ML032930103

to WCAP-15813-NP, Structural Integrity Evaluation of Reactor Vessel Upper Head Penetrations to Support Continued Operation: Millstone, Unit 2.

ML032930103
Person / Time
Site:	Millstone
Issue date:	08/31/2003
From:	Ching Ng, Swamy S Westinghouse
To:	Document Control Desk, Office of Nuclear Reactor Regulation
References
FOIA/PA-2005-0108 WCAP-15813-NP, Rev. 0
Download: ML032930103 (106)

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l Westinghouse Non-Proprietary Class 3 WCAP-15813-NP August 2003 Revision 0 Structural Integrity Evaluation of Reactor Vessel Upper Head Penetrations to Support Continued Operation:

Millstone Unit 2

• @Westinghouse

Westinghouse Non-Proprietary aass 3 WCAP-15813-NP Revision 0 Structural Integrity Evaluation of Reactor Vessel Upper Head Penetrations to Support Continued Operation:

Millstone Unit 2 D. Tang August 2003 Reviewer: -

C.Ng Piping Analysis & Fracture Mechanics Approved: _PA i%/ aft £/J i S. A. Swarny, Manager

, Piping Analysis & Fracture Mechanics This coincides with Rev. 1 of Prop Version.

Westinghouse Electric Company LLC P.O. Box 355 Pittsburgh, PA 15230-0355

©2003 Westinghouse Electric Company LLC All Rights Reserved

iii TABLE OF CONTENTS 1 INTRODUCTION ............................................... 1-1 2 HISTORY OF CRACKING IN HEAD PENETRATION NOZZLES ...................................... 2-1 3 OVERALL TECHNICAL APPROACH ................................................ 3-1 3.1 PENETRATION NOZZLE STRESS ANALYSIS ............................................... 3-1 3.2 FLAW TOLERANCE APPROACH ............................................... 3-1 4 MATERIAL PROPERTIES, FABRICATION HISTORY AND CRACK GROWTH PREDICTION.. 4-1 4.1 MATERIALS AND FABRICATION .4-1 4.2 CRACK GROWTH PREDICTION .............................. :4-1 5 STRESS ANALYSIS ............................... 5-1 5.1 OBJECTIVES OF THE ANALYSIS ....................................................... 5-1 5.2 MODEL ....................................................... 5-1 5.3 STRESS ANALYSIS RESULTS - OUTERMOST CEDM PENETRATION

- NOZZLE (42.5 DEGREES) .................. 5-1 5.4 STRESS ANALYSIS RESU LTS 37.1 DEGREE AND 29.1 DEGREE CEDM PENETRATION NOZZLES.....................................................................................5-2 5.5 STRESS ANALYSIS RESULTS-CENTERCEDM PENETRAT ION NOZZLE. .......... 5-2 5.6 STRESS ANALYSIS RESULTS: ICI NOZZLES ..................................................... 5-2 5.7 STRESS ANALYSIS RESULTS: HEAD VENT NOZZLE ........................................ 5-2 6 FLAW EVALUATION CHARTS .............................. 6-1

6.1 INTRODUCTION

.................................. 6-1 6.2 OVERALLAPPROACH .................................. 6-1

6.3 RESULTS

AXIAL FLAWS .................................. 6-2 6.4 CIRCUMFERENTIAL CRACK PROPAGATION....................................................6-4 6.5 FLAW ACCEPTANCE CRITERIA .................................. 6-5 7

SUMMARY

AND EXAMPLE PROBLEMS .. ............................ 7-1 7.1 SAFETY ASSESSMENT .............................. 7-1 7.2 EXAMPLE PROBLEMS .............................. 7-2 8 REFERENCES .............................. 8-1 APPENDIX A ALLOWABLE AREAS OF LACK OF FUSION: WELD FUSION ZONES .. A-1 APPENDIX B CEDM AND ICI NOZZLE HOOP STRESS VS DISTANCE FROM BOTTOM OF WELD PLOTS .B-1 August 2003 5874(copy).dac.100603 5874(copy).doc-100603 August 2003

iv LIST OF TABLES Table 1-1 Millstone Unit 2 Head Penetration Nozzles with the Intersection Angles Identified .......................................................... 1-3 Table 4-1 Millstone Unit 2 Reactor Vessel Head Penetration Nozzle Information ........................ 4-6 Table 6-1 Summary of R.V. Head Penetration Acceptance Criteria (Limits for Future Growth) ....6-8 Table 6-2 Millstone Unit 2 Penetration Geometries .......................................................... 6-8 Table 7-1 Example Problem Inputs: Initial Flaw Sizes and Locations .................... .................... 7-5 August 2003 5574.doc-100603 5974.doc-100603 August 2003

V LIST OF FIGURES K> Figure 1-1 Reactor Vessel Control Element Drive Mechanism (CEDM) Penetration ..................... 1-4 Figure 1-2 Location of Head Penetrations for Millstone Unit 2.................................................... 1-5 Figure 2-1 EDF Plant RN Closure Head CEDM Penetrations - Penetrations with Cracking ......... 2-4 Figure 2-2 Inspection Results for U.S. CRDM1CEDM Penetration Nozzle (Spring 2003) ............. 2-5 Figure 3-1 Schematic of a Head Penetration Flaw Growth Chart for Part-Through Flaws ............. 3-3 Figure 3-2 Schematic of a Head Penetration Flaw Tolerance Chart for Through-Wall Flaws .......... 3-4 Figure 4-1 Yield Strength of the Various Heats of Alloy 600 Used in Fabricating the Millstone Unit 2 and French Head Penetrations .4-7 Figure 4-2 Carbon Content of the Various Heats of Alloy 600 Used in Fabricating the Millstone Unit 2 and French Head Penetrations .4-8 Figure 4-3 Screened Laboratory Data for Alloy 600 with the MRP Recommended Curve (Note that the Modified Scott Model is also Shown) .4-9 Figure 4-4 Model for PWSCC Growth Rates in Alloy 600 in Primary Water Environments (325°C),

