted about two

times the cycle life of the case with the constant aspect ratio. Therefore the realistic
trcatment of surface crack growth ieads to a significant gain in the design life compared withthe constant aspect ratio model.

6-5

ITER EDA Central_ Soi~_npi~ pes_i~!P_In the ITER EDA(Engineering Design Activity), Incoloy 908 is assumed as the central
soienoid conduit material, and other design parameters are listed in Table 4. It is evident thatthe BACS design significantly reduces stress and fatigue burdens. The design requirement ofoperational cycle is a I O0,000 stress cycles that corresponds to a 50,000 plasma pulse cycles
with a safety factor of 2. This value of cycle correspond to a 200,000 stress cycles when pre-bias stress and EOB stress are considered. In this case, EOB stress is also assumed to be 0.8

times the pre-bias stress as in the previous CDA case. The results of predicted conduit
lifetime is shown in Table 5. When the aspect ratio is assumed to remain constant during thecrack growih, the predicted fatigue life is about 1 80,000 cycles. When the aspect ratio
evolution is taken into account using Eqs. (3) and (4), the life is predicted to be about 395,000

cycles. The new prediction, thus, shows that the design requirement of ITER EDA issatisfied with a signifrcant margin, with the BACS design.

5. Conclusion

Fatigue behavior of surface cracks in a superconductor conduit material, a Ni-Fe superalloy,

is predicted based on an empirica] model that takes account of the evolution of crack shapeduring the crack growth. The model predicts that the aspect ratio (a/c) of surface flaws
asymptotes to I .O, i,e., to a semi-circular geometry under the cyclic tension loading during

plasma pulse cycling offusion magnets. The measured fatigue lives of Incoloy 908 surface
cracked tension specimen agree well with the model prediction within a factor of two. The
results implicate that fatigue life can be predicted at the greater accuracy by taking into
account of evolving aspect ratio with fatigue crack growth. The consideration of the aspectratio evolution of surface crack during fatigue cycle results in an increase in the calculated

cycle life of the conduit material by more than a factor of two, when applied to the central

solenoid magnet design fbr the Intemational Therrnonuclear Experimental Reactor (ITER).

Nevertheless the Bucked Against Central Solenoid(BACS) design is required to meet the

6-6

ITER design requirement on fatigue life.

Acknowledgement
Authors would like to express their deep gratitude to the Korea Science and Engineering
Foundation for the financial support.

References[l] R.L. Tobler, I.S. Hwang and M.M. Steeves, Fracture Mechanics Behavior of a Ni-Fe
Superalloy Sheath for Superconducting Fusion Magnet, Part I : property measurement,
submitted to Fusion Engineering and Desigr~ 1995

[2] L.S. Toma, I.S. Hwang, M.M. Steeves, and R.N. Randall, Thermomechanical Process
Effects on Hardness and Grain Size in Incoloy Ailoy 908, Adv. Cryo. Eng. Maters., Vol.
40(1 994) 1307-13 14.

[3] IAEA. ITER Magnet System, ITER Document Series N0.26 , 1990

[4] J.C. Newrnan. Jr. and I.S. Raju, An Empirical Stress-Intensity Factor Equation for theSurface Crack. Eng. Fract. Mechs., Vol. 1 5, 185-192, 1981 .

[5] K. Nishiok~~ K. Hirakawa and I. Kitaur~~ Fatigue Crack Propagation Behaviors ofVarious

Steels, Sumimoto Search No. 17, 39-55, 1977.

[6] D.L. Corn, A Study of Cracking Techniques for Obtaining Partial Thickuess Cracks ofPre-selected Depths and Shapes, Eng. Fract. Mechs., Vol. 3, 45-52, 1971 .

[7] J.G. Merkle, A Review of Some of the Existing Stress-Intensity Factor Solution for Part-through Surface Cracks, U.S.Atomic Energy Commision ORNL-TM-3983, 1973.

6-7

[8] A. Nyilas, Fatigue Charaterization of Jacket Materials at 4 K. Kfl( Report on NET Study
Contract 89-201, 1 990.

[9] T.W. Orange, Fractire Testing with Surface Crack Specimens Joumal of Testmg and
Evaluation. Vol 3, No. 5, 335-342, Sep. 1975

[ I O] N. Mitchell, Fatigue Assessment of the Structural Steel in Superconducting Coils for

Tokamaks, Fus. Eng. Des., Vol. 19(1992) 225-233.

[1 1 J Z. Mei, C.R. Krenn, and J.W. Morris, Jr, Initiation and Growth of Small Fatigue Cracks

in a Ni-Base Superalloy, Metall. Trans. A. vol 26A(1995) 2063-2073 .

6-8

Table l . Surface cracked tension test results for a wide range of aspect ratio.

Table 2. Test data for two tension fatigue test(specimen #203 and #204)
conducted at I O Hz and R=0. I at room temperature.

6-9

Table 3. Geometric and material properties of the Central Solenoid conduit usedin the fatigue calculation for ITER CDA design.

Table 4. Geometric and material properties of the Central Solenoid conduit usedin fatigue calculation for the ITER EDA design.

Tabie 5. Results of calculated fatigue life of the Central Solenoid conduit for

ITER CDA and EDA design.

6- lO

I

.

1'F

2h

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lp

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lt

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B
2c

.

b

t

Fig. I Surface flaw with a semi-elliptical geometry in finite plane.

6- Il

1 . OE-3

c+

Based on a/c = (a/c)jnitial
Basecl on a/c = (alc)final

Ccrnpact TesiOO Test (Ref.1)/ SCTEq.(3)1 .OE4c~i . -~~

c+. ' ' . '

~5

z
~
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(O)~F,F./

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(e)+

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34

5

A K avg [MPa ml/2]
Fig. 2 Crack growih rate in transverse direction vs. Stress intensity factor andpredicted power-law curve for SCT specimen #203 and #204 and Paris law correlationof compact specimens obtained for Incoloy 908 at room temperature.

6+12

1 .OOlr-- -~~-i~- - ~~ ,~~- ;F~::~~~~-- '~~_

c ~:~;~ ~\/ '/ *~ _

0.80

,

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0.60

+
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,

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,

0.40

#201 predicted

/
,
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#201 measured#202 predicted#202 measured#203 predicted#203 measured#204 predicted#204 measured#208 predicted,#208 measured

,

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0.20

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,

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o .oO

0.20

0.40

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0.60

0.80

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(a)
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#206 predicted
[

08

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t206 measured
t207 predicted

t207 measured

!

'l~
/'/1P

0.6

.l
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0.4

l

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02

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o o

0.2

0.4

alB

0.6

0.8

(b)
Fig. 3 Changes ofaspect ratios with crack growth as predicted by the optimized cc for

specimens under (a) pure tension load and (b) pure bending load.

6- 13

10

2. O

~
~ 1.8
~

=0) I .6:~~+cU

~L I .4~ I .2+~~~ll I O

+

'\

++
+

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+

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Fig. 40.2

0.6ratio (a/c~
Measured 0.4
final aspect

The ratio of a calculated fatigue life to a measured fatigue life
to the final crack depth, af for all specimens ofTable I .

6 - 14

0.8

1.0

7

Study on the

Effect of the Crack Length on

M. Kikuchi*(Science Univ. of Tokyo)

the JIC ValueSTUDY ON THE EFFECT OF THE CRACK LENGTON THE JIC VALUE
Masanori Kikuchi
Science University of Tokyo, Noda, Chiba, 278, Japan

ABSTRACTThe ductile fracture tests are carried out using CT specimen, three point

bend specimen and CCrspecimens made of A533B steel and Aluminum alloy
with different crack lengths. It is shown that the apparent Jrc value increases

due to the decrease of the crack tip constraint. It is pointcd out that theincrease of the apparent Jlc value is partly due to the error of theconventional equation to estimate the J value. Bascd on the FEM analyses,
these apparent Jlc values are corrected and are compared with the valid Jlc
values. The good co-relation betwcen apparent Jlc value and thc Q factor,
proposed by O'Dowd and Shih, is shown for every specimens.

1. Introduction
The J integral concept is one of the most important parameters in theelastic-plastic fracture mechanics. The HRR solution[1,2] gives the theoreticalbasis for the J integral fracture criterion. It means that when the crack tip

stress fields agree with the HRR solution, the J integral is the dominantparameter which describes the strcss, strain and displaccment fields at thecrack tip uniquely.
By the authors previous papers[3,4], it has becn shown that the crack tip

stress fields don't agree with HRRsolutions well in many cracked plates. In

such cases, the J integral is not a unique dominant parameter and thefracturc criterion may bccomc differcnt. It is called the constraint cffect.

O'Dowd and Shih [5,6] proposed Q factor as the second term, whichcorresponds to thc stress triaxiality at thc crack tip. Anderson [7], Jun [8],and Nikishikov [9] also proposed other paramcters which improves theestimation of the crack tip fields in the clastic-plastic fields, respectively.

In this papcr, the fracturc toughncss tests of A533B stecl and Aluminumalloy specimens are conducted using scveral kinds of spccimen
configurations and crack lcngths. Thc cffect of thc initial crack lcngth on theapparcnt fracture toughncss is studied. Then the finite clcment analyscs are

conductcd and thc crack strcss ficlds arc comparcd with I~R solutions. Thc
apparcnt fracturc toughncss valucs arc corrected bascd on the numcrical

7-1

results, and the rclation between Q factor and thc corrected apparent
fracture toughncss is obtaincd.

2. Fracture toughness tests.
Figure I shows the shapes and sizes of the test specimens made of A533B
steel. The three point bend specimen is made in three cases with diffcrent
initial crack lengths. They are changed in a/W=0.1, 0.3 and 0.55. In the CCr
specimen, the crack length is changes in two cases, as 2a/W=0.4 and 0.6. The

same three point bcnd specimens are made by Aluminum alloy 2025-T6.
The crack length is also changed in three cases as same as Fig.1. The CT
specimens with a/W=0.3 and 0.6 are also made by the same aluminum alloy.
The mechanical properties of these two materials are shown in Table l. Bothmaterials show ductile behaviors at fracture.

The fracture toughness tcsts are conducted based on the JSME testing
standards[10]. As there is no testing standard for CCTspecimen, the similar
method as CT and three point bend specimens are applied to this specimen.
The JIC value is defined as the J value at the initiation of the stable crack

growth. The results of every specimens for both materials are shown in
Table 2. As they don't show the constant value for each material, they arecalled the apparent Jlc value, Jlcap. The JSME standard states that the valid
Jlc value should satisfy the following two conditions.

b 2: 25
JlacP (1 )afslap

B :~afs25 Jlc (Z)Where b is the ligament length and afs is the flo,w stress. In the A533B
specimens, the three point bend specimen with a/W=0.55 satisfies these
conditions, and in aluminum alloy specimens, three point bend specimen

with a/W=0.55 and CT specimen with a/W=0.6 satisfy these conditions.
Then the valid Jrc values are determined as 135kN/m for A533B steel, and6.5kN/m for aluminum alloy, respectively. It is noticcd that other JICap
values are much larger than the valid values for both materials. As theinitial crack length becomes small, the apparent Jlcvalue becomes larger. In
both CCr specimens, the Jrcap values of two specimens are nearly equal tocach other, and both values are about 1.37 times larger that the valid one.
The effect of thc initial crack lcngth is not rccognized.

The increase of Jlcap valucs by the decre,ase of the initial crack length isimportant for the application of the fracture mechanics to the practicalproblcms. In the real structure, almost all of the rcal cracks are shallow

7-2

surface cracks. The crack length of thcm are comparatively smaller than thc
thickness of the plate. The constraint effect is inevitablc. Then the method
to estimate the apparent increase of the Jlcvalue is preferable.

3. Microscopic observation of dimple fracture of Aluminum alloy
specimen.
In the 2025-T6 aluminum alloy specimen, the ductile fracture occurs dueto the nucleation, growth and coalescence of voids. Clear dimple patterns areobserved on the fracture surface, as shown in Figure 2. This is the fracturesurface of CT specimen with a/W=0.1. The diameter of these dimples aremeasured. The results are shown in Table 3. In this table, number Nmeans
the number of dimples in 0.12mm2 area at the crack tip and at midpoint of
the specimen. In every spccimens, the number of dimples and average
dimple diameters are nearly same values.
Then the diameters of preceding voids are also measured. Preceding
voids are the void nucleated at the crack tip just before the dimple fracture,

as shown in Figure 3. Crack growth occurs by the 'coalescence of thesepreceding voids. The diameter of a preceding void indicates the criticalvalue of void coalescence, which is one of the important parameters for the
dimple fracture. They are nucleated in front of the crack tip during stablecrack growth. They are observed by introducing a fatigue cracking after
stable crack growth test. Figure 4 shows the photo of preceding voids. They
are observed independently on the fatigue crack surface. Their number anddiameters in 0.12mm2 area at the crack tip are also measured. The results
are shown in Table 4. Similar to Table 3, these values are nearly same forevery speclmens.
The results of Tables 3 and 4 show that these microscopic fracture
process as void nucleation, growth and coalescence, are not affected by thechange of the initial crack length. It means that the microscopic dimple
fracture mechanism docs not changc due to the change of the constraint at

the crack tip. The constraint effect appears only on the macroscopic
parameters as J integral.

4. Numerical analyscs by FEM.
The elastic-plastic FEM analyses are conductcd for thcse three types of
specimens. Figure 5 is an example of the numerical model of 3PB specimen.
By the symmctry of thc structure, a quarter part of the whole specimen isanalyzcd. As thc prc-crack is introduccd by thc fatigue loading, the crack

front configuration has some curvature. It's shapc ia measured in the
7-3experiment and is uscd in the FEM modelling.
The stress-strain relations of these two metals are approximated by thefollowing equation.

ae ~=_+(E )n_[~,)
for a~:aoa,Where ao is the yield stress, E' and n arc constants. They are determined asE'=400Mpa, n=7 for aluminum alloy, and 1055Mpa and 7 for A533B steel,rcspectively. These relations are also approximated by the Ramberg-Osgood[ na 4 type equations.
8 O +aaO
Where a and n are 1.48 and 8 for Aluminum alloy, and 4.1 and 8 for A533B
steel, respectively. They are used to calculate HRR fields at the crack tip.

4.1 Crack tip stress field.
Figure 6 shows the crack tip stress fields of 3PB and CT specimens madeof A533B steel. They are the results at the midplane of the specimen
thickness when the J integral at this location is nearly 135 kN/m, the valid
Jlc valuc. The solid line is the HRR solution. The strcss field of 3PB specimen
with a/W=0.55 agrees well with HRR solution. It is rcasonable because thisspecimen gives the valid JIC value. As the a/W value decreases, the crack
tip stress deviates from the HRR solution gradually. In the CCT specimens,
the crack tip stresses of both specimens with different 2a/W values agree

well, and both are largely different from HRR solution. Obviously, these
deviations of the stress fields from HRR solution have relations with theapparent increase of JIC values in these specimens.

OtDowd and Shih proposed Q factor as the second parameter for thecrack tip stress field, which is d~fined by the next equation.

at r=ZJ
ao' O=0 (5)Q=aee oO
Where (a OO)HRR is the a 66 by the HRR solution, and r is the distance fromthe crack tip. Qfactor is calculated and is shown in Figure 7. It shows thc Q
factor distribution along the crack front. In every specimens, the absolutevalue of Q is small in the center of the specimen and increases near the freesurface. The largest Qvalue is given by CCTspecimens.

4.2 Correction of Jrcap value.
Thc J value is cvaluated using the conventional cquation in the

7-4experiment. They are:for 3PB and CT specimens,J - (1 ~ v2)K2 + 11Apl (6)and for CCrspecimcn,

J-(1 ~ v2)K + I PdA PA (7)where the first term is the elastic part, and the second term is the plastic
part of the J integral, respectively. The first term is evaluated by the stressintensity factor by refcrring the Handbook[11], and the second term isobtained by the load-displacement record obtained experimentally. ~ inequation (6) is thc constant defined by the specimen configuration.
Originally, these equations are used for the deep cracked specimen. In thisstudy, short cracked specimens are used. Then the numerical error of the J

integral due to these c6nventional equations should be evaluated. The J
integral is evaluated using eqs.(6) and (7) based on the load-displacement
relation obtained by the numerical analyses, and they are compared withthose by the contour integration. The results of A533B steel specimens areshown in Figure 8. The ordinate is the J by the contour integration and theabscissa is the J by the conventional equation. Both J values of deep cracked
3PB specimens agree very well. But the rcsults of short cracked 3PB and CCr
specimens show that both values arc diffcrcnt from each other largely. It isconsidered that the contour integration gives the correct J value. It meansthat the J value is not correctly evaluatcd by the convehtional equation forthese specimens. Then the experimental Jlcap values are corrected using
these results. The Jlcap value of 3PB specimcn with a/W=0.1 is 339 kN/m,
which corresponds to 222 kN/m by the contour integration. By the similarway, every Jrcap values are corrected and thc results are shown in Table 5.
The Jlcap values of aluminum alloy specimens are also correctcd. As noticed
by this Table, the valid Jlc value changcs little by this correction, and othervalues change largely.
The relation between the corrected Jlcap value and the Qfactor is shownin Figure 9. The Jlcap value is normalized by the valid Jlc value of eachmaterial. It is noticed that the relations of two metals are very similar toThough somc amount of data scattcr cxist, they arecach other.
approximated by a single line. Ofcoursc, the number of data is not enough.
But this result suggests that by obtaining cnough data, we could estimate thc

Jlcap valuc from thc Q value. The problem is that wc should conduct FEM
7-5analyses for calculating Q factor. But recently, the computer capacity has
dcveloped rapidly, and it is not difficult now to conduct three dimcnsionalelastic-plastic analysis using adequately fine mesh pattern. For this goal, thedata of shallow surface crack problem are needed. This is our future target.

5. Concluding remarks.
The J dominant crack tip fields do not exists in many practical crack
problems, and the J integral is not the unique fracture parameter in these
problems. It seems the relation between Jlcap value and Q factor enables toestimate the Jlcap value for these cases. Other methods proposed by otherauthors[7-9] should also be studied from the practical viewpoint. So far, the

new and reliablc fracture criterion has not been proposed yet. For theestablishment of new fracture criterion, the mechanisms of the dimple
fracture should be studied more precisely.
Then many researchers'
corporation concerning experimental, theoretical and numerical studies may
be needed.