With Supporting Data from Standard Steel, Huntington, and Sandvik Materials . 4-10 Figure 4-5 Surnmmary of Temperature Effects on PWSCC Growth Rates for Alloy 600 in Primary Water . 4-11 Figure 5-1 Finite Element Model of CEDM Penetration (42.5 Degrees) . 5-3 Figure 5-2 Three Dimenstional Model of the ICI Penetration Nozzle (54.8 Degrees) . 5-4 Figure 5-3 Stress Distribution at Steady State Conditions: Outermost CEDM Penetration Nozzle (42.5 Degrees) (Hoop Stress is the Top Figure; Axial Stress is the Bottom Figure) . 5-5 Figure 5-4 Axial Stress Distribution at Steady State Conditions for the Outermost CEDM (42.5 Degrees) Penetration, Along a Plane Oriented Parallel to, and Just Above, the Attachment Weld .5-6 Figure 5-5 Stress Distribution at Steady State Conditions for the 37.1 Degrees CEDM Penetration (Hoop Stress is the Top Figure; Axial Stress is the Bottom Figure) .5-7 Figure 5-6 Stress Distribution at Steady State Conditions for the 29.1 Degrees CEDM Penetration (Hoop Stress is the Top Figure; Axial Stress is the Bottom Figure) .5-8 Figure 5-7 Stress Distribution at Steady State Conditions for the Center CEDM Penetration (Hoop Stress is the Top Figure; Axial Stress is the Bottom Figure) .5-9 Figure 5-8 Stress Distribution at Steady State Conditions for ICI Penetration (Hoop Stress is the Top Figure; Axial Stress is the Bottom Figure) ............................... 5-10 Figure 5-9 Head Vent Nozzle Finite Element Model ............................... 5-11 Figure 5-10 Stress Contours in the Head Wnt Nozzle as a Result of Residual Stresses and Operating Pressure (Hoop Stress is the Top Figure; Axial Stress is the Bottom Figure).5-12 Figure 6-1 Stress Intensity Factor for a Through-Wall Circumferential Flaw in a Head Penetration Nozzle .6-9 Figure 6-2 Inside, Axial Surface Flaws, .5" Below the Attachment Weld, Nozzle Uphill Side -

Crack Growth Predictions .6-10 Figure 6-3 Inside, Axial Surface Flaws, .5" Below the Attachment Weld, Nozzle Downhill Side -

Crack Growth Predictions .6-11 Figure 6-4 Inside, Axial Surface Flaws, At the Attachment Weld, Nozzle Uphill Side - Crack Growth Predictions .6-12 Figure 6-5 Inside, Axial Surface Flaws, At the Attachment Weld, Nozzle Downhill Side - Crack Growth Predictions .6-13

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Vi LIST OF FIGURES (Cont.)

Figure 6-6 Inside, Axial Surface Flaws, .5" Above the Attachment Weld, Nozzle Uphill Side -

Crack Growth Predictions ................................................... 6-14 Figure 6-7 Inside, Axial Surface Flaws, .5" Above the Attachment Weld, Nozzle Downhill Side -

Crack Growth Predictions ................ 6-15 Figure 6-8 Outside, Axial Surface Flaws, Below the Attachment Weld, Nozzle Uphill Side -

Crack Growth Predictions ................ 6-16 Figure 6-9 Outside, Axial Surface Flaws, Below the Attachment Weld, Nozzle Downhill Side -

Crack Growth Predictions ................ 6-17 Figure 6-10 Outside, CircumferentialSurface Flaws, Along the Top of the Attachment Weld -

Crack Growth Predictions (MRP Factor of 2.0 Included) ...................................... 6-18 Figure 6-11 Inside, Axial Surface Flaws, At the Attachment Weld, Head Vent, Nozzle Downhill Side - Crack Growth Predictions.: ...................... 6-19 Figure 6-12 Through-Wall Axial Flaws Located in the Center CEDM (0.0 Degrees) Penetration, Uphill and Downhill Side - Crack Growth Predictions .................................... 6-20 Figure 6-13 Through-Wall Axial Flaws Located in the 29.1 Degree Row of Penetrations, Downhill Side - Crack Growth Predictions .................... 6-21 Figure 6-14 Through-Wall Axial Flaws Located in the 29.1 Degree Row of Penetrations, Uphill Side -Crack Growth Predictions .................... 6-22 Figure 6-15 Through-Wall Axial Flaws Located in the 37.1 Degree Row of Penetrations, Downhill Side - Crack Growth Predictions .................... 6-23 Figure 6-16 Through-Wall Axial Flaws Located in the 37.1 Degree Row of Penetrations, Uphill Side - Crack Growth Predictions .................... 6-24 Figure 6-17 Through-Wall Axial Flaws Located in the 42.5 Degree Row of Penetrations, Downhill Side - Crack Growth Predictions .................... 6-25 Figure 6-18 Through-Wall Axial Flaws Located in the 42.5 Degree Row of Penetrations, Uphill Side - Crack Growth Predictions .................... 6-26 Figure 6-19 Through-Wall Axial Flaws Located in the 54.8 Degree Row of Penetrations, Downhill Side - Crack Growth Predictions ..................... 6-27 Figure 6-20 Through-Wall Axial Flaws Located in the 54.8 Degree Row of Penetrations, Uphill Side - Crack Growth Predictions ..................... 6-28 Figure 6-21 Through-Wall Circumferential Flaws Near the Top of the Attachment Weld for CEDM and ICI Nozzles - Crack Growth Predictions (MRP Factor of 2.0 Included) .6-29 Figure 6-22 Section XI Flaw Proximity Rules for Surface Flaws (Figure IWA-3400-1) .6-30 Figure 6-23 Definition of "Circumferential".6-31 Figure 6-24 Schematic of Head Penetration Geometry .6-32 Figure 7-1 Example Problem I .B-2 Figure 7-2 Example Problem 2 .B-3 Figure 7-3 Example Problem 3. B4 Figure 74A Example Problem 4 (Also See Figure 7-4B) .B-5 Figure 74B Example Problem 4 (Also See Figure 74A) .B-6 Figure 7-5 Example Problem 5.B-7 Figure A-I Typical Head Penetration . A-3 Figure A-2 Allowable Regions of Lack of Fusion for the Outermost Penetration Tube to Weld Fusion Zone: Detailed View ..................... A-4 584dc100 Augst00 5874.doc-100603 August 2003

vii LIST OF FIGURES (Cont.)

Figure A-3 Allowable Regions of Lack of Fusion for the Outermost Penetration Tube to Weld Fusion Zone .................................................... A-5 Figure A4 Allowable Regions of Lack of Fusion for all Penetrations: Weld to Vessel Fusion Zone .A-6 Figure A-5 Allowable Regions of Lack of Fusion for the Weld to Vessel Fusion Zone .A-7 Figure B-I Hoop Stress Distribution Below the Weld Downhill and Uphill Side (00 CEDM Penetration Nozzle)................................................................................ B-2 Figure B-2 Hoop Stress Distribution Below the Weld Downhill Side (29.10 CEDM Penetration Nozzle) ...................... B-3 Figure B-3 Hoop Stress Distribution Below the Weld Uphill Side (29.1° CEDM Penetration Nozzle) ...................... B4 Figure B4 Hoop Stress Distribution Below the Weld Downhill Side (37.10 CEDM Penetration Nozzle) ...................... B-5 Figure B-5 Hoop Stress Distribution Below the Weld Uphill Side (37.10 CEDM Penetration Nozzle) ...................... B-6 Figure B-6 Hoop Stress Distribution Below the Weld Downhill Side (42.50 CEDM Penetration Nozzle) ...................... B-7 Figure B-7 Hoop Stress Distribution Below the Weld Uphill Side (42.50 CEDM Penetration Nozzle) ...................... B-8 Figure B-8 Hoop Stress Distribution Below the Weld Downhill side (54.8° ICI Penetration Nozzle)................................................................................. B-9 Figure B-9 Hoop Stress Distribution Below the Weld Uphill side (54.8° ICI Penetration Nozzle)............................................................................... 1B-10 K>

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1-1 I INTRODUCTION In September of 1991, a leak was discovered in the Reactor Vessel Control Rod Drive Mechanism (CRDM) head penetration region of an operating plant. This has led to the question of whether such a leak could occur at the Millstone Unit 2 Control Element Drive Mechanism (CEDM), In-Core Instrumentation (ICI), or head vent nozzle penetrations. Note that the designation CRDM (Westinghouse and French designs) and CEDM (Combustion Engineering Design) are synonymous. The geometry of interest is shown in Figure 1-1. Throughout this report, the penetration rows have been identified by their angle of intersection with the head. The location of head penetrations for Millstone Unit 2 is shown in Figure 1-2 and the angles for each

,penetration are identified in Table 1-1.

The CEDM leak resulted from cracking in Alloy 600 base metal, which occurred in the outermost penetrations of a number of operating plants as discussed in Section 2. This outermost CEDM location, as well as a number of intermediate CEDM locations, the ICI nozzles, and the head vent were chosen for fracture mechanics analyses to support continued safe operation of Millstone Unit 2 if such cracking were to be found. The dimensions of the CEDM penetrations are all identical, with a 3.850 inch Outside Diameter (OD) and a wall thickness of 0.566 inch [bOA, IOC]. The ICI penetrations have an OD of 5.563 inch and wall thickness of 0.469 inch [10A, IOC]. The head vent OD is 1.050 inch and the wall thickness is 0.154 inch [bOA, IOC]. All of these dimensions are summarized in Table 6-2.

The basis of the fracture analysis was a detailed three-dimensional elastic-plastic finite element analysis of several penetration locations, as described in detail in Section 5. The fracture analysis was carried out using crack growth rates recommended by the EPRI Materials Reliability Program (MRP). These rates are consistent with service experience. The results are presented in the form of flaw tolerance charts for both surface and through wall flaws. If indications are found, the charts will determine the allowable service life of safe operation. The service life calculated in the flaw tolerance charts are all in Effective Full Power Years (EFPY).

Introduction August 2003 Revision 0

1-2 The flaw tolerance handbook WCAP-15813 Revision 1 is updated to reflect the following development in the Primary Water Stress Corrosion Cracking of Alloy 600:

1. New crack growth rates for evaluation Primary Water Stress Corrosion Cracking in CRDM penetration nozzle provided in MRP-55 Rev. I (Reference 4H).

2. MRP-55 Rev. I recommends the use of a factor of 2.0 for OD flaws above the attachment weld to account for uncertainty in the chemical environment of the penetration nozzle annulus. This factor was not included for OD flaws above the attachment weld in the original WCAP report.

3. Based on recent experiences with vessel head inspections of a number of power plants NRC has been requesting hoop distribution below the attachment weld if 100% of the penetration nozzle is not feasible. These hoop stress distributions will be included in the updated flaw tolerance handbook and can be used to support the submittal of relaxation request from the NRC Issued Order EA-03-009 for less than 100% inspection coverage of the penetration nozzle.

K>

August 2003 Introduction August 2003 Revision 0

1-3 Table 1-1 Millstone Unit 2 Head Penetration Nozzles, with the Intersection Angles Identified I I V YI 7 I Nozzle Angle Nozzle Angle Nozzle Angle No. Type (Degrees) No. Type (Degrees) No. Type (Degrees)