References[1] JW. Hutchinson, J. Mech. Phys. Solids, vol.16 (1968), pp.13-31[2] J.R. Rice and G.F. Rosengren, Ibid, vol.16 (1968), pp.1-13[3] M. Kikuchi, Int. J. Fracture, vol.58 (1992), pp.273-283[4] M. Kikuchi, Int. J. Pressure Vessel and Piping, 63 (1995), PP.315-326[5] N.P. O'Dowd and C.F. Shih, J. Mech. Phys. Solids, vol.39, N0.8 (1991),

pp.989-1015[6] N.P. O'Dowd and C.F. Shih, Ibid, vol.40, (1992), pp.939-963[7] T.L. Anderson, Int. J. Fracture, vol.41 (1989), pp.79-104[8] S. Jun, Engineering Fracture Mechanics, vol.44 (1993), pp.789-806[9] G.P. Nikishkov, A. Bruckner-Foit and D. Munz, Engineering Fracture

Mechanics, vol.52, N.4 (1995), pp.685-701[10] JSME S-OOl, Standard Method of Test for Elastic-Plastic Fracture
Toughness Jrc (Supplement, Ist Edition), (1992)[1 Il Y. Murakami et al., STRESS INTENSITY FACTORS HANDBOOI(. Pcrgamon
Press

7-6

2a

-'~ - -'-' 'W=ZO~'W=16

14

a =1.6, 4.8, 8.8B= B=4.5
2a=8, IZ
(afW=0.1, 0.3 ,0.55) (2afW=0.4, 0.6)

Three-point-bend (3PB) specimen

CCT specimen

Fig. I Shapes and sizes of the test specimens made of A533B s'teelTable 2

Apparent Fracture Toughness values

preceding void:d7\\fatigue crack.' ' /

crack

Fig.2

Dimple Fracture Surface (2025-T6)

7-7

OO
Fig. 3Preceding VoidsFig. 4 Photo of Preceding Voids (2025-T6)
Table 3

Nulnber and Diameter of Table 4 Number and Diameter of

tl

\\

~V

x

z

~~lg. 5

Example of the Numerical ~. Iodelof :3PB Specimen,' -:~

5e 3pB a/W=0.1
o 3pB aNV=0.3
o 3pB a/W=0.55

4o

A CCT 2a/W=0.4
A CCT 2a/W=0.6~~o~cb

3

o

b

\

qD
ID

b

oo
..o ooe ~ a.e

2

HRR

-1

ee o oo oe .e o 8
~~l~A

AA

a

A AA A aaA ee

e 3pB a/W=0.1
o 3pB aNV=0.3@ 3pB aNV=0.55

JL

1

.2

J ~ 1 35kN/m
2Y/B=0. 1o12

JL CCT a/W=0.4

J ~ 1 35kN/m

A CCT a/W=0.6

34567

o 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9

8

2Y18

rl( Jlgo)

Fig. 6 Crack Tip Stress Fields of 3PB and

Fig. 7 Distributions of Q Factor at the

CCT specimens made of A533B steel

Crack Tip (A533B steel)

400

f
300

~E
zJ'~ 200
~
1
_o

8 100

JA

Sample line

\

~

~

da
dA

A
A
dCL

e

,e
eee
e
e 3PB a/W'0.1
ee

o 3PB alW'0.3
e 3pB aMl=0.55
A CCT 2alW'0.4

A CCT 2aNV~.6
1 OO 200 300 400

Conventonal J (kNlm)

Fig. 8 Relation between J by the contour
mtegration and J by the conventional equation

7-9

1

4

o A533B steel(average Q for thickness)

e2025-T6 aluminum alloy(two dimension ,plane strain)o3

O~ CCT
0.4

i,_~

a
g

~2

3

o

3PB I O
PB

1

0.55

0.55 0.6o-1Q
Fig. 9 Relation between Corrected
Jlrc Value and Q Factor

7 - lO~•3CT

8Development of New Z-factors for the Evaluation of
Circumferencial Surface Crack in Nuclear Pipings

theY.H. Choi*, Y.K. Chung, Y.W. Park, J.B. Lee
G. Wilkowski (Battelle Memorial Institute)(KINS),

Development of New Z-Factors for the Evaluation of the
Circumferential Surface Crack in Nuclear Piping s

Y H Chor Y K Chung Y W Park J B Lee and G. Wilkowski"(*) Korea Institute of Nuclear Safety, Korea(*) Battelle Memorial Institute, USA

Abstract
The purpose of this study is to develop new Z-Factors to evaluate the behavior of acircumierential surface crack in nuclear pipings. Z-Factor is a load multiplier used in the

Z-Factor method, which is one of the ASME Code Sec. XJ's recomrnendations for theestimation of a surface crack in nuclear pipings. It has been reported that the load
carrying capacities predicted from the current ASME Code Z-Factors, are not well inagreement with the experimental results for nuclear pipings with a surface crack. In thisstudy, new Z-Factors for feiTitic base metal, ferritic Submerged Arc Welding(SAW) weld

metal, austenitic base metal, and austenitic SAW weld metal are obtained by use of
SC.TNP method based on GE/EPRI method. The desirability of both the SC.TNP method
and the new Z-Factors is exarnined using the results from 48 pipe fracture experiments
for nuclear pipings with a circumferential surface crack. The results show that theSC.TNP method is good for describing the circurnferential surface crack behavior and thenew Z-Factors are well in agreement with the measured Z-Factors for both feilritic andaustenitic pipings.

1 . Introduction

In the design stage of pipings in nuclear power plants, it is assurned that there is no
crack in the pipings.[1] There, however, are many micro-cracks in nuclear pipings becauseof material inhomogeneity and welding process problems. During plant operation, somemicro-cracks may grow into surface cracks and/or through-wall cracks which give
adverse effect on piping integrity.

ASME Boiler & Pressure Vessel Code Sec. XJ "Rules for Inservice Inspection of Nuclear

Power Plant Components" (hereaiter we denote this code as ASME Code ) requlres thesafety evaluation of cracks which are detected during inservice inspection.[2-4] The
Z-Factor method is one of the ASME Code's recornmendations to evaluate the behavior ofa cirrurrLferential surface crack in nuclear pipings. The Z-Factor is a load multiplier tocompensate piastic load with elasto-plastic load. It has bcen reported that the load
carrying capacities predicted from the ASME Code Z-Factor method are not well inagreement with the experirnental results for nuclear pipings with a suface crack.[5-7] The

reason for the disagreement is that the cunent ASME Code 'Z-Factors are not exact to8-l

predict the surface crack behavior in nuclear pipings.[7]The purpose of this study is to develop new Z-Factors to evaluate the behavior of acircurnferential surface crack in nuclear pipings. Many crack evaluation methods[5-17] suchas J integral method including GE/EPRI method[5-7, 11-17], R6 method[8], and DPFAD
method[9-lO] have been proposed to describe the suface crack bchavior in pipings. In thisstudy, the SC.TNP method[14-16], which is based on the GE/EPRI method, is used todevelop new Z-Factors for ferritic base metal, ferritic Submerged Arc Welding(SAW) weld
metal, austenitic base metal, austenitic Gas Tungsten Arc Welding(GTAW) weld metal,
austenitic Gas Metal Arc Welding(GMAW) weld metal, and austenitic SAW weld metal.
The desirability of both the SC.TNP method and the new Z-Factors is exanained using
the results from 48 pipe fracture experiments for nuclear pipings with a circurnferential
surface crack.[5-7]2. ASME Code Z-Factor Method
In the ASME Code, the following criterton rs speclfied for a crrcurnferenuai suface
crack in ferritic and austenitic pipings.

where Pb and S. are the applied stress and the allowable stress respectively. The
allowable stress, S., Is detennined by the ASME Code Z-Factor method as

J_ P p) p~(1 Z(SF) ' (2)S.= (SF)( Z ~ ' ~lwhere lyb, P., P*, and (SF) are the bending stress at incipient plastic collapse, thethermal stress (P* = O for austenitic pipings), the membrane stress, and the safety factor.
The above equation is basically similar to the equation of lirnit load method in the ASME
Code. The Z-Factors in Eq. (2) for the ferritic and austenitic pipings are given in the

ASME Code as follows : (Hereafter we denote Zasme for the Z-Factor given in the

ASME Code.)
F 'tic Base Metal

F 'i AWWI M
i' B M

Zasme-3 = 1.0

TAW M WWI M(5)

Austenitic SAW Weld M tal

Zasme-4=1.30 [1+ 0.010 • (OD-4)] ,

(6)

8-2

where OD is the pipe outside diameter(inches). The A in Eqs.' (3) and (4) is a function of
R~t which is the ratio of the pipe mean radius(R*) to the pipe wall thickness(t) , given as(we denote Aasme for the A given in the ASME Code)5~ (t)
R~ ~lO)
Aasme-[OR~- . ( t -1250.25]o.25lO~(t)
R* ~20=[04(R~\ 3.0]0.25' ~ t j~As seen in Eqs. (3)-(8), both Zasme-1 and Zasrne-2
diameter and R~t while Zasme-4 is a function of thevalue of Zasme-3 is I because fully plastic behavior is

(7)

(8)

are functions of the pipe outside
pipe outside diameter only. The
assurned for the austenitic basemetal and GTAW/GMAW weld metal.
3. Development of New Z-Factors
3.1 Definition

The Z - Factorof Z-Factor
as a load multiplier is defined as

ML

Z= M EP- max(9)where ML and MEP- are the lirnit load under fully plastic condition and the maximum
load under elasto-plastic condition respectively.

The limit load for a circumferential surface crack, ML, under fully plastic condition is
given as [2](D e+e~n
M =2 a/R~ t 2sinP ~( SinetP_
{ ( -~.
t )}
crf

(10)

(ll)

'r

(2) 6+p> nML=2cr/R~tfl2 ~ }a Sin e

T

l

2~T '
where e is

a

T

(12)

P

( 13)
,

the half crack length, e is the angle

from (-)y-axis to neutral plane,

8-3

af

IS

the

flow stress, and a is the crack depth. The cross section geometry of piping with acircumferential suface crack is shown in Fig. l. P* in Eqs. (ll) and (13) is the stt~ss
due to the pipe intemal pressure (p), given byR~

P~ = p (R -R•) '(14)

where Ri and I~ are the pipe inside radius and the pipe outside radius respectively.

The maximum load under elasto-plastic condition, MEP-~ax, is obtained from the SC.TNP
method as following section.

3 2 Determmation of MEP-max using SC.TNP Method
As mentioned above, many methods such as the J integral method including GE/EPRI
method[5-7, 11-17], R6 method[8], and DPFAD method[9-10] have been proposed todescribe the circumierential surface crack behavior in pipings under elasto-plastic

condition. Based on the GE/EPRI method, Ahmad et al.[15-16] proposed SC.TNP
method(~urface C;rack for lhi~: Eipe) to evaluate the circumferential surface crack
behavior.

In the SC.TNP method based on the GE/EPRI method, J is divided into elastic
component of J (J.) and plastic component of J (Jp). For a circurnferential surface crack in
pipings, the Jp is expressed as[13]a V~
T a'eo'ao'(1- )'a'hl'- J t 6 n+1(15)

where 80 and oo are the reference stinin and the reference stress respectively, hl is the
GE/EPRI coefficient given in the GE/EPRI Handbook[13], b is the uncracked ligaJrnent anda and n are the strain hardening coefficient and the strain hardening exponent respectively,
which are deterrnined from the following stress(a) vs. strain(8) relation.

:= a() .a(16)

Under the applied moment M, the stress a in Eq. (15) can be expressed as[15-16]

M sinl+ +cos
2t A 2 T) ''4 ' R2~. t' H

8-4

~L
( 17)

where p is the angle from x-axis to neutral plane(i.e. p=1c-e), Y is the angle from y-axis
as shown in the Fig. l, and H* and h3 are numerical coefficients given in the reference[15] and the GE/EPRI Handbook[13] respectively.
Using the Eqs. (15)-(17), and J-R curve of pipe material, the relation between momentand pipe rotation can be calculated by numeric iterations. The maximum load underelasto-plastic condition(MEP-) can be easily deterrnined form the relation.

3.3 Reference Material Properties used in the Z Factor Calculation
In order to calculate Z-Factor for nuclear pipings, the reference material properties of
the related nuclear pipings has to be given first.
Table I represents the reference tensile properties for ferritic and austenitic pipings usedin this study.[5-7, 17] These tensile properties were obtained from base metals of both thefeivitic pipings and the austenitic pipings. The same tensile properties obtained from thebase metals are also used for weld metals for the purpose of conservatism.
Fig. 2 shows the reference J-R curves for the ferritic base metal and the ferritic SAW
weld metal, while Fig. 3 shows those for the austenitic base metal including austenitic

GTAW/GMAW weld metals'and the austenitic SAW weld metal.[5-7, 17]

3.4 New Z-Factors for Ferritic Pipings (Znew-1 and Znew-2)
For a given material, Z-Factors are functions of fracture parameters such as pipe
outside diameter(OD), Rm/t, the crack depth, and crack length. In this study, Z-Factors forgiven materials are calculated for the above four fracture parameters. The pipe isunpressurized in all calculations.

The effects of pipe outside diameter and R~:t on the new Z-Factor for the fefritic basemetaJ (Znew-1) are investigated for 3 cases of pipe outside diameters (OD=4.5, 16, and 42

inches) with Rm/t=10. and 3 cases of R~/t (Rm/t=5, 10, and 20) with OD=42 inches
respectively.

For a given pipe outside diameter and R~:t, Z-Factors for both 4 cases of crack depth(a/t=0.1, 0.3, 0.5, and 0.75) and 4 cases of crack length (e/7t=0.1, 0.25, 0.5, and l) arecalculated.

For OD=16 inches with Rm/t=10, the limit moments, the maximurn moments obtained
from the SC.TNP method, and the Z-Factors are shown in Figs. 4, 5, and 6 respectively.
As shown in Fig. 6, the new Z-Factors are not sensitive to either the crack depth or thecrack length.

We define the maximum Z value among 16 cases in Flg. 6 as the new Z-Factor
corresponding to the given pipe outside diameter and Rm/t (i.e. OD=16 inches and Rm/t= lO)

for conservatisrn The ASME Code Z-Factor in the figure shows a constant value bccause
the ASME Code Z-Factor is not a function of crack depth or crack length.
The other Z-Factors corresponding to the other pipe outside diameters and/or the otherR. ~t values can be detemined by same method.
Fig. 7 shows the new, Z-Factors for the fdritic base metal as a function of the pipe

outside diameter with Rm/t=10. The new. Z-Factors are iower than the ASME Code,8 - 5

Z-Factor by about 40%. The figure also shows that the increasing rate of the new
Z-Factor decreases gradually as the pipe outside diameter increases, while the ASME
Code Z-Factor increases linearly as the pipe outside diarneter increases.

The new Z-Factor for ferritic base metal can be fitted up to OD=42 inches as (hereafter

we denote Znew for the new Z-Factor)' (OD-4) - 0.0002•A'-lwhere Anev-1 is a new A value for ferritic base metal, which can be obtained from therelation between Anew-1 and Rm/t as shown in Fig. 8. The Anew-1 value decreases as the
Rm/t value increases, while the Aasme values mcreases as the Rm/t value decreases Anew l
can be fitted as

5~ R_ ~20-= (t)

A,,ew I [ R,,$O ' 0125 + 0'875]~2.791(19)

New Z-Factors for ferritic SAW weld metal (Znew-2) can be obtained through same
method used in the determination of Znew-1. The effects of crack depth, crack length,
pipe outside diameter, and R~t on Znew-2 are also investigated.
Znew-2 as a function of the pipe outside diameter is shown in Fig. 9. The Znew-2
also show sirnilar 'tendency to Znew-1 The Z-Factor for feni:tic SAW weld metal can be
fitted up to OD=42 inches as_ ' (OD-4)-0.000176 • A_2 ' (OD

Zn~v-2=1.35[O 742+0 0134 A 2

where Ancw-2 is new A value for the ferritic SAW weld metal given as: 5~ R* ~20-= (t)

Ancu 2[ O R~
' 1161 + 0'16l]~0.383(21)

3.5 New Z-Factors for Austenitic Pipings(Znew-3 and Znew 4)

New Z-Factors for austenitic base metal and austenitic GTAW/GMAW weld
metal(Znew-3), and austenitic SAW weld metal(Znew-4) are also obtained through similar
methods used in the Z-Factor calculation of f~itic pipings.
The effects of crack depth(a/t=0.1, 0.3, 0.5, and 0.75), crack length (8/7c=0.1, 0.25, 0.5, andD, pipe outside diameter(OD=4.5, 16, and 42 inches with R~t=10), and R~t(R~t=5, lO,
and 20 with OD=42 inches) on Znew-3 and Znew-4 are also investigated.
Figs. 10 and ll show Znew-3 and Znew-4 as a function of the pipe outside diametar
with R~t=10, respectively. As shown in the figures, the values of Znew-3 are similar toZasme-3 although Znew-3 under OD=20 inches is slightly lower than l. The values of
Znew-4 are slightly higher than Zasme-4.
The difference between the new Z-Factors
and the ASME Code Z-Factors in austenitic pipings is smaller than those in felvitic
pipings because the austenitic pipings is more ductile than the ferritic pipings. The

8-6

Z-Factors can be fitted up to OD=42 inches as

Znew-3 = 1.00[0.729 +0.0217 • A~u'-3 ' (OD-4) -0.0002

Znew-4=1.30[1.048+0.0203 • Aneu'-4 ' (OD-4)-0.00
where A**-3 and A~e~-4 are new A values for austemtic base metal and austemuc
SAW weld metal respectively, given as_ = [(O .+06080.392]-0.512)-= ( )[ O . 0778

+ 0.222]-0.369: 5~

4. Results and Discussions
4.1. I)ipe Fracture Experiments

In the last decade, a nurnber of pipe fracture experiments have been performed fornuclear pipings with a surface crack and/or a through-wall crack as parts of piping
research programs such as Degraded Piping Program[5], IPIRG(International Piping
Integrity Research Group) Program[6], and Short Crack Prograrn[7].

Table 2 is a summary of 48 pipe fracture experirnents obtained from the above
programs. Experiment nurnber, material specification, pipe outside diarneter(OD), pipe wall
thickness(t), crack depth(a/t, ~6), crack length(el7t, %), yield stress(Sy), ultimate tensilestress(S~), reference strain(80), sttain hardening coefflcient(aL) and stiain hardening

exponent(n) given in Eq. (16), J-R curve coefficients(Jlc, C, m) from the relation of
J=Jlc+C(Aa)~, where the unit of Aa is [mm], design stress intensity(S~) from the ASME

Code Sec.m[1], pipe intemal pressure(p), test temperature(T), and maximum momentmeasured from the pipe fracture experiments(M~) are given in the table.
Thirteen surface crack experirnents for fdritic base metal, 2 for fefritic SAW weld

metal, 25 for austenitic base metal, and 4 for austenitic GTAW weld metal, 4 foraustenitic SAW weld metal were perfonned under four point bending load. The pipe
outside diameter ranged from 114.3 mrn(4.5 inches) to 711.2 mm(28 inches). Twentye)q~riments were perforrned with pipe internal pressure ranging from 1.55 MPa to 39.3

MPa. Test temperature ranged hom room temperature(21 'O to PWR operating
temperature(288 'O. The potential drop method was used to determine the crack growthamount.
Detailed contents including ex~mental facilities, e~imental methods, data processing,and crack evaluation methods are provided in the references [5-7].

4.2 Desirability of the SC.TNP method

8-7

Table 3 represents the maximum moment measured from the pipe fracture experiments
and the inaximum moments predicted from the SC.TNP method[14-16], the R6 method[8],and the DPFAD method[9-lO]. In the table, fracture ratio is also shown to compare theresults from the above three methods with the measured maximum moment. In order toconsider the pipe internal pressure effect, the fracture ratio(FR) is defmed as a stressterm, given by

apred+ o:p)

The aexp and op are the measured stress from experiments and the calculated stressinduced by pipe internal pressure respectively. The o~d is the predicted stress obtained

from the SC.TNP method, the R6 method, or the DPFAD method. The ae~b opretl, and ap
are given as

M R*
I
R~

The I is the moment of inertia given by(.

= )t R4 - R{}
4

For th(; ferritic pipings, the average fracture ratios obtained from the SC.TNP method,

the R6 method, and the DPFAD method are 0.99, I .42, and 1.42 respectively. For theaustenitic pipings, the average fracture ratios obtained from the SC.TNP method, the R6
method, and the DPFAD method are 1.lO, 1.39, and 1.31 respectively. This results showthat the SC.TNP method gives good results to predict the maximum load for both ferritic
and austenitic pipings. The well known R6 and DPFAD method give conservative results
by about 30-40%.

4.3 Desirability of new Z-Factors
The measured Z-Factors, Zmeas, from the experiments can be easily obtained from thepipe fracture experiments using the relation of

Mex p

8-8

where ML and Mexp are the limit moment calculated from Eqs. ( lO)-(13) and the maximum
moment measured from the pipe fracture experiment respectively.