I CEDM 0.0 27 CEDM 29.1 53 CEDM 37.1 2 CEDM 11.0 28 CEDM 29.1 54 CEDM 37.1 3 CEDM 11.0 29 CEDM 29.1 55 CEDM 37.1 4 CEDM 11.0 30 CEDM 29.1 56 CEDM 37.1 5 CEDM 11.0 31 CEDM 29.1 57 CEDM 37.1 6 CEDM 11.0 32 CEDM 29.1 58 CEDM 42.5 7 CEDM 11.0 33 CEDM 29.1 59 CEDM 42.5 8 CEDM 11.0 34 CEDM 35.6 60 CEDM 42.5 9 CEDM 11.0 35 CEDM 35.6 61 CEDM 42.5 10 CEDM 22.4 36 CEDM 35.6 62 CEDM 42.5 11 CEDM 22.4 37 CEDM 35.6 63 CEDM 42.5 12 CEDM 22.4 38 CEDM 35.6 64 CEDM 42.5 13 CEDM 22.4 39 CEDM 35.6 65 CEDM 42.5 14 CEDM 23.9 40 CEDM 35.6 66 CEDM 42.5 K' 15 CEDM 23.9 41 CEDM 35.6 67 CEDM 42.5 16 CEDM 23.9 42 CEDM 38.5 68 CEDM 42.5 17 CEDM 23.9 43 CEDM 38.5 69 CEDM 42.5 18 CEDM 25.3 44 CEDM 38.5 70 ICI 54.8 19 CEDM 25.3 45 CEDM 38.5 71 IC I 54.8 20 CEDM 25.3 46 CEDM 38.5 72 T CI 54.8 21 CEDM 25.3 47 CEDM 38.5 73 1IC 54.8 22 CEDM 25.3 48 CEDM 38.5 74 T CI 54.8 23 CEDM 25.3 49 CEDM 38.5 75 ICI 54.8 24 CEDM 25.3 50 CEDM 37.1 76 ICI 54.8 25 CEDM 25.3 51 CEDM 37.1 77 ICI 54.8 26 CEDM 29.1 52 CEDM 37.1 Introduction August 2003 Revision 0

1-4

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By D l

~Downhilll

[.~j \Cladding (00) J Weld Head Penetration Nozzle Figure 1-1 Reactor Vessel Control Element Drive Mechanism (CEDM) Penetration K>~

Introduction August 2003 Revision 0

1-5

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Figure 1-2 Location of Head Penetrations for Millstone Unit 2 Introduction August 2003 Revision 0

2-1 2 HISTORY OF CRACKING IN HEAD PENETRATION NOZZLES In September of 1991, leakage was reported from the reactor vessel CRDM head penetration region of a French plant, Bugey Unit 3. Bugey 3 is a 920 megawatt three-loop Pressurized Water Reactor (PWR) plant which had just completed its tenth fuel cycle. The leak occurred during a post ten year hydrotest conducted at a pressure of approximately 3000 psi (204 bar) and a temperature of 194 0 F (90'C). The leak was detected by metal microphones, which are located on the top and bottom heads. The leak rate was estimated to be approximately 0.7 liter/hour. The location of the leak was subsequently established on a peripheral penetration with an active control rod (H-14), as seen in Figure 2-1.

The control rod drive mechanism and thermal sleeve were removed from this location to allow further examination. A study of the head penetration revealed the presence of longitudinal cracks near the head penetration attachment weld. Penetrant and ultrasonic testing confirmed the cracks.

The cracked penetration was fabricated from Alloy 600 bar stock (SB-166), and has an outside diameter of 4 inches (10.16 cm) and an inside diameter of 2.75 inches (7.0 cm).

As a result of this finding, all of the control rod drive mechanisms and thermal sleeves at Bugey 3 were removed for inspection of the head penetrations. Only two penetrations were found to have cracks, as shown in Figure 2-1.

An inspection of a sample of penetrations at three additional plants were planned and conducted during the winter of 1991-92. These plants were Bugey 4, Fessenheim 1,and Paluel 3. The three outermost rows of penetrations at each of these plants were examined, and further cracking was found in two of the three plants.

At Bugey 4, eight of the 64 penetrations examined were found to contain axial cracks, while only one of the 26 penetrations examined at Fessenheim 1 was cracked. The locations of all the cracked penetrations are shown in Figure 2-1. At the time, none of the 17 CRDM penetrations inspected at Paluel 3 showed indications of cracking, however subsequent inspections of the French plants have confirmed at least one crack in each operating plant.

Thus far, the cracking in reactor vessel heads not designed by Babcock and Wilcox (B&W) has been consistent in both its location and extent. All cracks discovered by nondestructive examination have been oriented axially, and have been located in the bottom portion of the penetration in the vicinity of the partial penetration attachment weld to the vessel head as shown schematically in Figure 1-1.

ace History of Cracking in Head Penetration Nozzles

2-2 Non-destructive examinations of the leaking CRDM nozzles showed that most of the cracks were axially oriented, originating on the outside surface of the nozzles below the J-groove weld and propagating primarily in the nozzle base material to an elevation above the top of the J-groove weld. Leakage could then pass through the annulus to the top of the head where it was detected by visual inspection. In some cases the cracks initiated in the weld metal or propagated into the weld metal, and in a few cases the cracks propagated through the nozzle wall thickness to the inside surface.

[

]acle History of Cracking in Head Penetration Nozzles August 2003 Revision 0

2-3 The cracking has now been confirmed to be primary water stress corrosion cracking. Relatively high residual stresses are produced in the outermost CRDM penetrations due to the welding process. Other important factors which affect this process are temperature and time, with higher temperatures and longer times being more detrimental. The inspection findings for U.S. plants are shown in Figure 2-2. From this figure, it is interesting to note that low percentage of CE-fabricated vessels have been found cracked (only 9 of 1332 penetrations UT and/or ET inspected in CE-fabricated vessels have shown cracking whereas 117 of 628 penetrations in non-CE-fabricated vessels experienced cracking or leakage). In addition, no cracks in head vent have been found History of Cracking in Head Penetration Nozzles

2-4 270' 270-O 0.

T-F 0 Lee~k"n Pefietr-.tion 90

• Cra-cked Penetrtioton 9'~ 0 Ciackced Penetration BUGEY 3 BUGEY 4 270e 90-

• Cr-acked Penetratmon FESSENHEIM I Figure 2-1 EDF Plant RN Closure Head CEDM Penetrations- Penetrations with Cracking K-o zeu ut2 0 of Cracking Hitof e d P ntato

_Crci giin Head Penetration Nozzles August 2003 History Revision 0

) / 2-5 CRDMICEDM Penetrations Inspected CRDM/CEDM Inspection Results 3500- 1400 3000- 1200 2500- 1000 E

0 I0 200- 800

0. 0.

600 i

E E z a Z

1000- 400, 200 01 CE Vessels Non-CE Vessels CE Vessels Non-CE Vessels I

• UT and/or ET

. _ _. _. ... _ . A....

URMV Only

• Not YAt Insoacted I

I t UT/FT wtih Nn Cracks

• Cracked {not Ibaker)

• Laker I Figure 2-2 Inspection Results for U.S. CRDMICEDM Penetration Nozzle (Spring 2003) co!

History of Cracking in Head Penetration Nozzles August 2003 Revision 0

3-1 3 OVERALL TECHNICAL APPROACH K> The primary goal of this work is to provide technical justification for the continued safe operation of Millstone Unit 2 in the event that cracking is discovered during inservice inspections of the Alloy 600 reactor vessel head penetrations.

3.1 PENETRATION NOZZLE STRESS ANALYSIS Three-dimensional elastic-plastic finite element stress analyses applicable to Millstone Unit 2 was performed to determine the stresses in the head penetration region [6A, 6B, 6C]. These analyses have considered the pressure loads associated with steady state operation, as well as the residual stresses that are produced by the fabrication process.

3.2 FLAW TOLERANCE APPROACH A flaw tolerance approach has been developed to allow continued safe operation until an appropriate time for repair, or the end of plant life. The approach is based on the prediction of future growth of detected flaws, to ensure that such flaws would remain stable.

If an indication is discovered during in-service inspection, its size can be compared with the flaw size considered as allowable for continued service. This "allowable" flaw size is determined from the actual loading (including mechanical and residual loads) on the head penetration for Millstone Unit 2. Acceptance criteria are discussed in Section 6.5.

The time for the observed crack to reach the allowable crack size determines the length of time the plant can remain online before repair, if required. For the crack growth calculation, a best estimate is needed and no additional margns are necessary.

The results of the evaluation are presented in terms of simple flaw tolerance charts. The charts graphically show the time required to reach the allowable length or depth, which represents additional service life before repair. This result is a function of the loading on the particular head penetration as well as the circumferential location of the crack in the penetration nozzle.

Overall Technical Approach August 2003 Revision 0

3-2 Schematic drawings of the head penetration flaw tolerance charts are presented as Figures 3-1 and

.I 3-2. These two types of charts can be used to provide estimates of the remaining service life before a leak would develop from an observed crack. For example, if a part-through flaw was discovered, the user would first refer to Figure 3-1, to determine the time (tp) which would be remaining before the crack would penetrate the wall or reach the allowable depth (t.) (e.g. a/t =

0.75). Once the crack penetrates the wall, the time (tB) required to reach an allowable crack length would be determined from Figure 3-2. The total time remaining would then be the simple sum:

Time remaining = tp + tB Another way to determine the allowable time of operation with a part-through flaw would be to use Figure 3-2 directly, in effect assuming the part-through flaw is a through-wall flaw. This approach would be more conservative than that above, and the time remaining would then be:

Time remaining= to Overall Technical Approach August 2003 Revision 0

3-3 Flaw Becomes Through - Wall 1.0 _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _

cLj= .75

---

-7 '_

Q. . Allowable Time (IAs,) Bef a}

U- I-

.3. Reaching Allowab Depth

/gL Detected Indication I Allowable Time Before Depth Wall Penetration. II Time ( Months )

Figure 3-1 Schematic of a Head Penetration Flaw Growth Chart for Part Through Flaws Overall Technical Approach ----

August 2003 Revision 0

3-4 Critical Length ( Excessive Leakage)

Allowable Length U.

Detected Indication C~~Lnt Leng)

I rime ( Months )

Figure 3-2 Schematic of a Head Penetration Flaw Tolerance Chart for Through-Wall Flaws Overall Technical Approach August 2003 Revision 0

4-1 4 MATERIAL PROPERTIES, FABRICATION HISTORY AND CRACK GROWTH PREDICTION 4.1 MATERIALS AND FABRICATION The reactor vessel for Millstone Unit 2 was manufactured by Combustion Engineering with the head penetration nozzles from material produced by Huntington Alloys in the USA. The carbon content and mechanical properties of the Alloy 600 material used to fabricate the Millstone Unit 2 vessel are provided in Table 4-1. The material CMTRs were used to obtain fie chemistry and mechanical properties for the vessel head penetrations. The CMTRs for the material do not indicate the heat treatment of the material. However, Westinghouse records indicate that the materials were annealed for one hour at a temperature of 17000 - 18000 F, followed by a water quench. Figures 4-1 and 4-2 illustrate the yield strengths and carbon content, based on percent of heats, of the head adapter penetrations in the Millstone Unit 2 vessels relative to a sample of the French head adapters which have experienced cracking. Note that the French plants do not contain the large diameter ICIs, but only tubes the size of CEDMs. The general trends for the head adapter penetrations in Millstone Unit 2 are a higher carbon content, higher mill annealing temperature and lower yield strength relative to those on the French vessels. These factors should all have a beneficial effect on the material resistance to PWSCC in the head penetrations.

4.2 CRACK GROWTH PREDICTION The cracks in the penetration region have been determined to result from primary water stress corrosion cracking in the Alloy 600 base metal, and in some cases the Alloy 182 weld metal.

There are a number of available measurements of static load crack growth rates in primary water environment, and in this section the available results will be compared and a representative growth rate established.

Direct measurements of SCC growth rates in Alloy 600 are relatively rare, and care should be used in interpreting the results because the materials may be excessively cold worked, or the loadings applied may be near or exceeding the limit load of the tube, meaning there will be an interaction between tearing and crack growth. In these cases the crack growth rates may not be representative of service conditions.

The effort to develop a reliable crack growth rate model for Alloy 600 began in the spring of 1992, when the Westinghouse Owners Group was developing a safety case to support continued operation of plants. At the time there was no available crack growth rate data for head penetration materials, and only a few publications existed on growth rates of Alloy 600 in any product form.

The best available publication was found to be that of Peter Scott of Framatome, who had developed a growth rate model for PWR steam generator materials [1]. His model was based on a study of results obtained by Mcllree and Smialowska [2] who had tested short steam generator tubes which had been flattened into thin compact specimens.