The measured Z-Factors(Zmeas), the ASME Z-Factors (Zasme) obtained from Eqs.
(3)-(8), and new Z-Factors(Znew) obtained irom Eqs. (18)-(25) for 48 pipe fracturee~irnents are also given in Table 3.
The average values of Znew/Zmeas are 0.974 and O.~6 for fefritic pipings and austenitic
pipings respectively, while the total average values of Zasme/Zmeas are 1.350 and 1.133
for feftitic pipings and austenitic pipings respectively. This result shows that the new

Z-Factors obtained from the SC.TNP method are well in agreement with the measured
Z-Factors for both ferritic and austenitic pipings. However, it has to be pointed out thatthe new Z-Factors give slightly non-conservative results within about 496. Fig. 12 shows
the comparison of Zasme/Zmeas with Znew/Zmeas for the 48 pipe fracture experiments.

5. Conclusions
In this study, new Z-Factors to evaluate the behavior of a circumferential surface crackin nuclear pipings are developed by use of the SC.TNP method. The desira:bility of boththe SC.TNP method and the new Z-Factors are examined using the results from 48 pipe
fracture experiments for the nuclear pipings with a circumferential surface crack. The
conclusions of this study are as follows :( D The SC.TNP method is good for describing the circumferential surface crack behavior
in nuclear pipings.

(2) The new Z-Factors obtained from the SC.TNP method are well in agreement with themeasured Z-Factors for both feiritic and austenitic pipings.

Referencesl. ASME, Rules for Construction of Nuclear Power Plant Components, ASME Boiler andPressure Vessel Code, Section nl, Division 1, 1992.

2. ASME, Rules for Inservice Inspection of Nuclear Power Plant Components, ASlv~
Boiler and Pressure Vessel Code, Section XI, 1992.
3. D. M. Norris, et al., Evaluation of flaws in austenitic steel piping, EPRI Repoli~ Electtic

Power Research Institute, NP-46eo, 1986.
4. D. M. Norris, et al., Evaluation of flaws in fetTitic piping, EPRI Report, Electric Power
Research Institute, NP-6045, 1988.

5. G. M. Wilkowski, et al., Degraded piping program- Phase II, USNRC Repolt US
Nuclear Regiulatory Comrnission, NUREG/CR-4082, 1989.
6. G. M. Wilk. owski, et al., Intemational piping integrity research group - final report,
BMI Report, Battelle Memorial Institute, 1993.
7. G. M. Wilkowski, et al., Fracture behavior of short circumferentially suface-cracked

8-9pipe, USNRC Report, US Nuclear Regulatory Commission, NU~G/CR-6298, 1995.
8.

9.

I. Milne, et al., Assessment of the integrity of structures containing defects, CEGB

Report; CEGB, R/H~6-Rev 3, 1986.
J. M. Bloom, Validation of the deformation plasticity failure assessment diagram(DPFAD) approach - the case of an axial flaw in a pressurbed cylinder, Trans.
ASME, Journal of Pressure Vessel Technology ll2 (19eo) 213-217.

J. M. Bloom, Extension of the failure assessment diagram approach semi-elliptical fiaw
in pressuri~: cylinder, ASME winter amual meeting, 1983.
ll.
P. C. Paris, H. Tada. AL Zahoor, and H. Emst The theory of instability of the tearing
mode of elastic-plastic crack growth, ASTM STP 668, 1970, pp. 5-36.
12. J. W. Hutchinson and P. C. Paris, Stability analysis of J-controlled crack growth,lO.

ASTM STP 668, 1979, pp. 37-53.
13.

V. Kumer, et al., An engineedng approach for elastic-plastic fracture analysis, General

Electiic company corporate research and development EPRl Report, Electric Power
14.

Research Institute, NP-1931, 1981.
J. M. Bloom and S. N. Malik, Procedure for the assessment of the integrity of nuclear
pressure vessels and piping containing defetts, EPRl Report, Electric Power ResearchInstitute, NP-2431, 1982.

15.

P. M. Scott and J. Ahmad, Experirnental and analytical assessment of circumferentiallysurface-cracked pipes under bending, USNRC Report, US Nuclear Regulatory
Colrunission, NUREG/CR-4872, 1987.
16.

J. Aihmad, et al ., A Predictive J-estimation method for circumf erentiallysuface-crackod pipes of power-low hardening material in pure bending, in Advancedin Fracture and Fatigue For The 1990's - Load history effects of fracture resistance",

17.

Proceeding of ASME PVP Conference, Honolulu in Hawaii, 1989
S. Rahrnan, N. Ghadiali, D. Paul, and G. Wilkowski, Probabilistic pipe fractureevaluations for leak-rate-detection application, USNRC Report, US Nuclear Regulatory
Cornmission, NUREG/CR-6004, 1995.

Table I Reference Tensile Properties of Ferritic and Austenitic Pipings

8- lO

Table 2. Test Matrix of 48 Pipe Fracture Experiments for Ferritic and Austenitic Pipings

8- 11

T8ble 3.

Measured and Predicted Maximum Moments, Fracture Ratios, and Z-Facto,s for 48 Pipe Fracture

Experiments' : not caleulated

8- 12

fY

- F•,tl-~•U•-- f•,le8AWVWUd

I

I

b

~,
x
.p
p

Ndd

d-

,
o
o

ng. 1. Cross Section Geometry of Pipi:U with•osoFilg. a Refenence J-R Cwves for F~itic Base

Ciramfe~ltial Surface C~kM~l auf F~Titic SAW Weld Metal--- hN~SAWlf~Nd-A~t 8-5M•U

,

~Ls

10

4- d~.1
1

JF 'alO!S ~• af~J5 1- Qn'0.7S

0'8

e

~
~
;_ o'e

~-

~

,

= 0'4

~

2

o i2

o

o

10

~!-~

(o

so

o
o

02

0.4

0.8

o. 8

1

e'l

Filg. a RefiQ~ce J-R Cwves for Austeoitic BaseM~l andiAustenitic SAW Weld Metal

Flg. 4. Iinrit Mement far Faritic Base Metal

with OD=16 inches a~/t*lO)

J~ a~.1
1

e* 'tlO.S -1-_alM.7S-- •,~,sJ~ .~0.1

JF tll0.8~' aA'o.s

1- Qal0.7S - ASalE:

3'~ o'a

E

i:

=
~'

2

~ a'8

~

:E

~~

e 0'4

i oJa

1

o

o
o

02

ol

o.e

0.8

1

~,

o

o-

o.e

o. J'

0.8

1

e,,t

Fig. 5. I~xilrum Ma~t fcr F~ritic Base Wetal

with OD-16 inct~ (R~•lO)

8- 13

Fig. a Z-Factors for F~itic Base Metal with

OD•16 itnjw (RJt•10)

-~s,•)so ZC~~~ -- ~~,~mR

a~~U --- A~~~n•Ft

-~sf'

2

2

S
C1

'L

N

--g

.-1'-a--•1-1c,'d- ~~

o
o

10

20

OO

o
40

Rp Ol•,••:. b~1-

so

s

10

Flg. 7. Cemmjson of New Z-Factar ~leW-D

with ASME A value(Aaslne) fcr Fenitic
Base Metal

F~Titic Base Metal-~~a•,

20

Filg. 8. ~l~riscn of New A vaiue(Anew-D

with ASME Z-Faior ~a~~D fGr

3

IS
R,,"

o ZQ~~Od --- ZQ~~Ft-~~,•,

2azQ~,~--- Z~~,1•Rt

2

~

~1

*
~

t~

--G

.__..-,o-•1-

1-'-'-a.'ra~~

,~,

:!r

,,-

1

o

o

o

10

2O Ob,•K.
3O Il~
PW

4o

o

co

Fig. 9. Comparison of New Z-Factor ~aew-2)

2Oa•,-•:,
9Q t•tAp•

IO

Fetritic SAW Weld Metal

Austalitic Base Metala z(Nw~ .-- ZQ~Cu~R-z~alE'

5o

Flg. lO. Coa~rison of New Z-Factor (Znew-3)
with ASME Z-Factor (Z~smr3) forwith ASh~ Z-Factor (~smr2) for3

4Qazl•,,~,wIZ~Zr,w-

2

o
CP

~-,•••a~__..p~1-"oo ooL;ooP
ORDCP
d'
oco o

OO

.:

~1

o'

~

o

C~O CLO2

1o D lOllllllbjlril 'll,ll lllll ll ,lll, Il

l

o

o

o

10

20

t " 1" " " 1"'1"I"'1""I'l"III"'1"'1rl""II'l'll"'1'1'e("('('c('d'eaO

Jo

Rpa•1••:. I~-

E~lm.t ~lll~so

F~• Znew/Zmeas
Ia Comparison
for 48 Pipe Fractureof Zasl~neas andFlg. Il. Ccn~ison of New Z-Factar (Znew-4)

with ASN~ Z-Factor ~asmr4) for
Austenidc SAW Wdd Metal

Ex~iments

8- 14

9

Requirements for Pressure Boundary Integrity of Operating
Nuclear Plant-Japanese Standards
Technology)
H. Kobayashi* (Tokyo Institute ofPresented at International Workshop on the

Integrity of Nuclear Combonents,

May 7-9, l996 , Korea Institute of Nuclear

Safety, Taejon, Korea

N ew Maintenance Code for Operating

Nucl ear Power Plants in Japan (Draft)
by

I ida

(Japan Power Engineering and Inspection Corporation)

Kobayashi(Tokyo Institute of Technology)

Y.

Imamu ra(Mitsubishi Heavy Industries, Ltd.)K.

Hasegawa (Hitachi, Ltd.)K.

Abstract

As of the beginning of 1995, forty-eight (48) Iight water reactors (LWRs) are
operating in Japan, twenty-six (26) of which are boiling water reactors and twenty-

two (22) pressurized water reactors, and more are under planning or construction
phase. These LWRS constitute the main body of nuclear power generation units in Japan,

although there are three other power reactors of different kinds: a gas-cooled
reactor, an advanced thermal reactor and a fast breeder reactor. The oldest ones

among the LWRS were commissioned in 1970, experiencing more than twenty years ofoperation. Consideration for the increasing aged plants brings up an important
technological task to IRaintain and improve the reliability of operating nuclear

plants, which would be essential, technically and socially, for nuclear power
generatian in Japan to keep growing.

With these backgrounds, the Japan Power Engineering and Inspection Corporation(JAPEIC), under entrustment from the Hinistry of International Trade and Industry(MITI), constituted a technical committee, named the Committee on Nuclear Plant

Operation and Maintenance Standards (POMS Committee) in the fall of 1993, to develop
a maintenance code for operating nuclear power plants.

Slnce the Japanese law requlres nuclear power plants to undergo an annual
9- linspection, the plants are shut down annually as scheduled for two or three month

period during which refueling operation and inspection of the facility, systems andcomponents are carried out. The inspection activities include in-service inspection(ISI) in which the integrity of pressure retaining conponents is examined by non-destrYctive examination techniques and in-service testing (IST) in which systems andactive components are functionally examined. Some of the examinations are required
and sometimes attended by the government authorities, while the others are autonomous
ones based on the utility' s judgement. The JAPEIC' s POMS Committee is intended tocomprehensively examine annual inspection requirements which should be provided by

the government codes and/or private voluntary standards, and to construct a new
application system for the requirements to be fitted in. Due to the restriction of
the period given to the Committee for the first phase being until March, I996, theCommittee is concentrating its efforts on the area of ISI for the period. The
facilities, systems and components to be examined by the Committee activities arethose covered by MITI Notification N0.501, the technical standard for nuclear power
plants, which roughly correspond to Class l, 2, 3 and MC Components specified in ASME

Boiler and Pressure Vessel Code, Sections m and XI.
The new maintenance code (draft) comprises three parts, Non-Destructive
Examination and Inspection (NDE) Rules, Flaw Evaluation Rules and Repair/Replacement

Rules. This paper presents the summary of a draft of the'NDE Rules and Flaw
Evaluation Rules. Major differences betveen Japanese Code and ASME Code areemphasized.

Refferences"(1) Y. Imamura and K. Iida, Development of Nuclear Plant Operation and Maintenance

Code in Japan, 1995 ASME/JSHE Pressure Vessels and Piping Conference, Panel Session"A Look at Current Activities of ASME Section XI" , July 23-27, 1995, Honolulu,
Hawaii, USA.

(2) K. Iida and K. Hasegawa, Acceptance Standard for Pipes in New Maintenance Code(Draft) for LWR in Japan, Inquiry Number ISI-94-21, ASME Code Section XI SG Eval.
Stand., Harch 6, I996, Charlotte, NC, USA.

9-2

N ew Mainten ance Codefo r Oper at ing Nuc l ear Powe r Pl an tSin Japan (Draft)by

K. Iida(JAPEIC)

H. Kobayashi(TIT)

Y. Imamura(nH I )

K. Hasegawa(Hitachi)

New Malntenance Code in Japanis developed by the Committee on"Nuclear Plant Operation and Naintenance Standards in Japan'(l994 - I995, Chairman : Prof. K. Ilda)
The committee is organized by"The Japan Pover Engineering and Inspection Corporation (JAPEIC)"
under the sponsorship of"The Ninistry of International Trade and Industry (XITI)"

9-3

Design/Manufacture andMaintenance/Inspection Codes
for Nuclear Power Plants

ASMEBo iler &Pressure Vessel CodeSec. XI

Rules for Inservice Inspectionof Nuclear Power Plant Components~

Requirements for Pressure Boundary Integrityof Operating Nuclear Plants

9-4

Maintenance Codes in
Foreign and Domestic Countries

ISI : In-Service Inspection

NDE: Non-Destructive Examination and Inspectlon

New Malntenance Code in JapanR Non-Destructive Examination and Inspection Rules

R Flaw Evaluation Rules@ Repair/Replacement Rules

9-5Impact of Flaw Evaluation

Rules

O Design/Manufacture Code
Flaw Detection -' Repair/Replacenent

O ~aintenance/Inspection Code
Flaw Acceptance - Continued Operation

Mainten a nce Code for Operatin g Nuc lear
Po w er P lant s inJapan (Dr aft)

General Rules
Te rm ino I ogy

Inspection

Ru I es

Articlel General

Article2 Requ i reme nts

Article3 Requirements f o r Class 2 Colrponents ( JWB)

Article4 Requirements f o r Class 3 Conponents( JvC)

Article 5

ArticleRequirealents

for

for

( JWA)
C•lassComponentslClass 4 Colponents ( JVD)

6 Requirements f o r Colponent Suppo rt s ( JVF)

Evaluation Rules
Re pai r IRep I ac emet

R equ i rements

Rules

9-6

Inspe ction Interv al S and Progr am
Japanese Code = ASMECode Program B
,

~

~

~

Perc ent a g e of UT in I

SI (per 10 ye

Japanese CodeDV

Dissililar Velds

SD

Structural Discontinuity

GV

General Velds

9-7

~

ars)

Perc entage of UT in ISI (perASMECode

10 years)

Flaw Ev aluatlon Procedure
In-Service Inspection(Flav Detection/Sizing)
F I aw

Yes

Characterization

Acceptance Standards(Allovable Flaw Size)NOF I av

Grovth Evaluation

Fracture Evaluation

Acceptance Criteria(Critical Flav Size)

No

Yes

Acceptablefor Continued
Operation

Acceptable for Continued
Ope rat i on
during End-ofEvaluation Period

9-8Colponent
lust beRepaired/ReplacedFl aw Detectability/SizingAccuracy

by Ultrasonic Examination
Reference : Proving Tests by Nuclear Power Engineering Co.
ObjectFerrltlc Vessel Steels (t 50-250nm Clad & Unclad)
Ferritic Pipe Steels, Austenitic 'F I awPlpe Steels (t 5 = 5 Omm )~: Surface Planar Flavs

Detectability

Sizing Accuracy ( :i: 3a)

Ac cept an ce Standards (Allowable Flaw Size)

O Over Detectability by NDE
O Validation by Fracture Hechanics Evaluation
O Continuity between Ferritic Vessel and Pipe Standards
O Combined Ferritic and Austenitic Pipe Standards
O ~inimum Alldvable Flaw Depth a=1.5mm
O LiRit of Aspect Ratio a/ Q >0.06
O Subsurface Flaw a/t = Surface Flaw a/tx 1.04Y
9-9

Technlc a l Basi

Allowabl e Fl aw Size

s offor Pr~ e's s u ire Vesselso ASHE Code, Sec . nl
Postulated Flaw Size,a/t=25% for a/ Q =1/6

O ASME Code,Sec . XIAllowable Flaw Size,

for a/ Q =1/6, t=100-300mn

a/t=2 . 5%

Constant K for Conbinations of a/t

and a/ Ql

Al I ovabl e

x postulated

Flaw Size

F I awSize

10

O Over Detectability by
O Validation by Fracture

NDE -~ Bare uargin
Mechanics Evaluation -~ Sufficient

HarginAll owable Pl an ar Fl aws for Vessels

9- lO

t

a
F I awQ

a/1=0' S ~,

~EE
~
c~CL
~

16

o. 35

12
o. 2s

;,:'

C

O

JO

8

,:';

~
o. 15

4

O. 06

Wall Thickness, t (mm)Allowable Planar Flaws for Vessels

9-llInconsi stency of Allowable Flaw Sizes
for Pipes in ASME CodeOI n con s i stencyof Allowable Flav Sizes for Ferritic Vessels(V) and Pipes(P)a. > a+O Comparison of Allovable Flaw Sizes for Ferritic(F) and Austenitic(A) Pipes

F : Low Toughness, A : High Toughness

All owabl e Pla na r Flaw sf or Pi p es
Combined Ferritic

and Austenitic

9*12

Pipe Standards

Wall Thickness, t (inch)

3.0 4.0

1 .O 2.0

o

0.4
1 0.0

~:u

a/1=0.5~~

EE 8.00.3 '!