Material Properties, Fabrication History and Crack Growth Prediction August 2003 Revision 0

4-2 An equation was fitted to the data of reference [2] for the results obtained in water chemistries that fell within the standard specification for PWR primary water. Results for chemistries outside the specification were not used. The following equation was fitted to the data at 3300 C (6260 F):

.t=2.8xlO (K-9)1* 6 m/sec (4-1) dt where:

K is in MPa4JM-The next step was to correct these results for the effects of cold work. Based on work by Cassagne and Gelpi [3], Scott concluded that dividing the above equation by a factor of 10 would be appropriate to account for the effects of cold work. The crack growth law for 330'C (626TF) then becomes:

da=2.8x10-12(K -9) 1.16m/sec (4-2) dt Scott further corrected this law for the effects of temperature. This forms the basis for the PWR Materials Reliability Program (MRP) recommended crack growth rate (CGR) curve for the evaluation of SCC where a power-law dependence on stress intensity factor was assumed [4H].

The MRP recommended CGR curve was used in this report for determining the primary water stress corrosion crack growth rate and a brief discussion on this recommended curve is as follows:

[

Iaxc.e Prediction August 2003 Material Properties, Fabrication Material Properties, History and Fabrication History Crack Growth and Crack Growth Prediction August 2003 Revision 0

4-3 I~~~~~~~~~~~~~~~~~~~~

There is a general agreement that crack growth in Alloy 600 materials in the primary water environment can be modeled using a stress intensity factor relationship wth differences in temperature accounted for by an activation energy (Arrhenius) model for thermally controlled processes. Figure 4-3 shows the recommended CGR curve along with the laboratory data from Huntington materials used to develop the curve.

[

I ac, Material Properties, Fabrication History and Crack Growth Prediction August 2003 Revision 0

4-4 The applicability of the MRP recommended model to head penetrations was recently confirmed by two independent approaches. The first was a collection of all available data from Standard Steel and Huntington Alloys materials tested over the past ten years [4H]. The results are shown in Figure 4-3, along with the Scott model for the test temperature.

The MRP crack growth curve was structured to bound 75 percent of the 26 heats for which test results were available. Fits were done on the results for each heat, and the constant term was determined for each heat. This was done to eliminate the concern that the curve might be biased from a large number of results from a single heat. The 7 5 th percentile was then determined from these results. The MRP expert panel on crack growth endorsed the resulting curve unanimously in a meeting on March 6h and 7h 2002. This approach is consistent with the Section XI flaw evaluation philosophy, which is to make a best estimate prediction of future growth of a flaw.

Margins are incorporated in the allowable flaw sizes. The entire data set is shown in Figure 4-3, where the data have been adjusted to a single temperature of 325 0 C.

A second independent set of data were used to validate the model, and these data were obtained from the two inspections carried out on penetration no. 75 of D.C. Cook Unit 2, which was first found to be cracked in 1994 [4G]. The plant operated for one fuel cycle before the penetration was repaired in 1996 and the flaw was measured again before being repaired. These results were used to estimate the PWSCC growth rate for both the length of the flaw and its depth. These two points are also shown in Figure 4-4, and are consistent with the laboratory data for Huntington materials. In fact, Figure 4-4 demonstrates that the MRP model is rearly an upper bound for these materials. The D.C. Cook Unit 2 penetrations were made from Huntington materials.

Since Millstone Unit 2 operates at a temperature of 312'C (5940 F) in the head region [11], and the crack growth rate is strongly affected by temperature, a temperature adjustment is necessary.

This temperature correction was obtained from study of both laboratory and field data for stress corrosion crack growth rates for Alloy 600 in primary water environments. The available data showing the effect of temperature are summarized in Figure 4-5. Most of the results shown here are from steam generator tube materials, with several sets of data from operating plants, and results from two heats of materials tested in a laboratory [4A].

Material Properties, Fabrication History and Crack Growth Prediction August 2003 Revision 0

4-5 Study of the data shown in Figure 4-5 results in an activation energy of 31-33 Kcal/mole, which can then be used to adjust for the lower operating temperature. This value is slightly lower than the generally accepted activation energy of 44-50 Kcal/mole used to characterize the effect of temperature on crack initiation, but the trend of the actual data for many different sources is unmistakable.

].C' 'Therefore the following crack growth rate model was used for the Millstone Unit 2 head penetration for crack growth in all the cases analyzed.

da =1.51 x 101-2 (K 9)1.16 MlSeC dt Material Properties, Fabrication History and Crack Growth Prediction August 2003 Revision 0

4-6 where:

K = applied stress intensity factor, in MPa m This equation implies a threshold for cracking susceptibility, Kscc = 9 MPa 4 . The crack growth rate is applicable to propagation in both axial and circumferential directions.

(a.ce Material Properties, Fabrication History and Crack Growth Prediction August 2003 Revision 0

4 ) V-7 40 35 l EdF (11 Heats)

_30 LI Millstone 2 (7 Heats) 0 25

.v 20 0

~15 CL 5

0 Yield Strength (ksi)

Figure 4-1 Yield Strength of the Various Heats of Alloy 600 Used in Fabricating the Millstone Unit 2 and French Head Penetrations Material Properties, Fabrication History and Crack Growth Prediction August 20030 Revision

) )

4-8 60

• EdF (11I Heats) 50

_.O% lOI Millstone 2 (7 Heats)

° 40 0

I

, 30 0

, 20 I~~S....

fIL 0~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~ An'

,,tz' 10 0

Carbon Content (Weight %)

Figure 4-2 Carbon Content of the Various Heats of Alloy 600 Used in Fabricating the Millstone Unit 2 and French Head Penetrations Material Properties, Fabrication History and Crack Growth Prediction Material Properties, Fabrication History and Crack Growth Prediction August 2003

4-9 Figure 4-3 Screene d Laboratory Data for Alloy 600 with the MRP Recommended Curve.