'lafl
o 9'u

c
a
o
Q

0.25

Q

~
'o

0.35

J:~

;

;,1

0.2 ,5E

LL

Zs 4.0

~

o
'1'

,9;

;

0.15

~2

o
0.06

2.0

1 20

1 oo

80

40

Wau Thickness, t (mm)Allowable Planar Flaws for Ferritic
and Austenitic Pipes
Wall Thickness, t (inch)
4.0

3.0

1 .O 2.0//
0.4

a/1 = O.5

1 0,0

~EE

~
co

/ / ./''8- 8,00.3 'o_// ./ /.~~a// .~/ '

l' /

~ 6,0
/

UL
J:(a

ll

// . /~ 4,0///~ l'/,~/V] / ~/1~7/ /

2.0

1~(l

/ .l
~//

an = o.ool'l'./' - ~
:/~

'

'

/

//

'' '

/

~
02 D:
2S

a/1 = 0.06

a';~

InseNice Examination~
0.1

Japanese Code--- Ferritc
errtPipec lpe)Austenitic Pipe

20

/
/ /

c~
~a

60

ASME Code80 1 Oo

1 20

Wall Thickness, t (mm)Comparison of Allowable Flaws
between Japanese and ASMECodes
9- 13

Who I e Constitut ion of

Evaluation

Ru I e s

q:)
I

:~

.4

Cooponents

1 Components

Accepta nc e Cr iteria

Fl aw Gr owt h Eva luat lon

9-15

SystenTransients

Expected
af te rConponent

Flav Discovery

must beRepai redEnd-of -Fatigue

Flav to'Crack

be EvaluatedGrowthadAnalysis

Evaluation
Periodor Replaced

No
a*>

10af

Yes

AcceptableNo
al > 2af

Yes

Flaw Size

Evaluation Period~OlInitiation

Initiation and
Sma I I est

K[. Curve

K:c 'Critical Flaw

Kl* Curve

Size for

Severest
Norl a lTransientOperation until
End-of -af~;

for Contiriued

Norual, Upset

Severest

Test Conditions

Accident

aL

Cond ition

ArrestSmallest

Fracture

Critical Flav

Mechanics

Size for

Analysis

Accident
Conditionsatnust Arrest
within 75% ofVall Thickness

Enhanced Nondestructive Examination

Flaw Eva luat lon Sequence of Pressure Ves sels (ASME Code)Fr ac ture Toughness of Ferrltlc Vessels
Initiation and Arrest Toughness (mpafm)
KI= = 36.49 + 22.79 exp[0.036(T - RTNDT)]
K[. = 29.45 + 13.68 exp[0.0261(T - RTNDT)]

Neutron lrradiation Embrittlement ('C)

RTNOT = [RTNOT]' + ARTNDT + n
Parent Xaterials

ARTNDT = [CF]f0.29-0,04 l'. f[CF] = -16 + 1210P + 215C. + 77fC.'Ni
Ve IdsARTNDT = [CF]f0.25-0.lo l.. f[CF] = 29 - 24Si - 61Ni + 30lfC.•Ni
f(x 10t9 n/cm2, E>1 NeV) = f. exp(-0.24a/25.4)

N = 2cr[ARTNDT]Fr ac ture Eva luati

9-17

on for Pipes

Initial

Flaw Sizeal

PipeStresses

Fatigue Crack G row th

Flaw G rowth Analysis

IGSCCG row thInspection~O

ReduceInspection
Interval('~

Normal Conditions
Emergency and
Faulted Conditions

Safety Factor

Interval

Final Flaw Size

Critical Flaw Sizeaf

a* : Normal Conditions
ao : Emergency andl63

Faulted Conditions

NO

Acceptance Criterion
a~
ao

Yes
Repair/Replace

F I awContinued Operation
Acceptable

Evaluation Sequenceof Pipes (ASMECode)Fracture Cr iteria andFracture Eva lu at ion of Pipes

Fracture Cr iteria for

9- 19

Aus tenitic Pipes

Limit Loa.d Evaluation for

Ferritic and Austenitic Pipes
TablesAllovable End-of-Evaluation Period Flaw Depth to Thickness Ratio for

R Circumferential Flaws, Normal Operating (Including Upset and Test)@ Circumferential Flavs, Emergency and Faulted Conditions@ Axial Flaws, Normal Operating (Including Upset and Test) Conditions@ Axial Flavs, Emergency and Faulted Conditions

Screeni ng Cr iteria for Ferrltlc Plpes
IS JlcAvai I ab I e?

Obtain Jlc frol ~-@Use Actual JI CG)

Des ign Values ofS. and S*

.2

Limit

Load

Screening Criteria
SC=X . ' /S. 'SC ~: O. 2

EPFN

9-20

-->:SC~l.8:: LEFX :ASHE Code(NO Considerationin Japanese Code)

Conditions

Elast lc-Plastic Fracture Toughness
of Ferritic Pipes
G) Actual JI= by Tests~) Lower Bounds of JI= Data

@ JI= = 1.296 CVN
JI= : kJ/m2, CVN : J(Lower Bounds), T~20~
I~ Circulferential Flavs :lAxial Flaws : JI=(A) = J*=(C)2

Elastic-Plastlc Fracture Mechanlcs
Evaluation for Ferritic and Austenltlc Plpes
O Circumferential Flaws
Tables of Limit Load Evaluation

P. + Pb + P./SF Z[P* ~ Pb + p./SF]~
Ferritic Pipes

Z = 0.2885 Iog OD + 0.9573
Austenitic Pipes

TIG, SNAV : Z = 0.292 Iog OD + 0.986

O Axial Flaws - Failure Assessment Diagraa

9-21

Failure Assessnient Diagratus
Ferrit'ic and Austenitic Pip

f or
es

R Ferritic Pipes, Circumferential and Axial Flaws@Austenitic Pipes, Circumferential Flaws
C Austenitic Pipes, Axial Flaws(Preparing)@Austenitic Cast Steel(Ferrite~~20%), Circumferential and(Preparing)

Future Problems
O Official Codificaion of Draft
XITI Notification

Japan Electric Association Code/Guide

O Construction of International Code
Unified Japanese and ASNE Codes

Revised Edition of ASXE Code

9-22

Axial Flaws

10 The Development Status of Mechanical
Nuclear Power Application in Korea

N.H. Kim*, J.S. Nah (KOPEC)

Com ponent Code forTHE DEVELOPMENT STATUS OFMECHANICAL COMPONENT CODE FOR NUCLEAPOWER APPLICATION IN KOREA
NAM-HA KIM*, JIN-soo NAH
Electric power mdustry code Development proiect'Korea power En9ineerin9 company, mc.

seouL Korea

ABSTRACTAs of the present date, the majority of the mechanical components
installed in Korea's Power Plants have been designed and manufactured inaccordance with various foreign codes, even though many engineers areaccustomed to the usage of the foreign codes, certain amount of problems
occur because of foreign language, different units of measurement, terminology,

and administrative requirements. Because of the difficulties experienced inapplying differing codes, the Korean engineering community has been putting
forth efforts in the development of the Korea Electric Power Industry Code,described KEPIC hereinafter, to resolve these problems.

This paper discusses the status of the present KEPIC - Mechanical
Components Codes and discusses future development of KEPIC-Mechanical.

10 - 1INTRODUCTION
In 1986, South Korea for the first time achieved a favorable balance-of-

payments ratio in foreign trade, and the favorable balance has increased
rapidly. This traditional agrarian country has transformed into a modemized,

urban, industrial nation. This industrial growth and enhanced manufacturing/
engineering ability has necessitated the development of the KEPIC.

Since early 1992, the engineering community has diligently pursued thedevelopment of the KEPIC to resolve numerous difficulties encountered in theimplementation of foreign code requirements. As a result of these efforts,
KEPIC for the areas of mechanical, electrical, structural, fire protection andquality assurance was published in November, 1995. The developed KEPIC-

Mechanical embraces the general methodology and principles of ASME B&PV

Code. This decision was based on our present and historical involvementwith the ASME Code as being the Construction Code of Record on many ofour projects. The KEPIC-Mechanical is divided into two Parts; namely, Level

I is the adaptation of the ASME Code, Section El:, Division I for steel

components in the Nuclear Steam Supply System (NSSS). Level H contains
the Balance 6f Plant (BOP) and Turbine/Generator systems. Table I provides

the frame of these two levels. This paper introduces ~ the developed
KEPIC-Mechanical for nuclear and non-nuclear components and discusses thescope of future development for KEPIC- Mechanical.

TABLE 1. FRAME OF KEPIC-MECHANICAL

l0-2

NUCLEAR COMPONENTS
Nuclear Component Code covers design and construction' rules for NSSS

mechanical components for Pressurized Water Reactor (PWR) Plants.
Safety-Related pressure retaining components, such as pressure vessels, piping,pumps, and valves fall within its jurisdiction. Nuclear Component Code iscomposed of the following three sections: 1) General Rules; 2) Component
Code; and 3) Component Service Code.

The General Requirements, Subsection MNA of Nuclear Component Codefollows the structural format and contents of the ASME Code, Section HI,
Division l, Subsection NCA: with the exception of concrete components, whichis part of KE;PIC-CiviVStructural Code.

General Re uirements
Classification of Components, MNA-2000, follows the guidelines ofSubsection NCA-2000. Responsibilities and Duties, MNA-3000, is a simplified

version of NCA-3000, such as owner, component manufacturers, material
manufacturers, component installers, and authorized inspection agencies
complies with the Korean Industrial System.

Quality Assurance, MNA-4000, meets the requirements of Subsection
NCA-4000. The General Requirements differs from Subsection NCA by theaddition of Documents, MNA-6000. This section delineates the appropriate

documentation and records required by variou s organizations in theconstruction of nuclear components. Design Specification and Design Reportshall be certified by Korean Mechanical Professional Engineer.

The Korea Electric Association will be responsible for the Certification of

Authorization in accordance with MNA-8000.

Table 2 shows a comparison between the General Requirements of MNA

and NCA.
lO - 3

TABLE 2. COMPARISON GENERALREQUIREMENTSOF

Com onent Code Nuclear
In the preparation of the Nuclear Component Code, an analogy wasdeliberately created between its structure and those of the ASME Code. Thiswas done to alleviate confusion with engineering disciplines working between

the two codes. As you see in Table 3 this was accomplished by the additionof "M" to each Subsectron of the ASME Code, Section lll, Division 1.

The structures are comparable to the ASME Code numbering system, of
which the articles, paragraphs, and the same reference principles are used.

The Subsections of Nuclear Component Codes are a direct correlation of the
ASME Code, Section Ill, Division 1, 1992 Edition. However, there does exist

some variances between the two, such as the use of American Standards
versus the use of Korean Standards. One example, is the caiibration of

instruments and equipment, is perforrned in accordance with the Korea
Standard Research Institute. ASME Code, Section 11 , V , and D( is directly
10 L 4

applied without translation to support the Nuclear Component Codes. Table 3
shows the full listing of the applicable Nuclear Component Code in Korea.

Table 3. CONTENTS OF NUCLEAR COMPONENTS CODENON-NUCLEAR COMPONENTS
The Non-Nuclear Component Code has evolved around the structure andcontents of the Korean Standards. The Non-Nuclear Component Codepresently focuses on the steam generating power facilities, but eventually will
include components for petro-chemical plants and refineries.

The Non-Nuclear Component Code is composed of two sections. The
primary section is the Component Code. Which delineates the general andtechnical requirements for materials, design, fabrication, examination, andtesting for construction. The secondary section is the Service Code, which

specifies the standards to be commonly used for manufacturing. Table 4
outlines the relationship of the Component Code and Service Code.
lo - 5

Table 4. Outline of Non-Nuclear Component CodeCom onent Code Non-Nuclear
The Component Code is classified into General Requirements (MGA) andTechnical Requirements (MGX). The format and contents follows the ASME
Section Lll Code.

General ReQuirements are in accordance with Korean Standard (KS) A 9000

Series. Quality System Standard, which were adapted from the ISO 9~
Series for the purpose of globalization. Table 5 shows the comparisons

between ASME Code, Section 111 , Subsection NCA and the General
Recluirements of MGA.

TABLE 5. COMPARISON OF GENERAL REQUIREMENTS

lO - 6

The philosophy of ASlvlE Section Vul, Division l, was used as a guideline

for the development of Non-Nuclear Components.

The contents of

Non-Nuclear Components are limited to welded structures. The following is asummary of Non-Nuclear Components Code:o The Article of Contents follows ASME Section HI, Subsection ND.
o Design requirements are expressed in SI units and the terminology of

Korean Standards.
o Material requirements for carbon steels, Iow alloy steels, and high alloysteels, which were used in Korean Plants.
o Welding processes are selected for specific fabrication reQuirements.

The contents of the Component Code are made, as a rule, and shown in
Table 6.

TABLE 6. CONTENTS OF THE NON-NUCLEAR COMPONENT CODE

Service Code Non-Nuclear
The S ervice Codes contain material specif ications, nondestructive
examinations, and welding Qualifications to pressure retaining parts as shownon Table 4.

Material Specifications are divided into four subsections, same as the ASME

Code, Section 11 , as follows: l0-7

o lvIDF - Ferrous Material Specificationso MDN - Nonferrous Material Specifications

o MDW - Specifications for Welding Materialo MDP - Properties
Material Specifications consists of 25 ferrous material specifications, 18
nonferrous material specifications, and 19 welding material specifications, whichwere mainly utilized in the construction of nuclear and thermal power plants.

The code will be supplemented to meet engineering needs and as technology
develops within our industry. In the determination of allowable stress values
for pressure parts, the maximum allowable stress values are established as thelowest value obtained from the criteria of ASME Code, Section H , Part D,
Appendix I "Non Mandatory Basrs for Establishing Stress Values."

Nondestructive examination (NDE) contains NDE methods and qualification

reQuirements of NDE personnel. NDE requirements and methods havebasically incorporated the ASME Code, Section V an:d the Korean Standards,
which are applied to the construction of power plant components.

For the preparation of welding Qualifications, ASME Section D(, Part QW,

was used as a guideline. However, forrnat and contents differ from the
ASME Code. Welding qualification incorporated the welding data, specified inPart QW, into the applicable paragraphs of the welding procedure and welding

perforrnance qualification. This change was done to provide convenience forthe end user.

Qualification of welding materials follows the general

reqrurements of AWS A5.01, "Filler Metal Procurement Guidelines."

DISCUSSION
At present, there exist two component codes for Pressure retaining parts;()ne for nuclear and the other for non-nuclear. It is our future plan to merge

Lhe two into one major Pressure component code in which one Service Codewill be applicable for both. This will be done with a conscientious effort by
10 - 8the Korean Mechanical Engineers. The main objective and reasoning for thisconsolidation is to bring within the engineering realm a more user friendly

code while delineating the technical requirements of both.

The

KEPIC-Mechanical will evolve through our engineering community ' sexperience and to be paced with the continuing technical betterments
conceming design, construction, and operation of the mechanical components.

It is now being developed for other Mechanical Component Codes such as

Turbine Generator, HVAC, Crane including addional Pressure Components
Code. Table 7 shows outline of the Code in Preparation.

TABLE 7. OUTLINE OF KEPIC-MECHANICAL IN PREPARATION

10 - 9

KE:PIC-lIECH JWICAL
ATTAcmENT. LIST OFPRESENT

l0-10

SESSION C

ll Progress of Component

Program at JAERIK. Shibata* (JAERI)

Aging and Structural

Integrity

Research

Pr~ess of Canlment Ageing and Structura l
Int~ity Rtsea:th Prognl at JAERI
Katsuyuki SHIBATA
Reactor Component Reliability Laboratory, Departnent of Reactor
Safety ResearclL Japan Atomic Energy Research Institute,
Tokai~nura~ Naka-guT~ Ibaraki-ken, 319-ll Japan

Abstract

At JAERI, the safety research of nuclear facilities is conducted in accordancewith the Fiv~Year Safety Research Program of the Nuclear ~fety CommissionL
Ageing and st~uctural integrity research of JAERI has been oonducted as one of
the safety research prescribed in this Program. From the viev points of plantsafety and replacenent technology, Reactor Pressure Vessel. Concrete structure andElectrical cables were identified as important key components to be investigatedin the ageing research This paper overvievs the progress and results related toageing and structural integrity research currently performe(Ll. Piping Reliability Test progr~[l]An extensive piping research program had been conducted to demonstrate thesafety and reliability of LVR piping [l]. Three major tests, Pipe fatigue test. LBB

test and pipe rupture test, vere performet
Pipe fatigue test vas perfonned to prove the integrity of primary piping

during the service period(Fi& D. Flat plate specimens vith one or tvo surface
cracks were teste(L From the test results a procedure to predict the multiple
fatigue crack grovth were proposecL To verify this procedure, straight and bend
pipes vith multiple cracks were testedL It vas confil~ed that the LVR pipings

vith a cra~k grovn during service period have sufficient fracture resistance(Fig 2). _LBB test, vhich involves the fracture test of circunferentially cracked piping
under hending and the I eak-rate test through fatigue cracked pipin& vas perfor-

med
to exaJnine the LBB of piping(Fi~ 3-Fi& 6). It •apipings can be jutified for pipes with 4-inch diameter or larger(Fi~ 7, Table D.
In additiorL Pipe ruptdre test was carried out to denonstrate the effective-

ness of protective device against postulated pipe rupture.
ll-l

2 Peserch Prograln on Integrity of Aged LVR oomponents
In the Five-year Safety Research Program of NSC for the period ~96 -~OO. theresearch area related to the ageing is identified to be one of the important
issue. JAERI has oonducted the research on this area since late 1980~s. From theviev points of plant safety and repare/replacement technology, Reactor Pressure
Vessel(RPV), Concrete structure and Electrical cables were identified as key components to be investigate(LObjectives of the research is to improve and validate the prediction methodology, to evaluate the integrity of aged components and to utilize the results on

regulatory issue. Subjects under investigation are described belov. Host activities are concerned with the ageing issue of RPV.

l) Ageing mechanism and prediction methodology(D Irradiation embrittlement of RPV• Irradiation tests using J~TR(Japan ~aterial Testing R
Fi& 9)

Fluence. Flux, Temperature Chemical compositions [2]• Evaluationof fracture toughness of irradiated PV stee
Correlation between mechanical properties [3]• tionInvestigationof embrittlement of decommissioned
Reactor) RPV [4] (Fi& 10, Fi& ID
liechanical and metallurgical tests of trepans cut from JPDR RPV is underway under th.e cooperation with ORNL.

(2) Enviromuentally assisted cracking of RPV• Investigation of JPDR RPV cracking• Corrosion fatigue test in high temperature water envir
Influence of temperature, flow velocity, DO, crack branchinng phenomena(3) Ageing of concrete structures• Investigation of ageing of the biological shield wall of• Investigation of ageing of the JPDR contairunent ves
2) Inspection and evaluation of ageing• magnetic
Improvedcrack detection technique by detecting thflux of ECT [7] (Fi& 12 -Fi& 14)• Detection of material degradation by llli(~agnetic Inter• Reconstitution technique on irradiated Charpy impa
(Surface Activated Joinig) (Cooperation with IHD [9] (Fi& 15-Fi& 17)

3) Evaluation methodology for component integrity• Integrity evaluation of an aged RPV based on EPFX
• Reliability analysiall-2of an aged RPV based on proba

Pef erences[l] K. ~ribata T. Isozaki et al. : Results of reliability test program on lightvater reactor pipin& Nucl. En~ Des. , 153, 1994

[2] IL Suzuki and Y. Idei : Effects of neutron flux and irradiation temperatureon

irradiation embrittlement of A533B steels, Presented at Asu Symposium on
Effects of Radiation on laterials, 1994, Sun Valley (to be published in AST~

STP127D[3] K. Onizawa and li. Suznki : Fracture touglmess evaluation in the Transition

Region of reactor pressure vessel steel, IAEA specialists meeting on irradiation embrittlement and mitigation, Oct. 1995, Espoo, Finland[4] IL Suzuki. Y. Idei, IC Onizawa : Investigation on irradiation embrittlement ofreactor pressure vessel steel using decommissioned JPDR, IAEA Specialists
meeting on Techrology for Lifetime lianagement of NPP, Nov. 1994, Tokyo[5] .N. Nakajima and N. Ebine:Effect of local enviromuental condition on crack

growth behavior in RPV steels of elevated temperature water enviromuent, IAEA
Specialists meeting on Technology for Lifetime ~nagement of NPP. Nov. 1994
[6] Y. Idei et al. :lechanical properties of JPDR biological shield concrete,
JAERI-1 90-205, 1990

[7] N. Ebine, S. Takeuchi and K. Ara :Eddy current testing using a parallelogramprobe of the distorted magnetic flux detection type. J. Hagnetics Society ofJapan. Vol. 20, No. 2, 1996

[8] K. Ara N. Nakajima and N. Ebine : A new method of nondestractive measurement forassessment of material degradation of aged reactor pressure vessels, Presen-

ted at Japan~entral Europ Joint Workshop on Advanced Compiting in Engioneerin& 1994, Poland[9] ~ Onizawa ~ FukayEL Y. Nishiyama et al. : Development of reconstitution techni-que of Charpy impact specimens by surface-activated j oining for reactor

pressure vessel surveillance, IAEA Specialists meeting on Technology for
Lifetime lanagement of NPP, Nov. 1994, Tokyo

ll-3

Fi& lF I owof the pipe

fat i gue

E 3.S

test

3. 5

cl)

\CL

JL4-0

.-~ .~.

.~Ll:stress of piping
Fracture stres{with a fatigue crack after+E: 3.0
3. o

,L

~J

40 Years operationo_ L5

Ai I owab I e

LS4'

cond i t i or~stress for Level C
Z 25Sat

1'

.~ LO

Lo

~.
o-

zlS

A I I owab I e

st ress

cond i t i or~

1 . 5Smfo r

Level 8L5PW~ PCL
1. O

Lo

O

o

o

4

Fi& 2

8 i2D i ame te r ( i nch)20t8

24 32 e28

Pm ' 1.5~l'
4

Fracture stress of pipjng with a fatigue crack(BWR PrimarYLoo p Reci rcul8tion Line and PWR

ll-4

o sch 160
8le12
D i ame te r

20
( i nch)24

after 4O years ope rat i oo
Pr imary Coo I ant

L i ne)2,

a2

Fig3 Flow of L8B verification

1 .a

test

STS 42
C8rbal 'teel pipet .1

Typc 304
Stainle$s

123-inch dilL 8•se Ntal. RT
A 6-indl di,L B1lse •etil. RT

stee I .

D12-inct,'di~e•$e Nt81. RT
I12-inch diL Veld l'tal". RT' GTA," SULWel2-i,~hdit B••e Ntil. 300'C
1)12-inch dit 'etd Ntal". 300'C

1-1)12-inch di~ tetd Nt•t". 3001C

\

:~

~

:~F::i.v~LC~o 0.8GLc r'ck

c'

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Nc :Col 18pse NoentHe :Col lapse a'ooent~p-e•eMp.
o. 2

Hc :Col llpse 'oent
H. :Col lapse 'DNnt

Aof Uncr8cked PiPe
o

o

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PIO
p-cJ8eMp.

of uncncked pipeO• 0.4 O.S OJ,
UD

o

O•
0.4
0.8
0.8
Ccmpar i soo between net-s•ct ion col llpse
Nofll iz•d cflck enel• ze IZ•r

Fi~ 4

ThraJtr•1 1 l
cr•dc

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,,,

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V 6-irdt di• 8•se Ntat 2851C Irl

IP

V 6-inch di•.8ase Ntll 2851C U'Io~ll2-inch di~'eld Ntll". RT~l2-irNh dil88se Ntll. 3001C:1 1.0'c.
o 0.8

D12-inct, dilL 88se ll'ta . RT

"S"At

JL 6-inch ditteld Ntal' .RT-,nch ,t8ase Nt81.R

A
6-itdl dit8Qse •.tll.RT
A 6-irdl diILVeld ll'tll".RToIn

NorNli:•d cr•de angle 2 e /Zlr

Fi& 5 Canpirison bet••en net-sect ion col lap
critericn
8nd te•t results of throughl,llt
cracked stainless steel pipes
criterion 8nd test results of through-•a
cracked c•rbon steel pipes

ll-5

2OQ

'i,.

'~

~

' !~' j '/~'~!'~;'-~~
i
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e

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C~

to

co

8ending stress ' b (IPa)

Prdieted le8k rate Cca/Iin)

Fi&6 Prediction of

COAand l eak

120

rate

1 20

\.

.Through~'811 crack anole,2e .,i, at the oQset of,fr8cture under PlttPb'L 5Sd. (P'design pressure)Throoghl811 crlck anGle,2e ,t't it the ort$et ol
fr8cture u'Ider. p0+pbll. 5Sa (pldesief' Pres$Ure)~)-o / O SctL 8aloe~~~-O~\o

o
c\'
e 80

/

, oo

e sctL 160

80'\ ~~ ~b,b \1~ac,"*o 60

50

\

:
J:

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:,

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C

,-

/

Le8k-detect8ble crack 8n9ie. 2 e i..t
8t 20 l/oin le8k rlte undef stre$s
Pb'G. 5Sa 8r'd operttiog pressuresure

20

~\
/ .\

de

. .\\ ~\

Lelk-detect,bl, cract 8nelq ze t"I 'at 20 Illi,1 Iellk rate nder stressH)~. 5Sa ind opl'rating pressuresure20

B~I~PLRop,~lu

8 1Di80Gt~r
2 1 c(inch)20 24 28

o

o

32

a

,

(a)PrimarY Loop Reci rculation LineFi& 7 Case studies of LBB

Table 1

Sumna r y

' t2Di80cter
18 20
24(inch)(b)PrimarY Coolant Linein pressure boundarY Pipings of BWR and PWR

results of piping

rel iabi I i tY

ll-6

test

28

32

Table 2

auieal colpcsiticn of the laterials usecL

(a)

(b)

300

300

z80

250

:Boo

zoo

150

150

lOO

lOO

80

60

E;

~h

(rt,5)

~e~

oao

15

~
~

~

sso

~o

-lOO

o

60

Test Temperature~50

lOQ 150

-lOO

('O

~ro
o
50
Test TeHLperatureroo 150('O

Absorbed energy transition curves before and after
irradiation to lxl018 and 2xl019n/c~2t (E>lMeV) at
290'C for (a) alloy L and (b) alloy tl. All data
are of dose rate of -lOl:an/aaa/s.

Fi& 8

120

'~ IOO
~)
~'
~,

ttd
.,.

co

.::I

CQ

.$
-cl'

co

40

.E~

20

~o

260

270

280

lrradiation temperature (1O
Fig g

~O

800

•Effects
irradiation
temperatureon ( a ) tr
tenperature shiftof,afterirradiatiOn to -2xlO:L9n/cm[1 Suzuki et al. :Effect of neutron flux and irradiation
telperature on irradiation embrittlement of A533B steels,
Presented at ASTI sylposiul. 1994, Sun Val ley(to be published in ASTK STP 127D

ll -7

J

V-no t ch
(F I ycu t )\...~-.' t; RB-1 :: --- B-. - - ' Overlay aaddh9c87/

Ch8 r py

SPec il'en u' I I tne\lHack-sa' \ R8_1-lU

RBRe_ t _2

L

'3
RB_
I~4~
l

R0-1

t

fl

RB-l

CbddingCut t e r

B

(C8rbon blade)i'lL2~~~LB

\('

l

/f

\2,l

TrepanFi& lOS-~

RB -1- 1oOverlay. to

~

U

tl'

s3_6s

Tens i I e

SPecilnen Lithingfrom Trepan cU t from JPDR RPVSpec i men machining procedure

03U

200
Expe r i menta lbresul ts expected

190

from JPDR Rrv i nvest i gat i on1) Flux effects on Radiation
Erbr i tt I ement-Ver i f i voat i on

,~
It,~

of Trend Curves

~
o
1o
~
~

~e

2) lhrough-thickness Attenuat i onof I rradiation Embrittlement

180

~~(b

o~~Ct9ooo
oo
o~)

c5

3) therma I Age i ng- I r rad i ation

::l
I=,

Embrlttlement Interaction4) Ameal ing Characteristics ofReal RPV material

(()

oo

170

A

. ..

160

•' after
as500Ctrepanned
anuealx lh[~. Suzuki and Y. Idei : Study on Through-thichness

150oAttenuation of lrradiation Embrittlement using

20

40

60

80

JPDR Pressure Vessel, JAERI-Res, 94-038, 1994 lDistance frem Inner Surface (mm)Fi& Illl -8

Through-thi;ckness measurement ofhardness befoTe and after Anneal ing

Scanni ng
di recti onzexciting coilY
P.X'~~~:1:.
":•'1~ .,,t:,, J:~,' ...
'~~ ~ '::

metal plateHp(a) excitatien by coil

H.;H.

X, y

ma. gnetl c(b) magnetc fleld distribution

Fi& 12 Principle of eddy current probe
detecting distorted mayletic flux

'~
'~E

>

i

0.5

0.4
0.3

a;
OD

S 0.2
-o
:~

dJ

=

a

+'

g

verti cal d~fferentlal pi ck~JP Coil
(90T-90T)x4 per coi 1

Fig 13 Geometry of parallelogra!n eddy
current probe

Outer holeposi tion Y=15mml~ . 5A p-p

f:20kHz-F c=1 .6(mm)-1- c=1 .O

(m)

•o. 1
o

20

40 60

Hole depth (%)
Fig 14 Pelation hetveen detected voltage and depth of outer surface flav
Dl. Ebine, et al. :Eddy current testing using a parallelograu probe of the distorted magnetic flux detection type, J. Iagnetics Society of Japaa Vol. 20, 1996]

ll-9

Contarninaton: Frlction force,

Polished Oxide flm, Ion, so on

partlele Absorbed 9as, Clean surface --: ~3 ~:, ,,,Surface- Activated Joined

Generic

SUrtace

ConCttion

Activation SUrface
pOuShing Surtace
Condition Treatment

Fi& 15Pr i nc i p I e~.

~~_ ::~. ::-~~_ :_~I

Jo i n i ng)of SAJ(Surface Act i vated

2L
h!$1

Broken Charpy

C~i~ C~] u~eto'~~n l~~rdat~ i~~{~:""'~~'~~ ' S N

Fu,tre VJ\btdl

800'~ 700
~-

~~]

:~

~

~

~ B~tab

~400

'OOO

Jolned'SAJ -Jol ned,l~vn~ r•~ttaect!u be'm vddll'
b7 Ila'IttJBel' t Hofmalnl"(ASTN STP u2S]
/ JIL!e Itld veldtaa,lby petrin et al,:

500

~300

'~ )!PRI I'P':tne)
by l'a8abta ct dl:cD••. ~;l I,aoer lr'Idia'

.___:
_ ...
~L8Tldl 8TP 12OC]"' _~__i:'i"/.v-:---------------------~.;.'. '-"v:-~-------------------'-----------.......tbe ,tlld vcldia,
b7 V,Lle t JLm'uudc'-•~:;1,./..aLSlr)( STP IZ04]

iOO

La

t

Lraffected !?~~,..~Y..SA,!.......

q~

Reconstituted
Reconst I tut I on
Fi& 16Char p y
P rocedu r e

~

lr.

~ 200

~

~a'lp bN~

8

D~t f' hterf"a L ()Fi& 17 Comparison wi th other method- Iarg6 unaffected zone by SAJ[K. Ohizawa, et. al. : Development of Reconstitution

Technique of Charpy Impact Specimens,
IAEA specialists meeting, Nov. 1994. Tokyo

11 - lOJ

12 Development of Expert System for Nuclear Piping Integrity
Y.J. Kim*, M.W. Suh, C.S. Seok, H.K. Jun(SungKyunKwan Univ.) Y.W. Park, Y.H. Choi, J.B. Lee (KINS)DEVELOPMENT OF EXPERT SYSTEM FOR
NUCLEAR PIPING INTEGRITY

Young-Jin Kim , Myung-Won Suh, Chang-Sung Seok. Hyun-Kyu Jun
Dept. of Mechanlcal Engineering

Sung Kyun Kwan Universlty
Suwon, Korea

Youn-Won Park, Young-Hwan Choi, Jeong-Bae Lee
Korea Instrtute of Nudear Safety

Daeion, Korea

A I ;_~ l:' R.,rL C'I'

Int:,•rity. 'I':,;~, l) Iper descnbes t]le tmt It'l~ ~nct ,.he ~,,•velopme_nt slrateg~l7jl': ~y~','m
consists of 5 pa]'ts; u~,c',~ inl.*?]-r,:t'e. duLabas:]. cnou ledge base, e\1)ert 2lnd i! "~rit]~' i.tc~~cloped, *'ef e' tl*ely in the database pfu't, n].!"le.'**' pioing n_laterial p,*)perties are stored an(: l,tc '.1*t'(Jtc.,,vn

lr:L't'rial properties are provided tl]rol'(J'i• inferring t-jlf' known material prol)el'or illf, rr;,1gf InaLerial pn)perties aT(: t,ttJ,t( m I t (]lowlcdg~ , Ise part Th most a' I ( f"I 'r;~a],]'-- ion Inethorl for given input c,~,ndttion Is~ 're('on'meltded in the expert part. Fi,1:tl;, . .Il'.;

integrity evrtluation part is develol~cl for the evaluaLion of piping integrity effectively.

INTRODUCTION
The expert system, as a paJ1 of artificial intelligence (AD, is being applied for the ptuvose ofdesign and diagnostic analysis in a variety of engineering fields.

In nuclear industry,a~)plications of AI technology to the integrity assess]nent of power plant facilities haveprogressed from carly 80's and several softwares such as DIAS (Oka]mxlo et al, 1987), ESR(Jovanovic et al 1991) anci RAMliNO (Lucia and Voita, 1991) have bcen cleveloped.

Potcnti•al loss of structural integrity due to aging of nuclear piping
offcct on the safety of nuclear power plants. Currently nuclear piping integrity evaluation ist)crfonncd ror clcfccts found during in-scrvicc inspection. Thc integrity cvaluation rc(tuires not
only' cngineering kno~vledge but also comprehensive juclgement basecl on the rielcl experience.

In {1 inajority of integrity evaluation softwares currently being used in nuclcar industry, theintcgrity cvalLtation is usually performed after tl]e m*aterial properties and tl]c ovaluation methtxl~Lrc t)rovidcd by' tl]c user. Howevcr the inaterial properties sucl] as fractLtrc resis~tancc cLtrves andrtill _strcss-s_ train cunv'cs' are not al~vays al"ailable for early-built nuclear p()~vcr t)lilnts. 'I'l]crefore

it is~ uscfttl to tl]c Ltscr if unknovvn material propcrties can he inferred from knobvn materialt)roperties. In addition, althougll tl]c sclection of tl]e intcgrity evaluation mQth()d I]as n sit(nific:tnt

offcc. 't (]n thc itccuracy of atnalysis rcsults, it is not Qasy for non-ox])crt u_scrs t() cl]()oso tl]c
12 - lapprol)riate meth(xl for given information.

'I'l]e objective of tl]is pal)er is to clevelol) the expert system cailed ~*PiES (~uclear Piping

Intcgrity Expert Syste]n) for nuclear piping integrity. Tl]is papcr doscribcs the structure anddcvclopmcnt stratcgies of :\;PiES sys~te]n.

K"q'uDGe 8As!
EVALUAT*O" PARTFIGURE l. S rRUC'I'lJRF_ Ot' NPil:iS ~YS'PEM~. -I I~tLre I ~']()ws the structure of NPlhS system ~~PIEc~ ~ysLem i,~ developed for PC (pers,~nai
omputer) and consists of 5 parts; user interface, (iat;ibase, knowledge base, expert and integrityc~aiuation part. Those are illustrated as followings.

User intcrface

The user interface part is developed under Winclol;vs 3.1 operating system environment toconnect thc user and N_ 'PiES syste]n effectively. Pull-down menu driven method is adopted andfrc(iucntiy used mcnu is developed as toolbar for user's convenience,:.

Databases_'teels sucl] as S_•\312 Tl)304. SA312 Tp316, S/\l06 Gr.C e()btaincd from domestic material test program such as Yongg~vang 31!~;'4 (Ki]n et al, 1993) withadditional infonnation obtained from literature survey. Various material constants such as
RiunbcTg-Osgood paramcters are dircctly pro~lidcd, or can bc ()t)Laincd by curvc-ritting therlt\~' dilLit f()r l)ropcr mnge. .'\n effcctivc ciatitt]•a_sc managemcnt s12-2

u"' ~~-'~OIN'~-u 4~~1 "~d r'~'t Y~o'--' * "' i'~' by *~ ?
o'A'I I 'KI '~-' +~ .."A" "tid r'~*~, y_: ~/L'~-u,._~~:IcT ~)(' ~~,~ t- '' DIN'8-K8w?

: ,,o'p' o'Jler j
f'L--~ il~5,u~~

.

[ -•
J-1 '

\~iJ

, ~.b"FI( URE 2. STRUCTURE OI•' IIYBRID DATAB
Figlt,;e 2 slioll s the structure of the ciaulh~Ise vvhich i~ devt'loped Ltsi]]g f,l]e hybrid datiaba~cfor]nat. .'\n ef[ective search for material property is 1)ossible by )_Ising the hybrid databasewhich !; or~ xi L' Ith DBMS (datl)ase in Ln~L.rt nent sy,;:cl l) ?tn•!
managc!'nt llt s, '~tcn]).

DBMS is Ittililic'd to scarch necessary matenal l)rol)ert/ from the $'tol*d m Itenal l)ro,)erty If
nec.'ssary m;il:.rlal plope:ly js not ioLuld in tl]e dtltabas,', KB~lS i\ '.Itilized :u prodict unkJ,(,v,'(Lmaterial propt-rty from the knowleclge which is stored in the knu'vledgt: l]ase. In additio]1, thelower bound tensile and fracture toughness values which are provided in tl]e EPRI report (LNorris
et al, 1988) are stored to compare with the predicted material property.

In the kno~vledge base, the following levels of knowledge are stored.

I.evel I : Knobvledge for inferring a-e curve from yicld strength

Level 2 : Knovvledgc for inferring o-s curve fromyield strength and tensilc strengtt].

I e,,el 3 : Knolvledge for inferring J-R curvc fromfracture strain.

Det uls of thc kno~~ Iedge base u 111 be expl uned In the following cl]apter.

E~(pert p_art

In tl]c cxpert part, prop(,r intcgrity cvaluirtion motl](xl for given input conditions is'
rccommendcd bascd on citl]er tl]e posscssion ratio of matcriai pro~rty or tl]c critcria givcn inAppcndix C and Appcndix H of AS\_ II: Scc. XI (AS.'_\ll:, 1992). Detilils ~vill he oxL)laincd in tl]of()llo~ving chaptor.

12 - 3

In teglity _ ~~y_aluatiolL Qarl

In tl]c evaiuation l)art, foLtr cvaluation modules such as LEF\. I (1_incar Elastic Fracturc\_ Icchanics_'), EI)F\. I (Elastic I)lastic Fracture \_ Icchanics), Ii]nit l0ad method and fatiguc analysis arcpn)videci for the l)urpose of piping integrity el"alLtation. Anct the EPFM morlule consists of throel)arts such as CDFD (Crack Driving Force Diagram), J/T and DPF_F\D (Deformation Plasticity
Failure Assessment Diagra]n).

_\PiES system is progra]n]ned by using the C~~ Ianguage which suptx)rts OOI) (Ot]Ject Oncnteci
Progra]nming) and is modulizecl for tl]e sake of easy expansion of the system.
The following analysis models are currently available in l~~PiES syste]n.

- .'\xiai surface crack uncler intemal pressure- Circumferential surfacc. crack L]nder tension- Circumferential througl]-wall crack uncier
tcnsi()n- Circumferential through- \~ all! crack undcr}'f)r~l:iT~g moment- C Ir'._'unlie*rential through-YLali crack underbcndi]Tg: r,.()ment anJ tensl "- Circunlrercntial surf,lc(' cr;1ck under ben.:il]g n.10nlt~n'- Circumferential surface ~ rack tmder t.ension and bending moment- T .ong axidJ; suri'ace crack In(ler inter']~tl r,ressure- A\'ia] through-~vail crack ur'(le]' intcrnal l)ressure

INFERRING UNKNOWN MATERIAL PROPERTIES
Although the tensile properties of nuclear grade steels are easily found in open literature, thematerial constants Cl. C2 of J-R curve are selclo]n founcl. The following empirical equation was

obtained by statistically correlating the tensile data and the J-R data of SA312 TP304 and SAI06
Gr.C (Kim et al, 1995).

Cl =1.687r~k(S,, Y*'* = " (1))..*_ = O.38 for carbon stcel' [ O.(;1 for stainless steel~~"herc E is Young's modulus, k is non-dimensional constant, S',, is defined as S,/E and 8,, isfracture strain.

In this rescarcl], tl]c prcdiction of J-R cur/e from tcnsile claLa (Sy, E, 8,.) vvas pcrfonuod. nndtl]e accuracy of nbovc cquati()ns is obL{tino(1 as 7O t~~). 1\ccordingly, vvc define tl]c certainty fiLct()r()f rulc. (Cl:tttl.E) as 0.7 to c:llculntc tl]c p()sscssion rzttio of matcrial property. 'I'hc procodurc for
cillcukLLing tl]e p()sses_'si()n nlti() ~v'ill t)c cxplilincd latcr.