(Note that the Modified Scott Model is also Shown)

Material Properties, Fabrication History and Crack Growth Prediction August 2003 Revision 0

4-10 (ass)

K-Figure 44 Model for PWSCC Growth Rates in Alloy 600 in Primary Water Environments (325 0C), With Supporting Data from Standard Steel, Huntington, and Sandvik Materials Note: That the data have been normalized to a temperature of 325 0C. The actual test temperatures are listed in parenthesis after the caption. For example, the Huntington data were obtained at temperature ranging from 315 0C to 331 0C.

Material Properties, Fabrication History and Crack Growth Prediction August 2003 Revision 0

4-11 TEMPERATURE, DEG. C 372! 352 333 315 298 21B2 1

1E-OE5 i i i i IOOO0 1E-09 100 I-z 0

Co G

G I I 1E-10 _ _ _ _ 0 _ _ _

_ _ -~ _ _ _ _ _ _ _

CI-lE-1l 1 C-)C lF-12

. .- . I n I I I I 0.00155 0.0016 0.00165 0.0017 0.00175 0.0018 RECIPROCAL TEMPERATURE, 1/DEG. K Figure 4-5 Summary of Temperature Effects on PWSCC Growth Rates for Alloy 600 in Primary Water Note:All symbols are for steam generator materials, except the solid circles, which are head penetration laboratory data.

Material Properties, Fabrication History and Crack Growth Prediction August 2003 Revision 0

5-1 5 STRESS ANALYSIS 5.1 OBJECTIVES OF THE ANALYSIS The objective of this analysis was to obtain accurate stresses in each CEDM or ICI housing and its immediate vicinity. To do so requires a three dimensional analysis which considers all the pertinent loadings on the penetration [6A, 6B, 6C]. An investigation of deformations at the lower end of the housing was also performed using the same model. Four CEDM locations were considered: the outermost CEDM row (42.5 degrees), rows at 37.1 degrees, 29.1 degrees, and the center location. In addition, the ICI penetration nozzles (54.8 degrees) were analyzed.

The analyses were used to provide information for the flaw tolerance evaluation in Section 6.

Also, the results of the stress analysis were compared to the findings from service experience, to help assess the causes of the observed cracking.

5.2 MODEL A three-dimensional finite element model comprised of isoparametric brick and wedge elements with mid-side nodes on each face was used to obtain the stresses and deflections. Views of CEDM and ICI models are shown in Figures 5-1 and 5-2 respectively. Taking advantage of the symmetry of the vessel head, only half of the CEDM penetrations were modeled. Similarly, only half of the center penetration was modeled.

In the models, the lower portion of the Control Element Drive Mechanism (CEDM) penetration nozzle, In-Core Instrumentation (ICI) nozzle, the head vent, the adjacent section of the vessel closure head, and the joining weld were modeled. The vessel to penetration nozzle weld was simulated with two weld passes. The penetration nozzle, and weld metal were modeled as Alloy 600 and the vessel head shell as carbon steel.

The only loads used in the analysis are the steady state operating loads. External loads, such as seismic loads, have been studied and have no impact since the penetration nozzles are captured by the full thickness of the reactor vessel head (about 7 and 1/2 inches of steel [IOB]) into which the penetrations are shrunk fit during construction. The area of interest is in the penetration near the attachment weld, which is unaffected by these external loads.

5.3 STRESS ANALYSIS RESULTS - OUTERMOST CEDM PENETRATION NOZZLE (42.5 DEGREES)

Figure 5-3 presents the hoop and axial stresses for the steady state condition for the outermost CEDM penetration.

[~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~

Stress Analysis August 2003 Revision 0

5-2

[

K>

Iac e 5A STRESS ANALYSIS RESULTS 37.1 DEGREE AND 29.1 DEGREE CEDM PENETRATION NOZZLES

]anb 5.5 STRESS ANALYSIS RESULTS-CENTER CEDM PENETRATION NOZZLE 5.6 STRESS ANALYSIS RESULTS: ICI NOZZLES The ICI nozzles are located further from the reactor vessel centerline than any of the other head penetration nozzles. They are larger in diameter and thinner than the CEDMs. The stress contour plots for the ICI nozzles at steady state conditions are shown in Figure 5-8.

5.7 STRESS ANALYSIS RESULTS: HEAD VENT NOZZLE The head vent is a smaller penetration than the CEDM or ICI head penetration nozzles, but is also constructed of Alloy 600 material, with a partial penetration weld at the inside of the reactor vessel head. The head vent is located 8.2 inches from the centerline of the head dome.

[

]sBcle Stress Analysis August 2003 Revision 0

5-3 Figure 5-1 Finite Element Model of CEDM Penetration (42.5 Degrees)

Stress Analysis August 2003 Revision 0

5-4 I"-

Figure 5-2 Three Dimensional Model of the ICI Penetration Nozzle (54.8 Degrees)

Stress Analysis August 2003 Revision 0

5-5 Figure 5-3 Stress Distributions at Steady State Conditions: Outermost CEDM Penetration Nozzle (42.5 Degrees) (Hoop Stress is the Top Figure; Axial Stress is the Bottom Figure)

GOLF Stress Analysis August 2003 Revision 0

5-6 XId 25 2002 12.19~01 P~a.U NO. 2 TrAF-~40o4 1SYTS--'OLU DlMX -..42 tBvA -1.L Z1V -2 ZF -?06'70 PREC-.EW H4IDDEr DMX -. 434 SMN -50 SAX -73a04

-- 45207,

-- 1~O00a 20003 1~00090

?MI1l2'CH0(42.4Gd1,CYC SS3G/.1~.E0,J- Operatintg Figure 5-4 Axial Stress Distribution at Steady State Conditions for the Outermost CEDM (42.5 Degrees) Penetration, Along a Plane Oriented Parallel to, and Just Above, the Attachment Weld C' 0 5:'

Stress Analysis August 2003 Revision 0

5-7

/I--- Figure 5-5 Stress Distribution at Steady State Condition for the 37.1 Degrees CEDM Penetration Nozzle (Hoop Stress is the Top Figure; Axial Stress is the Bottom Figure)

Co(P Stress Analysis August 2003 Revision 0

5-8 Figure 5-6 Stress Distribution at Steady State Condition for the 29.1 Degrees CEDM Penetration Nozzle (Hoop Stress is the Top Figure; Axial Stress is the Bottom Figure)