~Vl]cn tllc a-c cur\'o is unkn(]~vn, Ritmt]crg-Osgo(xl cons~Lant n valuc is ohtaincd ~Is (ISl()()in ttnd
~ktlik, I~)82):

12 - 4

l,n _ _i/n _ _ I "~~L ___2+1183____

S_1' 1.002+S'.1'/E ~ lil(1.002+S'_1'/[~•-S~l( I .OO2 + eo) = ( n/ 2.7183 In (1 .002 +eu)} (4)ancl the a value is obt'aincd asl In 1 2+S1' E (11") l"(( Sy/ E) ( I . 002 + S_v/ E) f

E n
- r 2.7183nS_1'
Sv(!\ Su (6))

J;- in(1 002 +eo) I 002 eo
eu( I . 002 + eo) "The prcdiction of o 8 cun/e fro]n tcnslle clata (Sy Su E) VF as perfonnecl and tl]e l( cur lcy (]rabove equ'ations is obtained as 90 (/o. .'\ccoi'dingly, we clefir.t::* thb CFRt'LE asb 0.9.

In a(Idition, various rulcs_ fcr infen~ing .JIC and Klc valt]es from tensile c[ata are Llst, stored In tl ekn(,bvleciga' basf, .

SELECTlor~ Ol' EVA]~ .U~ Arl'lON~'~Flli'rli(.L,' ' ~_ .~:ecttxl 1[tts'ed on tl]e 'iritcria glv':ll
An nt•i'l ~n"tc punng Integn'y e\alu rtlon
metllocl Is usuall)materia! property. This criteria can only be used when all infonnations are available. In thispaper the "possession ratio (PR) of material property" is proposed a_~; a nevv criteria for selecting

the most appropriate integrity evaluation method and the weight factor which is quantitatively
determined througl] the sensitivity analysis is adopted to eva]uate tl]e effect of material property
on fracture behavior.

v
In order to dctcnninc thc most appropriate evaluation metl]orl for given input information, ncw
methor.1 basect on tl]c PR v~'I]icl] is clefined as a ratio of given material propcrty and requircd

material property. The procedure for obtaining the PR can be cxplained as follovving.

Step I : Detcrmine tl]c scatter band of the material property from the clatabase.
Step 2 : Perfonn intcgrity cl/alu'ation for a long axial surface crack in a ph)c (Figurc 3) ~vith_--_ 2O, ~ 4Ot~6 variation in Inaterial property.

Step 3 : Dctermine tl]c bvcigl]t factor for rcspectivc matcrial propcrty rrom thc Intcgrtty
cvaluati()n rcsults.

Stcp 4 : Dotcrminc tl]c l)R from thc _summation of vveight ractors ror gil/on miltQriill l)ropcrty.

Scnsitiyity_J\nalysis _l.i.__._m_it_load_ l']ethod. 'I'hc liinit load for il long axinl surfncc cnlck in it l)i,)e, ns slu)~~ l] tn I Ig'urc:S, is cx,)ros'sed zts f,)II()vving (7,ti]o,)r, 1989).

12 - 5

long *xiai erackl"IGURE 3. GEOMETRY OI,' A LONG AXIAL CRACK

Pl. = SA ( I -x)/( I -x/1Vf2)} (8)bl' ; Iere

.if =\ I +(1 6114Rt)!
* - /,'\'f= flol,1' str/*'\~'(;over]1 r,~;~ InaLeri:ll property in the linlit load mctllod is flo~\' sJ~c:' s ',\'hlci] i'; dcfi]1 (. ':s :'.ber..*.

.* .

valuc oi' I :eid strength ('~,) and ultimate tensile strength (S,,). Flalure sll()\,~$ rhe I Inatlon ofli]nit loacl as a function of matenal sensltiljlty m the lumt ioad method Here f *~~*
maximu]n load carrying capacity of pipe with certain variation of material property while P .
that ~vith no variation of material property.

20

e Sy
l Su

~R 10

~
8
-

ee

~le,
\

.

~ -10

~
~
-20

50 40 -OO -20 -10 O 10 20 eo 40 50
Material sersitivlty pq

FIGIJRl: 4. SENSITIVITY ANAl.YSIS RESUl.1'S Ol: 1.lMIT l.OAD Ml:TllOD

12 - 6' is the~'a:xo isAs shobvn in tl]e figurc, the sensitivity due to the variation of S,, valtte is higher than tl]at duc.

to thc variation of S), value. The weight factor (WF) for resl)ective materi'al propcrty is dcrino(lfrom Figure 4 as following.
,'~ Sl.,
i= l

wherc n is_~ the total number of matcrial prol~!rty and SI_f mcans the slope of i th materiill
sensitivity vs load.

_'\ccorctingly the ~veight factors for respective material property are defined as given in ~Pa}]lc l.

TABLE l. WEIGIIT FACTOR FOR LIMIT LOAD METIIOD~,'1ateriatl , Sy S**

Propel~t\' =Vfeig_ ht , 0.7
Faictol

0.3

JZT_ ~n_d C~I_'~_D_ metl_1_od_. The J-integral for a long aJxl'? surface crack lit I t)llxt Is c\LJrtss( dag: (7ahoor, 1989)/=7rtf'S;:,1'CtaS, ", • . ~ /E) i hl(S,. 'S.,)"+1 ( lO,vL~ Fi( ' rL*~/' ~pk-"u/ : f:'u ~~ ~ ,~;~

f= .c_'eometn' .tnd'!,jrhl * = shape functiorl

and the J-R curve is expressed as following.

JR= Cl(da)c= ( I l)
Goveming material properties in the J/T and CDFD analyses are;- Ramberg-Osgood constants ; ct, n- Yielcl strength ; S~.

- \_ ,Iaterial constants of J-R curve ; Cl, C

Figure 5 shows the variation of miLximum l0ad as a function of material sensitil~'jty in thc Jfl'
method. t\s sho~vn in the rigure, the S,. value is_ the clominant factor and the ~veight rnctors forrespcctivc material property ',~e clefincd as given in Table 2.

12-7

OO

~

20

~R_

8
-)( 10

ea
-,n
C1

- Ca
- SY
v

6e

~O
\~

o

I

e

9.

l

,

~ _10

*E5

~

~ -20
~O

-SO 40 ~O -20 -10 O 10 20 30 40 50
Matertal senstivity pq

FIGURE 5. SENSITIVITY ANALYSIS RESULTS OF J/T METIIODTABLE 2. WEIGII~" FACTOR FOR J/T METHOD

D~~FAI)_ _nOelh~. In general, thc e(luution f()r DPFAD curve i.s expressed as

J/f(a.P) = l/K~ (12)_ f(acff' P) +J~~~~L2~1~ f(a. P) f(a. P)

Therefore, thc gol~~erning material properties in the DPFAD analysis are sa]ne as those of J/Tmethod. Figure (; sho\L's tl]c ~/ariation of maximum load as a function of material sens_ itivity in

the DPFAD method.
As sho~vn in the figure, the S,. value is the clominant factor as ~vell. The \vcight factors forrespective materi~rl l)rol)crt~' is definecl as given in Table 3.

I?IL~)_f_Inaterla L Qrop~erty.

Tl]c PR of matcriill l)roperty is compute(i based on the previously obtainecl bveigllt factor 'andcertainty filctor of rulc. Fis;rurc 7 shobvs an c\a]nple for computing the PR.
Assumo ti]ilt a, n. S~. Cl, C2 Ivaluos ilre rc(luircd for thc an'alysis, and only a, n, S_'. values_' aregiven. If tl]c inrercncc r()r ol)taining unkno\vn matcrial propcl~ty is

12-8

3o

e
20
~R

~ 10

l C1
A C2
v Sy

v

7

)(

~

~

o

aF

-10

~~

a

8

o

el

t

~
a'_

-20

v

~ 50 40 ~o •20 -10 o lo 20 30 40 50
Materlal sensltMty Pq

6 SI NSITIVITY ANALYSIS RESULTS OF DPFAD METIIOD
l•'1GURF
TABI.E 3 WEIGIIT FACTOR FOR DPI,'AD MI~TIIOD'~". " " ' ""~'I'~l~.?.-'*I~:'1r':: R•(.JJr.d r,W'rtld p,OPtlt•t:: ab n, 8y. C1. C,[' 'A'aJn' 9iv'o. n-. Fop~d"!:: abn~Sl
thoreeY•• I Ftrx' ,•d,r~,, No

PoN..1"eo r•to

~ 1•.O ~,FIGIJRE 7. FLOW ClIART FOR COMPIJTING PR
f;lilC(I(lLIC toin s_' u rf ic icn t

input infonnation, the current PR is coinputcd asl)R = a (91~6)~n ( 14'~,)+Sy (56'%,) =79'~6

12 - 9following.

If tl]e infcrQncc is successful, the bveigllt factor is redefinecl 'as follo~ving:

WFR=CFRl,1 E \ Wl h ( 13)~~'here,

WFR = redefined weight factor
Wl~s. = ~veigi]t factor clefined by sensitivity analysis
CFRlLE = certaint~.' factor of the rule ~vhich is used toinfe]' unknobvn material prope]~ties

In tl]e next phase, the PR is recomputed as follovving:

Step I : Redefinition of ~veight factor

Cl = 07 \ 70~~;) = 4.9%~' tep 2 : Calculation of PI~
PR rl (9'10)+n (1496)+~) (,~)696)+CI (4_9016)4C2 (9.X('/u)* 93.7'~~;)_\fter obtaining the possession. rat.io f)f material l:rol)erty fo' I'c's.~1)ective_ integrit}' rv,r,]uatlol'

metnods. thc meth()d vvhich has the highcst l)ossession n*tio is rt'-co]nnlencled as ttl ' rlost
appropri.~te integrit, *~valuatio.] Inethod.

Evaluation method based on ASME Sec. XI criteria
In Appendix C and Appendix H of AS~,IE Sec. XI, the evaiuation procedures for austenitic andferritic piping materials are given. The austenitic piping material is distinguished between

austenitic materiais with high fracture toughness ancl certain flux welds that have lower
toughness. Thc base metal and non flux welcl (GTAW, G\_ If\W) evaluation is based on a plastic

collapse failurc mech'anism and the allo~vable flabv sizes 'are generated from the limit load
analysis. The Mu\ ~vcld evaluation is based on an ttnstable crack tearing failure Inecl]anism ancl
the ailo~vablc flaw sizes are detennined using EPF\_ I anaiysis methods.

The fcrritic l)iping material is clistinguished by i]igh toughncss materiais (base met'aD andsubmerged arc ~velcls (SAW). Since thc l)reciicted railure mechanism for thc Mux ~vckls was
unstable, Thc flabv e\tension may occur at ioads lo~ver th'an the plastic collapse analys_ cs. The

applicable railurc modc is defined depending on material toughness, Ioacl type and magnitudc, andfla\v size, s'hapc and orientation. Finally, an appropriate evaluation method is rccommcndcd.

CASI~ S'I'UDII~S
To illustrate LI]e usefulness of N. 'I]iES s_ ystc]n, cnse studics ~verc perfor]ncd.

I:jgure 8 s~hows tl]e im~Llysis m(xlel (]f a pjpe (mcnn r(Idius R = 21(;(] mm, tl]icknes_'s t = 216mm), (at) bviLl] ,tn ilxi:tl scn]i-cllit)tic;tl crnck (dQt)th ;t = ~4 tnn], ct'itck longtl] 2c = l(;2 n]m), (I])

12-lO

vvitl] a

long axial crack (deptl] a

The material properties of the= ~~54 mm). The pipe is under intem'alpres s urc

of 17\_ II)a.

pipe arG rollowing.

I

I

,k-~'
'

t

t
,

1-,

,
t

j 2C

semi-elliptical crack

,

,

,

,

a:
II ~I):l. . .F. . ,..,.'-FIGURE 8. GEOME'rRY oF AXIAL SEMl-ELLlpTICAL CRACK

Sv ' 594 \- iPa
St 1: 703_ \_ IPa

E = ?03 GPa_? of 3~:] \,s.A( e, c-',= )
il oo
\

(

( 14)

JR:= 132(.J'!;)L " "' I (kNfnl/

Assuming that full material

properties are

not given,( l.))

the following

three case studies

v,' ereperf onned .

Case I:Sy value.

Inferring 0-8 curve

f ro mCase

II :Inferring a-8 curve

f rom Sy and Su va]ues.

Case

III:

Inferring J-R curve

f roin8c Value.

ve o

valu

Case :I'n 0-8This case study is when the yield strength (S,.) is given instcad of rull 0-8 curve. Figure 9sl]o~vs the screen of database bofore inference and Figure 10 shows tl]e screen after inference.;
Su = 891 \_ IPa, a = 1.64, n = 6.09 and E = 197 Gl)a.

EPF\. I analyses were perfomled based on 4 differcnt integrity evaluation methods (CDFD, J/rl',
DPFAD, Iimit lo'ad) and the results are sun]marize(.1 in Table 4. Ilere the safety factor is defincd ns'

ratio of operating I]ressurc P against ma\imt]n] Ioad P~,~*. .•\s sholv
clirfcrence bet~veen the Case I ~md refcrence case (when material prol)crtics ?u~e rL]lly givcn) is ~vitl]in2:~) '~~0.

12-ll

=L

Np,ES SU8 UOOULEMll•fi•,:D•,•b••• -,I__L:L~!

,

TtNS"JE TEs' DJLTA-- -rooG"NEs8 TEST OAIA

se fr~~,,.~ ~X5rFE~T

s' f5~~i~~bt IAUsTtN,'t,i. Fr~~t Nll'

Jb r~~ b,uN

LN/1' '

,dd ~X~E~t f5T~ 'o'll"
s' r5T~wN' Ii.,td f--I"-'a"f:ET~- f5'~~ :lu- r5~~,, fr~ !

,1ATtRuL P,,Of~nlY ,NfiRCNC: --

f5i~~
ol
l"
c2tN
~~~

--i :,

ru,,,,fNT P9ssllu: Ev,Lt,AnoN b,ET,,oos -- - -Fr iS•f~r Lt~~, ulslic mJu:]oac ItDwu5

r STA8T WftRol':r EIJN$TIC Pubsnc mACIUf,C NtC,w,tcs
r u,1,T LoN, /tNJd!Ys,s_ ___i~] ~~lI
n~FERENCE FOR
FIGURE9. DATABASESCREEN BEFORE
CASE___=uil

iL_

NPiES SUO MOOOLE:
U•tl~ritl
D_ Itlbl•C

TE"SLE TESI OIATA - - -- 70'JGr'~ss TEST OATJ'r - tl- isA]12_ 'P'04 sl iobN~-'
8'1' la'El~"d'w" fA'JS T t NIs' T t ,~G 5s' t"I'~ J'c rb~~ o""lil,',,.,d rEX;E~~ rl557;52 tL'u-'
i o ' Pl"Q
FrTr~
L

ttN'a'lf;i5~~ rrl~l, ; ,
gdb" f5rFE;~
c2 F~T~s' ' L,,1'1'
:o a

'l'

:i
dd~ ,]rl~
sJ"!

. F~r
"ATEn'AtPPoptR'Y INFEncNc!--- - a"'uNT PossiILJE5'~JJ~Tlo" t'u"aDs - -fr sTAl'l tNflEnt"ct , r: :::-

; F b'rtil '- ' " ': :1't!b 1 I11"~I:uLsl"Fr'Ac7u9t t'o'AN'$r EULSTC PutsT'c Ffw:T'x'E WEow'lcs

F L'wlr LOAD ~w:vs's .

[~I ~~]l
FIGURElO. DATABASE SCREEN
AFTER INFERENCEFORCASEl
TABLE 4. ANALYSISRESULTS OF CASE

Crtick~' hatpe

A\'itti~;.:' Cmi - El lipt icu l

DPFADl_[ITliL

CDFD
JfTl.()ll~

2.96

.-~-

l .7*ILLl f-t~'l .O I

A\i{ilI)1'FAD : I .(f)X i. -- --. --- ---- -tl.imlL l_o:id' :~.X2
s. ' I" ':~.'arctv ract,,r ~L'hen iuferrinX a-F.

~. ' I"":~.'are[\ f:act,,r Lvhcn irll material[)i r f .

2.J6 20! 2:)•1. 1 2

LOtiCl I--~rl .69 * :~l .b*2 1 X
l .(~S i i {)')) .tvll6 : 2. lcurve lr,]m S~ value.
t)n'llertics are full,' gj~'en.

••• SI:'- SI:"_ lOO
Sl"
12- 12

~;ase_J~ :_1_~f~_rr_i!]gJL0-8 curve from S an 'lL~al~~~~lu s'I'his casc study js when the yield strength (S~•) andgiventl]e ulti]nate
instead of full a-8 curve. After the inference, unknown material propcrties are ohtaincd

as; a=10.99. n=1.(;8 and E=197 GPa.

EPFll ana]yses vvere perfonned and the results are summarized in 'I'able 5. As sl]own
table. tl]o maxi]nttm difference between tl]e Case 11 and tl]e reference case is within 150~,.

TABLE 5. ANALYSIS RESULTS OF CASE ll: Crack Sh.alx? I \. fcthod SF ' : Diff. j
SF' .

: CDFD 2.86 1 2 ')oc: , 12 {' A:\i'.rl
Semi-Elliptical :jLimit Load: 3.59 i , ' 3.29 : 9 l!

.- -i I--J~: .CDFD. 77 lL- -- - ~ -- - - - l .6{] )1.85 l: ;, 'T I : .62
14

Long ~ixiLrl j DPFAD ' 1.68 : 1.66 ; ,'1--* .

!,__iLim]ri 3.06 .
__ _ ________Ii~)____ i_ _:_. __L_ __ _ ____J._, l_rJad;. . .__, ._ 3.~:~3lb~.~ F" : Safet~ * fi:L:'r whc:"tlfT=~:tl:. o-s ':u,1re i.,.'r~ L~T anf:""* vahtL :(!~F" :Safery fa(:tor lvhen all mar.erial properties are rUll~' given.

rhis case ' ~tL (a) sitov~:~ that the-,sld'litional _knov.'Ledg( of S, iml"r_ol~es tile._. :1~*1,:'is acc'ur (') in

the DPFAD mothocl and the limit load method.
Case 111 : I__nfe!Iin-R c v f m 8 val

This case study is when the J-R curve of the pipe is not given. From tensile property, theJ-R curve is obtained as; Cl = 321 (kiNlm), C2 ' 0.38.

EPF\_ I analyses ~l'ere perfonned and the results are summarized in Table 6. Since the limit
l0ad analysis is not affectod by tl]e J-R curve, it is not considered here. As sho~vn in tl]e table,

the maximum clifference between the Case 111 and the reference case is within 33%

TABLE 6. ANALYSIS RESULTS OF CASE 111• l,

Diff .

Crack Shape ~Iethod ' SF' i SF"

(oj,)

1

c:: 16 :CDFD 2 97 t .'_..)o,

AxiLalJfT' S.~ emi-Ellipticarl ~DI'FAD-i-,- ---13 .OO : _.-16:') L)1_)

CDFD 2.0S : I . (39 2~, I . 62l.oll~ A\ii_ll J I l,-17 ,-~-.-: Dl'FAD 2.2 1 : I .(~5{)~r. ~~~~ ,o'S. :.:).

S.~~1'~'~;~S~;al.'afet)Tf~'act,)r ~N~hen inferri]]~ i-R curve fr'Im E. value

Sl:": S.'aicty lact,,r lvhet] all material l]n)1)ertics

12-13

are rull,' ~ivcn.

in theEPF\_ ,1 analyses cannot t.)e l)rocosscd without material's j-R curve in a majority of soft~vare.

However, this case stucly shows tl]at NPiES system still provides analysis results vvith reasonableaccuracy even tl]ough J-R curve is not given.

CONCLUSION~PiES system has been developed to evaluate the nuclear piping integrity. The syste]n isdesigned such that the most appropriate integrity evaluation method for given infonnation is
selected based on the possession ratio of material property and the unlmown material property is

obtained through inferring the known material property. Several case studies were perfonned andthe usefulness of iNPiES system was demonstrated.

REFERENCE
AS\_ IE, 199 2, "ASME Boiler ancl pressure vessel code. Section Xl

Bloom, J.\_ I. and \_ •Ialik, S.N. ., 1982, "proceclurp for the assessmenr)ressure vessels and pjpjng containing defects", EPRl NI)-2431.

htxal, V.K., \. ~orrs, D.~/' I. and Scrver, \V I 1988, "I.eLtk-F;efore-Break : \l] mt !;rat(.dIL)p (, t ,1 ror 1•lgh ene gy pllun*. , Int. J. PrQs. Ves. &• Plpm~i \ ol 34 p

Kiln, Y .,., 19~,.}. "T,raclure
Choy, Y.S., Sec.[. C.S., yang, K.~/Z.= Ra. I.L, ?Lnd Ahri, H.Y..

3L.4 nucle'.',,' . cn

Trans. of the 12th S\. IiRT Conference. Vol.(;.. pT).435 - 440.

him, Y.J., Seok. C.S ., Cl]oy. Y.S and Chang ~ S 1995: "Prcdrctlon of fr lct l,~e *esistimcc'In e from tpnsll,(: data" AS\lE P\ I C*]lier, nce lvol 304 pp. 277-•284.
Kussmaul K., Jovanovic, .'\. ~m(t ~,IiLik~, K., 1991. "Expert system based life ass,*~s ne'tt of hlghtempcrature co]nponents", Trans. of the I Ith S\. IiRT Conference, Vol.SD1, pp.73-84.

Lucia, A.C. and Volta G 1991 "A knowledge based system for structural reliability
assessment", Trans. of the llth S~'IiRT Conference, Vol.SDl, pp.gl-96.

Okamodo, A., Kataol a S and Watanabe \1 1987 "The development of expert system fordefect identification and its assessment", Trans. of the 9th S~liRT Conference. Vol.B,
pp.247-252.

12-14

13

Recent Progress in Korean Nuclear PLIM Progeam
T.E. Jin*, H.J. Choi (KOPEC), I.S. Jeong, S.Y. Hong

(KEPCO)

Recent Progress in Korean Nuclear PLIM Program
Tae Eun Jin, Hyung Jip Choi, Korea PotL'er Engineering Com~ny
Ill-Seok Jeong, Sung- Yull Hoag, Korea E!(,ctric Power Corporation

Summary
The recent developments in Korean nuclear power plant lifetime management program which
has been performed for the last three years is introduced, together with the descriptions ofmaintenance activities. This leading Korean PLIM project has been focused on such areas asplant data survey, transient history, screening and prioritization of sv. stems, structures andcomponents (SSO, aging evaluation of major components and other supporting activities. The
current status is outlined in tems of major tasks including aging evaluation of thirteen major

components. The future long-term plan which eventually airns at maximizing the economic
benefit for both the utility and ,its customers is presented. Also described is, from the viewpointof plant life extension, the technical development

1 INTRODUCTION
Nuclear power plant (NPP) technologies in Korea have been remarkably evolved since thecorrunercial operation of Kori Unit I in 1978. As shown in Table l, ten nuclear power
plants are under commercial operation, while six nuclear plants are under construction andanother seven nuclear plants are scheduled to be constructed.

In the early era of NPP industry, a vast amount of efforts had been devoted in Korea
to accurnulate the technology and experiences related to the plant construction. As time

passes, plant aging and majntenance problems became a matter of concem in the NPP
indu*_~y. Under these circumstances, Plant Lifetime Management(PLIM) of nuclear power
plant attracts attention and is considered as an effective way to address this issue. This isbecause the PLIM has a strong possibility not only to solve the plant aging andmaintenance problems but also to provide the vision of extended plant operation beyond thedesign lifetirne.

As parc of the long term nuclear technology development program by Korea ElectricPower Corporation(KEPCO), the plant lifetime management project named "Nuclear Plant
Lifetime Improvement and Management(D" was started in November 1993 to cope with theaging and obsolescence of Kori Unit l. At this stage of PLIM phase I, the project aims at

a feasibility study of lifetime management of Kori Unit I together with the aging
evaluation of the thirteen maior components. The results of. the PLIM(1) are expected toinfluence the decision making of long-term Kori Unit I lifetime improvement. Subsequentiy ,the PLIM Phase n (Detail Lifetime Evaluation and Engineering) will be perforrned, inwhich the workscope and schedule are subjected to the outcomes of the feasibility studv. .

13-l

Table l. Status of Nuclear Power Plants in Korea

This paper thus introduces KEPCO 's basic PLIM strategy, some recent developments inthe on-going PLIM(D and other works related to the lifetime management of Kori Unit l.
To date, field data survey, systern/structures screening, compohents prioritization, fracture

mechanics test of Kori Unit I reactor pressure vessel surveillance coupons and component

aging evaluation of major components have been done. Other activities such as the
economic analysis, regulatory considerations and key technology reviews remain to bocompleted by the end of 19~.

2 KEPCO PLIM PROGRAM
The primary goal of KEPCO PLIM: is to operate a plant safely and economically up to
the plant specific design life. If the first goal is possible, then the operation of the nuclear

power plant beyond the design life to the optiunum life will be pursued as the second goal.
Furthermore in parallel with the PL~l proj ec~ key technologies rerluired for supporting thellfetime management are being developed. A specific feasibility study evaluates each plant's

optimum lifetime which then bccomes the target hie of the PLIM: efforts. The second goalof the PLIM program is to operate a plant up to the optimum llfetime. If such an optimum

13 - 2

lifetime is longer than the design life, additional activities to operate the plant beyond thedesign life shall be incorporated in the long term preventive or predictive maintenanceprogTam.
The master plan for PLIM including the lifetime extension of Kori LTnit I and other NPPs
in Korea is composed of three phases with Kori Unit I being regarded as a leading model plantof technology development. Such categorization generically stems from the level of details andrefinement that are to be accomplished during each phase of the project In Phase I, feasibilityof extending the lifetime of Kori Unit I in terms of technical, regulatorv. , economic aspects isestablished including aging evaluations of the prioritized thirteen major components. Phase H
program will perform the detailed lifetime evaiuation of the major components and other criticalcomponents screened in Phase I. The PLIM implementation plan for Phase m will berecommended based on the results obtained in the preceding Phases. For reference, overall

descriptions for the phased PLIM programs are outlined in Table 2. This long term plansubjected to changes in accordance with the outcomes of the PLIM(D feasibility studv. .

Table 2.

Three Phases of the PLIM Program

3 KORI UNlT 1 PLIM PROGl~M
3.1 PHASE I
The phase I workscope consists of the following ten tasks and is described in the sequence.

PLIM project plan and design llfe review
o Task lo Task 2 Screening major SSC's
o Task 3 Data survey and review
o Task 4 : Evaluation of reactor pressure vesselo Task 5 : Evaluation of major SSC's
o Task 6 Monitoring systems for PLIMo Task 7 Survey and review of PLIM regulation

13-3

o Task 8 : Economic evaluationo Task 9 : PLIM technology developmento Task lO: Feasibility study reports
3.1.1 Design life review

The major workscope of this task involves overall project planning and establishment ofdesign life for the Kori Unit l. Especially, the basis for 30-year design life of the Kori Unitl is reexamined, together with the provision of feasible amendments of its design life from 30

to 40 years. According to the archival documents survey, the design life of major components
in Kori Unit I including the reactor pressure vessel is confirmed to be 40 ~.,ears.
3.1.2 Screening and Prioritization

Critical component identifications for aging evaluation are an important part of the

PLIM Phase I efforts , because to identify which components are crucial to the plant
lifetime is necessary to ensure the proper focus of Phase I efforts at the beginning of the

PLIM prograrn. These critical components were identified though the application ofWestinghouse Owner's Group (WOG) screening and prioritization criteria to the Kori Unitl system, stmctures and components.
The screening process applies safety-related criteria which are based upon the US

NRC ' s license renewal rule (LR) and maintenance rule (MR) which are lOCFR54 and
10CFR50.65 respectively. Additionally, the screening process applies power production(PP)related criteria which are based on plant availability.

Aiter screening the Kori Unit I systems and structures, critical components andstructures were identified and prioritizes to detennine their relative importance.
Prioritization of Kori Unit I critical components applies ten attributes which were selected-

to assess the impact that either the replacement or refurbishment of these criticalcomponents would have on the decision to improve design life. For example, these
attributes are cost to replace or refurbish, impact on plant availability and radiation dose,etc .

Prioritization result shows a similar result to the previous experiences.

3.1.3 Data Survey
As a prerequisite to the evaluation of the plant aging status, a huge arnount of design

and field data of Kori Unit I accumulated since commercial operation should be surveyed
and reviewed. Even though tremendous man-power to re-produce useful data from the raw
materials is required, data survey is the most important j ob that has to be done in theprocess of compiling operating transient numbers. Such data required for the PLIM (D can
be classined as follows.

13 -4

o General methodology and technical references
o Operating transient historyo Component design specification and manufacturing datao Maintenance and inservice inspection data

3.1.4 Major Components Life evaluation

In order to evaluate the aging status of Kori Unit l, the following major components\vere selected at the beginning of the PLnvl: (D. All major components were ranked within
top twenties of the component prioritization results.

o Reactor pressure vesselo Reactor vessel internals

o Control rod drive mechanisms
o Reactor coolant system pipingo Reactor coolant system charging and safety injection nozzles
o Pressurizero Pressurizer surge and spray lines nozzles

o Reactor coolant pumpo Reactor pressure vessel supportso Turbineo Generatoro Containmento CablesSubsequently, the stressors and the degradation sites ' and mechanisms in conjunction

with the resulting failure modes and their operating history are identified through theappropriate tests and technical evaluations. In consequence, the task ends up withqu'.intitative evaluations of the plausible age-related degradation mechanisms, with the aidof proven technical papers collected by literature survey and the generic technicalprocedures published by the off-shore organizations who performed life management study
previously. Table 4 shows a life evaluation methodology with the reactor pressure vessel

as an example.
Special attention is paid to the reactor pressure vessel for its significant importance inthe nuclear PLIM prograrn. Fracture toughness test of the WOL specimens from the Kori
Unit I surveillance capsule irradiated for 34 EFPY showed that the fracture toughness ofKori Unit I beltline weld material is almost identical to that of other Linde 80 flux weld

metals and better than the criteria of the frac,ture resistance curve proposed by the US.

NRC draft regulatory guide DG-l023. For the comprehensive integritv. assessment of thereactor pressure vessel, the pressurized therrnal shock criteria of lOCFR50.61 is being
applied to the Kori Unit l.
In some cases where the desigh data are difficult to obtain, full engineering calculations

13-5

Table 4 - Summary of Evaluation Procedure for Reactor Pressure Vessel

are performed to establish the necessary data set The PLIM (D, for instance, conductedfatigue lifetime evaluation of the Kori Unit I pressurizer surge line nozzle with thecornmercial finite element package, NISA, in order to provide an evaluation procedure ofstress analysis and residual fatigue hfe. The result of this study demonstrated a very good

agreement with that of the vendor design stress report which did not show the calculatingprocedure like a black box. In the end, the residual fatigue life of the nozzle operated for
15 years was sufficient to meet the first goal of Kori Unit I pLIM, that is 40 years.

3.1.5 PLIM Regulation Review

Regulatory rules are intended to guide the nuclear PLIM in a proper wav. '. The
preliminary survey about overseas license renewal trend and rule development is necessary

to direct a domestic policy of rule making in the near future. The govemment body, theMinistry of Science and Technology, and its agenc.y, Korea Institute of Nuclear Safety, arecunlentiy in charge of nuclear power plant licensing and other license-related issues in
Korea KEPCO provides such regulatory bodies with information and interim results of the
PLIM study to help pro-actively the rule making.
3.1.6 Monitoring Systems for the PLIM:
Prior to and in parallel with the further detailed evaluation of major components, thedevelopment and utilization of monitoring s~ ystems deserve due considerations. In this task, the
13 - 6

currently available monitoring systems for the PLIM: are thus be studied, together with theidentification of degradation sites to be monitored.

3.2 RESEARCH ACTIVITIESIn addition to the feasibility study, five R&D items are under \va~.' in the areas ofradiation embrittiement corrosion and cracking , water chemis trv. , non-destructl ve
examination, and aging of cable and I&C. Namelv. ,o Utilization of small or reconstituted specimens of reactor pressure vessel materials

o Pb stress corrosion cracking of stearn generator tubes
o Evaluation of hideout return and stearn generator crevice conditionso Natural cracked small pipe specimen and defect signal analysis
o Destructive test of thermal and radiation exposed cable

3.3 COMPONENT REPLACEMENT
Kori Unit I steam generators, Westinghouse model WH-51, one of which has 3,388
Inconell 600MA tubes, have gone through a lot of maintenance works such as plugging,sleeving and chemical cleaning due to tube pitting/denting and primary water stresscorrosion cracking.

The feasibility study for the steam generator replacement at Kori Unit I has alreadybeen completed on a separate basis from the PLIM project currentl~., under consideration.
Taking the plausible plant lifetime extension into account, both the deterministic andprobabilistic economic evaluations of the related issues indicated that near-term replacementof steam generators with some plant rerating provides the most economically favorable
option. The other optional scenario aimed at avoiding or postponing steam generator
replacement by sleeving/plugging maintenance strategy turned out to be less cost-effective

strategy. The necessary ensuing actions will then be taken regarding the steam generator

replacement. KEPCO plans to replace it with Inconell 690TT tubes and stainless steel
broached support plate by 1998.

According to the field inspection of Kori Unit I Iow pressure turbine rotor disc, many
cracks were found at disc, dowel hole, and disc head due to moisture induced intergranular

stress corrosion cracking. To verify the safe operation of the turbine with the fracture

mechanics analysis, KEPCO provided a mitigation to re-inspect and repair the cracks
during the next outage and then, however, determined to replace the rotor and dial)hragmof the low pressure turbines with welding or mono-block type rotor.

4 SUMMARY
Nuclear PLIM is at present one of the most important tasks in Korean nuclear industrv.

as Kori Unit l, the first commercial nuclear unit is being aged. This papor introduces'

13 -7

KEPCO's basic PLIM strategy, Iong-term plan, current interirn results of Kori Unit l
PLII~I feasibility study and other related programs.

The feasibility study includes field data survey, screening and prioritization of the
SSC's, component aging evaluation, and stress and fatigue analv. sis of pressurizer surge

line nozzle. Fracture mechanics test of Kori Unit I reactor pressure vessel surveillancecoupon was completed and the pressurized thermal shock studv. is recornrnended. In theremaining period of PLIM (D project, further evaluation of major component aging, review
of regulatory issues, and economical evaluation of Kori Unit I PLIM program, are expected
to be done for the PLIM from the viewpoint of economic aspects.
Other projects that have to be considered for PLIM, such as process computer and I&C
upgrade, plant uprating, probabilistic safety analv. sis and reliabilitv. centered maintenance aregoing to be undertaken separatedly It is anticipated that PLIM will provide a good way oflong term life management of nuclear power plants in the cost-effective manner in Korea.

Referenc es

l. Korea Electric Power Corporation. 1994 Wbite Paper of Korean Nuclear Power Plants,
l 995

2, US. Nuclear Regulatory Commission. lOCFR Part 54, Recluirements for Renewal ofOperating Licenses for Nuclear Power Plants

3. US. Nuclear Regulatory Commission. 10CFR Part 50.65, Requirements for Monitoring
the Effectiveness of Maintenance at Nuclear Power Plants
4. U.S. Nuclear Regulatory Commission. raft Regulatory Guide DG- 1023. Evaluation of

Reactor Pressure Vessels with Charpy Upper-Shelf Energy Less than 50 ft-lb,
September 1993.
5. U.S. Nuclear Regulatory Comnaission. lOCFR Part 50.61, Fracture Toughness
Requirements for Protection Against Pressurized Thermal Shock

13-8

CONTENTS

1. Historical Evolution of Korean Regulatory System " '1

2. Nuclear Regulatory Structure in Korea " ' " ' '5

3. The Activities of the Korean Reguatory Organizatrons

8

4. Licensing procedure for NPPs " ' " " '11

5. Regulatory Inspection " ' " '13

6. Future Perspectives and Concluding Remarks
22

4t-2

1. Historical Evolution of Korean Regulatory System

Korea has one of the most dynamic nuclear power programs in the world.
During the last couple of decades, Korea has caJTied out a very ambitious nuclear
power program as part of the national energy policy aimed at reducing the extemalvulnerability and insuring against the global fossil fuel shortage. The first Nuclear

Power Plant (NPP) began its commercial operation in 1 978. Now 1 1 units are in

operation providing almost 50% of electricity in Korea. Seven units are underconstruction at the moment and 4 more units will be built by 2006. The current
status of NPPs in Korea is summarized in Tablel I .

Back in the bcginning period of NPP operations from 1971 to 1978, the frrst

commercial nuclear power program was implemented on a tumkey base. Kori Unitsl and 2 were ordered from Westinghouse, U.S.A. in 1969 and 1974, respectively
and Wolsong Unit I , the first CANDU plant, was ordered from the AECL, Canada
in 1 973. Contractors assumed overall responsibility for the construction schedule,
inspection, startup and performance of the plants. During this period, domestic laws

and regulations applicable to the licensing of NPP were not yet fully developed.
Therefore the vendor countries' Iaws and regulations such as 10 CFR, Reg. Guide
and Standard Review Plan (SRP) of USNRC were applied to the licensing review ofWestinghouse PWRs. As for the CANDU plant, Canadian laws and regulatory
requirements were applied as mandatory requirement. Construction perrnit (CP) andoperating license (OL) were applied simultaneously in March 1 976, as can be seenin the Canadian combined licensing approach and CP/OL was issued in 1 978.

From the early 1980s, six NPPs (Kori Unit 3&4, Yonggwang Units l&2,
Ulchin I &2) were constructed by employing a component approach with foreign
contractors. Contracts were awarded separately for major components of plants, thus
enabling more domestic industries to participate as subcontractors in the projects.

While on the regulatory and licensing side, the Nuclear Safety Center (NSC) wasestablished in December 1981 as a regulatory expert organization, which was thepredecessor of today's Korea Institute of Nuclear Safety CKINS). Two step
licensing system, construction perrnit (CP) and operating license (OL), was forrnallyincorporated into the law. However, the majority of important codes and standards

applicable in the vendor countries (U.S. and France) were still applied to thelicensing of these six plants only with somc appropriate modifications. As for

CANDU plant, thc Fina] Safcty Analysis Rcport (PSAR) was submittcd by Korca
Elcctric Powcr Co.(KEPCO) in 1982 accordin_~; to thc ncw]y amcndcd law and thc-1-

FSAR was rcvicwed once again at the NSC in ordcr to confirm thc dcsign safcty.

Starting from Yonggwang Units 3 & 4 contracts in 1987, Korea Power Co.
(KEPCO) had assumed the overall management and responsibility for construction
projects. The overriding priority for selecting suppliers was the condition oftransferring higher nuclear technology to Korea. The prime contractors weredomestic companies instead of foreign companies and several foreign companies wereselected as subcontractors. The same approach applied to the contracts for Ulchin
Units 3 & 4, and it will apply to the units that follow in the future.

lh the licensing of Yonggwang Units 3 & 4 for construction, some regulatory

difficulties emerged due to the scaled-down design from the reference plants of
Combustion Engineering Co. (CE) designed Palo Verde's System-80. With intensive
audit calculations and the third party verification in the area of safety analysis
together with the technical support of the USNRC and the IAEA, the CP was issued
to Yonggwang Units 3 & 4. Ulchjn Units 3 & 4, which are the first standard units

with substantially the same design features as Yonggwang Units 3 & 4 , werelicensed for construction with some conditions for implementing safety enhancement
regarding mid-loop operation, safety depressurization system and hydrogen ignitors aspart of severe accident mitigation, and ALARA. Recently, the licensing review for

CP of Yonggwang Units 5 & 6 has been performjng and several items has been
investigated to enhance the safety such as, Ievel control in CVCS, digitalization andduplication of the process control system, human factors in the remote shutdown
system, PSA for low power and shutdown, the filtered ve;nt system, etc.

Safety review for the construction permit(CP) of Wolsong units 2, and 3&4 hadbeen conducted since 1 99 1 , and the CPs were issued in August 1 992 for Wolsong

2, and in February 1994 for Wolsong 3&4. Operating licensing review for Wolsong
2 has been perforrning since May 1 995 and the OL is scheduled in August 1 996.
Regulatory efforts during the CP stage have been focused on the following aspects;

First, intensive review on the design changes/improvements compared with Wolsong
l . Second, applicability of the PWR safety issues to the units within a practicalmanner for the safety enhancement equal to or above international level. Third,
design suitability and environmental effect arising from the construction and operationof multiple, 4, units at Wolsong site. Several domestic and foreign expcrts wereinvited to participate in the special revicw arcas where high technology andaccumulatcd cxperience were requircd.

-2-

A significant progress has bccn madc so far in thc dcvc]opmcnt of Korcan
regulatory critcria, guides and procedures which arc applicablc to thc standardizcd
nuclear power plants, too. Industrial codes and standar(ls are under development andwill bc completed in several years. Since the operating nuclear power plants are
subject to deterioration due to ageing, appropriate aging management program suchas periodic safety review (PSR) with the emphasis on probabilistic safety assessment(PSA) has been actively proposed and studied at the KINS. This will lead to theestablishment of regulatory requirements for license renewal. In this regard too,

Korea will keep collaborating, with the advanced countries, particularly with thevendor countries such as U.S.A, France and Canada.

During the safety review of foreign designed nuclear power plants , thelicensing difficulties have been encountered mainly from diversified reactor types andvendor countries . The lessons have been leamed from taking appropriate measures
for the trouble-shooting of these problems such as difference Codes and Standards
and differences in Regulatory Procedures and Practices

In this regard. Korea's long-term nuclear research and development program

calls for a progressive design improvement for existing PWRS and PHWRS as well
as for those to be built up to 2006. After that, next generation reactors will bc
developed and introduced in Korea.
Therefore, pertinent regulatory positions will be established step by step, keeping

pace with the aforementioned nuclear R&D program, in terms of safety regulatory
and licensing system, procedures, design requirements, codes and standards, etc.

-3-Tab]c I .

*

**
***

+

N uc le~r

Powcr Plants

: Il
Units
In Operation
Under Construction : 7 Units

Planned : 2 Units
Rcactor typc wi]1 bc

determined by Nov. 1 996.

-4-

In

K orQ~

2. Nuclear Regulatory Structure in Korea
In order to cany out the nation's nuclear power projects successfully, nuclear

safety has to be assured through all stages of project devclopment, including site
selection, design, manufacturing, construction, oporation and decommissioning.

The basic concept of nuclear safety in Korea, as in other countries, is not only
to protect the plant crew and neighborirLg inhabitants from radiation hazards but also

to minimize the subsequent from radiation and keep the radiation effects as low asreasonably achievable. This concept is basically underlined in the Atomic Energy
Act of Korea, which provides the legal foundation for nuclear activities . Regulationand licensing of nuclear facilities in Korea are based on the provisions of theAtomic Energy Act, Enforcement Decree and Enforcement Regulation, Notice of the
Minister of the MO ST and Technical Specifications which are parts of the safety
analysis reports (refer to Table 2).

The ultbnate responsibility for the safety of a nuclear power plant (NPP) rests

with the operating organization. The govenment, the Ministry of Science andTechnology (MOST), has in nature a general responsibility for ensuring theprotection of public health and safety by the regulatory control and safety inspection

on a government level. The Korea Institute of Nuclear Safety (KINS), entrustedwith the regulatory works by the government, perforrns a detailed assessment of thetechnical submissions and an inspection of nuclear facilities as a technical expert
group.

In conformity with the atomic energy laws, the licensee should submit to the

MOST various documents demonstrating the adequacy of the proposed design. It isthen the task of KINS, to review the licensing documents and to determine whether
the design complies with the specified safety requirements for siting, construction
and operation of the proposed nuclear installations. The result of the technical review

and assessment is reported to the govemment and then the MOST issues a permit orlicense to the utility based on the KINS' assessment report.

Nuclear safety inspection and enforcement are of vital importance to ensurethat all activities at nuclear installations arc in conforrnity with thc rcgulatory

requirements and licensec commitments. Thc MOST, the KINS and thcir regiona]offices at cach nuclcar powcr plant sitc havc thc rcsponsibi]ity to assign inspectors
to manufacturing facilitics, conslruction sitcs and opcrating nuc]car instal]alions in-5-

order to conduct nuclear safety inspoctions.

The fundamental regulatory structure in

Korea isschematicallydescribcd in

Figure l.

Safety Authority

Licence

MOST

Application Reporting
Licence
Entrustment
Issuance
Inspection Application Technical Expert Group

Nuclear Industry K I N S
Inspection

NPPs, Fuel Cycle Facilities, etc. Technical Evaluations

Fig.1 FundamentaJ Framework-6-Tab_ l_~ 2_i_

L_~g_I~_1~tiY~_ S_ys.._.te_m_~f Atomic Energy Lavys-7-

3. The Activities of the Korean Regulatory Organizations

The nuclear regulatory organizations in Korea are composed of three parts,
namely, a national level decision-making body represented by the Atonaic Energy

Commission (AEC), a regulatory authority with enforcement power represented by

the Ministry of Science and Technology (M:OST) of the Korean Govemment and atechnical expert group established to support the MOST with its technical expertisein the development of nuclear regulatory policy and also in the enforcement of
nuclear safety laws and regulations , which is represented by the Korca Institute of

Nuclear Safety (KlNS).
The nuclear regulatory system is describcd schematically as follows :' President 'I

Prime Minister

Atomic Energy Cornmissio~l
Ministry of Science & Technology
Atomic' Energy Office

I

I

r~~
S afety Assessment

Officer :Atomic Energy

1

Policy Officer

Korea Institute of~N~u~'="I~~~afety ;Fig.

2 Nuclear Regulatory Organization-8-

in KoreaThc activitics of AEC, thc MOST and thc KINS m thc arca of safcty of NPPs canbc summarized as follows:o Atomrc Energy Commission (AEC)

The Atomic Energy Commission (AEC) is chaired by the Deputy Prime
Minister. The principal function of the Commission is decision-making on major
nuclear policy issues such as R & D on a national level, industrial projects, andsafety regulations .

The Special Committee on Nuclear Safety (SCNS) was established under theCommission in November 1 989. The SCNS perfonns delibcrations on theoutstanding safety issues identified or developed in the licensing review process orvarious inspections by the KINS. Based on the delibcration by the SCNS, the AEC
rules on the safety issues' contained in the license application documents , then alicense is issued by the Ministry of Science and Technology (MOST).

o Ministry of Science of Technology (MOST)
The MO ST is the governmental organization responsible for establishing andimplementing nuclear regulatory policies for the regulation of nuclear activities

related to power and research reactors and radiation applications. It is also
responsible for making nuclear research and development policies for peaceful usesof nuclear energy.

The MOST has the Atomic Energy Office under its framework, headed by theAssistant Minister who is supported by the Atomic Energy Policy Officer and the
Nuclear Safety Officer. The MOST operates the Resident Inspectors Office at eachplant site for the daily routine inspection of the plant.

o Korea Instrtute of Nuc]ear Safety (KINS)

Thc Korea Institute of Nuclear Safety was established in February 1990
through special legislation by the National Assemb]y. ,The KINS is a technicalcxpert group established to support thc MOST with its technical expertise in thedcvelopment of nuclear rcgulatory policy and also in the cnforccment of nuclcar
safcty laws and rcgulations. Entrusted by the govcmment (Ministry of Sciencc andTcchno]ogy) in accordancc with thc Atomic Encrgy Act, KINS is rcsponsiblc for :-9-

. Safety review and evaluation to assure the safety of nuclear installations;
Safety inspection for nuclear installations;. Technical standards development;. Radioisotopes and Radiation Generators Regulation ; and,. Development of regulatory policy.

Figure 3 Korea Institut~ ofNuclear Safety Organization- 10 -

4. Licensing Procedure for NPPs
The licensing procedure for NPP in Korea consists of two steps, in general,
i.e., Construction Permit (CP) and Operating License (OL). The detailed licensing

procedure for NPP is summarized in Figure 4
The basic philosophy bchind licensing reviews in Korea is to lead the industry

towards the achievement of higher safety levels for the proposed nuclear power
plant, compare to the previous ones by progressively incorporating the design
improvements. To achieve this goal, special attention should be paid to thefollowing items during the licensing reviews :- Feedback of domestic and foreign operating experiences .

- Application of new regulatory requirements.

- Incorporation of the improved design features of the evolutionary LWR 's- Evaluation of safety features against the severe accident.

- Establishment of well-defined Quality Assurance Program for maj or equipment
and components .

- Independent assessment of computer codes and methodology used in safety
analysis .

o Construction Permit (CP)The review objectives of CP are to conflrm the safety of the proposed NPP
design, which should meet the related regulatory codes and standards, and to review

the safety of the prelirninary designs which include the principles and concepts oft,he plant's design and the implementation of the regulatory criteria, and also toevaluate the environmental impact and the strategy for minimizing its effects.

The main documents required for the application of CP consist of Preliminary

Safety Analysis Report (PSAR) and Environmental Report (ER). Site Survey Rcportand Detailed Geological Survey Report are also required for the application of Early

Site Approval and Limited Work Authorization before the construction pcrmit.

e Operating License (OL)For thc OL, thc safcty of NPP opcration is cvaluatcd by confimlin~ lhat thc-n-final dcsign mcets the acccptancc critoria. Thc main documcnts rcquired of thc OL

application are Fina] Safety Analysis Report (FSAR), Technica] Specifications,Radiological Emergency Plan.
When changes or modifications are to be made on the nuclear facilities by theutility, a relevant Safety Analysis Report should be submitted, in accordance with theAtomic Energy Act, to the KINS through the MOST. The KINS performs atechnical review and reports to the MOST on the results. Then the MOST issues apermit to the applicant based on the KINS' report.

e CANDU Licensing
Korean regulatory system employs two-step licensing approach based on theprescriptive regulation similar to that of the U.S. However, the Canadian regulatory

system employs a one-step licensing approach based on the skills and technicalconsultations between the regulatory organization and the nuclear industry .
Particularly, for CANDU rcactors, the safety of plant design is reviewed continuously

even after the issuance of license in Canada. In Korca, however, a two-step
licensing approach is now official, even for C ANDU reactors starting from Wolsong

Unit 2 and the formats and contents of Safety Analysis Report are required tofollow the specifics prescribed in the U.S. Regulatory Guide I .70.

- 12 -

5. Regulatory Inspection
The regulatory inspection on the performance of a reactor facility is stipulated

in Article 16 of the Atomic Energy Act. The regulatory inspection embraces thepre-operational inspoction, the periodical inspection, the daily inspoction conducted byresident inspoctors, the special inspoction, and QA inspoction, which are explained in
more detail as follows :o Pre-operational Inspoction

The applicant shall be subj ect to the pre-operational inspection to prove thatthe performance of the reactor facility meets the safety requirements spocified in therelevant technical standards . If the utility, the licensee for construction andoporation, passes all the pre-operational inspections, then the Operating License isofficially issued by the MOST.
If the functional test rcsults are unsatisfactory and the inspectors decide thatproper corrections should certainly bc made to improve the performance of theequipment and components, official fmdings or recommendations are issued, which
are followed up with and resolved in order for the utility to pass the pre-operation
mspectlon.

e Regulatory Periodical Inspection (RPI)

In accordance with the provision of the Enforcement Decree of the Atomic
Energy Act, the licensee shall be subject to the regulatory periodical inspection,
which is usually conducted on an annual basis. This inspection should show that theperformance of the reactor facility, designed to withstand the pressure, radiation andother operating environments, is actually maintained in the state in which the reactor

facility passed the pre-operational inspection. The Government then issues a license
for operation.

The RPI consists of standard periodical inspection items (usually 50 - 60
items) which are established for each reactor type (refer to Table 3 & 4) andincluded are some spccial inspection items that are strategically selectcd bascd on thc

opcrating history and also on the cxpcrienccs gaincd from thc prcvious operating

cyclcs. Thc inspcction itcms arc dcvclopcd to covcr all thc tcchnica] arcas andopcrationa] aspccls of a plant. In thc coursc of RPl, thc inspcclors a]so rcvicw thc- 13 -

ncccssary uti]ity documcnts, obscrvc thc utility activitics and cvaluatc thc
maintenancc and tcst records.

e Daily Inspection by Resident Inspectors

Operational safety of an NPP is continuously monitored through daily
inspections at the plant site by the resident inspectors. The Resident Inspectors
Office at each plant site consists of several goverrLment officers and KINS personnel.

They monitor the safety parameters and review the station logs everyday to conflrm
whether the plant is oporated in compliance with the technical specifications. They
also routinely witness the safety-related functional tests such as start-up tests for

emergency diesel generators, etc . If an event occurs, the resident inspectors
investigate the event and report it immediately to the MOST. If the KlNS and the
MOST decide that the event is saf.-ty-significant then an in-depth investigation isnecessary, and a j oint special inspection team will bc organized and assigned to thesrte.

e Quality Assurance (QA) Inspection

Tbis QA Ihspection is performed on ' an annual basis by the KlNS staff tocheck whether the quality assurance activities of the utility are calxied out inaccordance with the QA program submitted to the regulatory authority.

- 14-

StageKorea Electric Power Corporation

MirListry of Scicncc &KJNS(Electric Utility)

Technolog yI

Apply for Prior Aprov

I

I

Assess Safetyof Construction Site

l

I

>Cl)
~~--•-':

Site Survey Report

I

:f:'

l-

CJl
I

~~~

I

Detailed Geologicai I

:3 ~'
FF
(t)

Survey Report t

I

l

I

l

Assess Safety

Apply for Limited Work

6' r

l

I

Asscssmcnl Rcport

Grant Approval of
Limitecf Work

l
I

I

I

l

CL

~
O
~

I

I

O1(JO

F~~)

Assessment Rcport

Grant Prior Approval ofConstruction Srte

Limited Work
(Foundation Work)

I

I
I
I

I

I

l

t

l
l

Pig. 4 Liccnsing Proccdurc for Nuclear Powcr Planls in Korcn

BStngeKorea Electric Power Corporation

Ministy of Science &TechJlo]ogy

KINS(Electric Utility)
I

A

I

l

I

l

Dclibcrace byI

f

Asse8s SAfety

Apply for Cont~Jction Permi

Atomic Eners,

Commissiont

I

Envilonmental Repo

PS AR

B

l

l

Assecsment ReportIll

Issuo Construct,on
Por ftI~

Submit Qu8Jity Assurance

As$ess Safety

Prograrns for ConstrlJction

(,b
I

~,

l
l
l

Ins tionl-

Assessment Report

I

Apply for Preoperationai

l

I

L
t

1'

f

,

t

l

I

l

Installation Work
I

I

I

l

I
I

I
I

t

Conductspecu

l

Ins uonRe rt I

Ord erCorrections or

Submit D~ign DrawingsoPig, 4

t

I

I

I

I

I

I

t

I

I

,

I

I

l

Complornents

Liocnsing Procedure for Nuclear Power Plnnts in Korea (continued)DStageKorca Electric Power Corporatioft

Ministy of Scicnce &KD~S(Electric Vtiiity)

Technology

I

c

I

l

l

l

Assess S8!ety

Atomic Encrs

IlAssessment Report

Operational Teehnica i

l

,

H

~t
l

~,
'~

Commissionl'peeifications

~S AR

l

Delibcrate by

I,

Apply for Operating License

~

D

I

I

Submit Quali~l~sturance
Prograrnt for ration

Is8uo Oporating
LlconseI

Atsest Safecyt,lAssessment Report,.

,

Appoint ar]d Notify
Responsible Person for
Reactor Operation

I

~

I

l

I

I

I

l

I

l

I

l

Commercial OperationlI

onductAppiy for Periodic8itInspection

Ins ctioht,Ilns ction Re

rt I

I

I

I

l

I

l

Order
Correctlons orFig, 4llIlComplernenls

Llccnsing Proecdure for Nuclcar Powcr Plaslts in Korca (continticd)

T~~rl._=~ ~__~~~nda_r~_~pri9~i9~l_Ins ection Items(900 MWe P_vy~)

A: Witness,- 18 -

B: Test Result Review

- 19 -

Tab_1_e i

S_t_an__dard P~llodical Ins ection Ite.=_.m_s(~4__NP_~_ ~HWR)
A: Witness,- 20 -

B: Test Result Review

- 21 -

6. Future Perspectives and Concluding Remarks
In Korea, " Continued Improvement of the Safety of Nuclear Power Plants"
has been set as national policy for the construction and operation of NPPs. Safety

related regulatory requirements under the Korea's atomic energy laws are strictly

observed in the design, fabrication, construction, and operation of NPPs. The
regulatory objective for nuclear safety can simply bc termed into how to maintainand improve, where appropriate, the safety level of nuclear power plants and how to
prepare for emergency situations. In order to achieve this objective, the regulatory
authority of Korean government has established a set of safety requirements in termsof laws and regulations, policies, guidelines and programs.

Safety regulations will bc improved further through th,e improvement ofregulatory systems, introduction of new regulatory concepts, efficient management ofregulatory organizations, and rationalization of technical standards. In particular, thefollowing improvements are mainly considered :(a) Strengthening of Q/A program in preparation of extensive localization ofcomponent design and manufacturing .
(b) Irirprovement of licensing system incorporating one-step licensing, whichallows certification for standard designs and comprehensive site approval.

(c) Severe accident rule making and criteria established for next-generation
reactors.

(d) Introduction of risk based regulation and license, renewal.

(e) Development of safety assessment code systems.

Nuclear safety research will be conducted m the areas of NPP safety,
environmental radioactivity, and the treatment and disposal of radioactive wastes, inorder to maintain and further improve the safety levels. The safety of NPPs can befurther improved by upgrading the qualification of operating personnel throughtraining and education, and by strengthening the nuclear safety culture.

Korea has carried out a very ambitious nuclear power program since the early

1970s with a strong commitment to nuclear power development as an integral partof the national energy policy. However, thc diversification of rcactor types andvendors has caused some difficulties in regulating and licensing thesc nuclear po\vcr

plants in Korea. Thesc difficultics have been coped with through a continucd cffon
to cstablish domcstic rcgu]atory positions and guidclincs as wcll as forTnin*'_' a closc- 22 -

vcndor country's

c oopcration

withparticular]ywith the IAEA.

rcgulatoryorgan i zations,v cn d ors

thcmsclvcs, andAs the standardization of nuclear power plants and the dcvelopmcnt of ncxt
generation reactors, and the establishment of industrial codes and standards applicable

to Korean nuclear power plants are making progress, the licensing system in Korea
is expected to be stabilized. All these efforts will be continued in the future with an

aim to make the nuclear power safer and more reliable.

- 23 -

KlN S ORGANIZATION
CHART
Audi tor

Prcsidcnt

Committee

Steering Committee

S af etyIt~~Vice PresidentlNuclear

NuclearS af ety

Regulator y

Review & Ass.
Div.

Inspection
Div.

Radiation &NuclearEnvironment

S af etyDiv.

Tech. Dlv.

Adrninistration

Div.