Stress Analysis August 2003 Revision 0

5-9 Figure 5-7 Stress Distribution at Steady State Condition for the Center CEDM Penetration Nozzle (Hoop Stress is the Top Figure; Axial Stress is the Bottom Figure)

C~o%

Stress Analysis August 2003 Revision 0

5-10 Figure 5-8 Stress Distribution at Steady State Conditions for the ICI Penetration (Hoop Stress is the Top Figure; Axial Stress is the Bottom Figure)

Stress Analysis August 2003 Revision 0

5-1 1 Figure 5-9 Head Vent Nozzle Finite Element Model Stress Analysis August 2003 Revision 0

5-12 Figure 5-10 Stress Contours in the Head Vent Nozzle as a Result of Residual Stresses and Operating Pressure (Hoop stress is the Top Figure; Axial Stress is the Bottom Figure)

CIO Stress Analysis August 2003 Revision 0

6-1 6 FLAW EVALUATION CHARTS

6.1 INTRODUCTION

The flaw evaluation charts were developed from the stress analysis of each of the penetration locations, as discussed in Section 5. The crack gowth law developed for Millstone Unit 2 in Section 4.2 was used for each case, and several flaw tolerance charts were developed for each penetration location. The first series of charts characterizes the growth of a part through flaw, and the second series of charts characterizes the growth of a through-wall flaw in the length direction.

The allowable safe operating life of the penetration nozzle may then be directly determined, using the combined results of the two charts. All times resulting from these calculations are effective full power years, since crack growth will only occur at operating temperatures.

6.2 OVERALL APPROACH The results of the three-dimensional stress analysis of the penetration locations were used directly in the flaw tolerance evaluation.

The crack growth evaluation for the part-through flaws was based on the worst stress distribution through the penetration wall at the location of interest of the penetration. The highest stressed location was found to be in the immediate vicinity of the weld for both the center and outermost penetrations.

The stress profile was represented by a cubic polynomial:

a(x) = AO + Ax+A2 x 2 + A3 x 3 (6-1) where:

x = the coordinate distance into the nozzle wall a = stress perpendicular to the plane of the crack Al = coefficients of the cubic polynomial fit For the surface flaw with length six times its depth, the stress intensity factor expression of Raju and Newman [SA] was used. The stress intensity factor K, (4) can be calculated anywhere along the crack front. The point of maximum crack depth is represented by 4) = 0, and this location was also found to be the point of maximum Ke for the cases considered here. The following expression is used for calculating K, (4), where 4) is the angular location around the crack. The units of K, (0) are ksi-Tin .

05 3 K1(0) =[- ,G Gj(a /c, a /t, t /R, 4)) Aj aj (6-2)

The boundary correction factors Go (4)), GI (P), 02 (4)) and 03 (4) are obtained by the procedure outlined in reference [5A]. The dimension "a"' is the crack depth, aid "c" is the semi crack Flaw Evaluation Charts August 2003 Revision 0

6-2 length, while "t" is the wall thickness. "R" is the inside radius of the tube, and "Q" is the shape factor.

[

1A.C.0

6.3 RESULTS

AXIAL FLAWS CEDM and ICI Surface Flaws The results of the calculated growth through the wall thickness of the CEDM penetration nozzles for surface flaws are shown in Figures 62 through 67 for inside surface flaws. For outside surface flaws the results are shown in Figures 6-8 and 6-9. Based on the discussion in MRP-55 report [4H], the use of stress intensity factors less than 15 MPa4;; involves assumption not currently substantiated by actual CGR data for neither CEDM nor ICI nozzle materials.

Therefore, these crack growth curves begin at a flaw depth that results in a stress intensity factor of 15 MPa4m', which exceeds the threshold value of 9MPa4i;. This may result in curves with Flaw Evaluation Charts August 2003 Revision 0

6-3 different initial flaw sizes, as seen for example in Figure 6-3. Note that results are only provided for the uphill and downhill sides of each penetration nozzle; the stresses for the regions 90 degrees from these locations are compressive. If flaws are found in such a location, the results for either the uphill or downhill location, whichever is closer, can be used.

Each of these figures allows the future allowable service time to be estimated graphically, as discussed in Section 3. Results are shown for each of the penetration nozzles analyzed in each of these figures. The stresses are much higher near the attachment weld than at 0.5 inch below or above it, so separate figures have been provided for these three regions. For more than 0.5 inch below the weld, the crack growth will eventually come to rest since the stresses are compressive as shown for CEDM nozzles in Appendix E. Also, the stresses are different on the downhill side of the penetration as opposed to the uphill side, so these two cross sections have also been treated separately.

CEDM and ICI Through-Wall flaws The projected crack growth of a through-wall flaw in the CEDM and ICI penetration nozzles are the primary concern in evaluating the structural integrity of head penetrations. In some cases, the through-wall flaw may be located sufficiently below the attachment weld that additional time may be required for the flaw to grow to the attachment weld. To provide a means to evaluate the duration of this additional time, a series of flaw tolerance charts for through-wall flaws were prepared.

Charts were prepared for each of the penetrations evaluated, for both the uphill and downhill locations, as shown in Figures 6-12 through 6-20. In each figure, the upper extremity of the thru-weld crack is measured from the bottom of the nozzle. Note that in all the cases, the crack slows down significantly as it grows above the weld, due to the decreasing magnitude of the stress field.

This provides further assurance that axial flaws will not extend to a critical length which exceeds 15 inches, regardless of the duration of crack growth.

Head Vent The only flaw tolerance chart that is necessary for the head vent region is for flaws at and above the weld, since there is no portion of the head vent which projpcts below the weld. Figure 6-11 provides the projected growth of a part through flaw in the head vent just above the attachment weld The growth through the wall is relatively rapid, because the thickness of the head vent is small.

Flaw Evaluation Charts August 2003 Revision 0

6-4 6A CIRCUMFERENTIAL CRACK PROPAGATION Since circumferentially oriented flaws have been found at five plants (Bugey 3, Oconee 2, Crystal River 3, Davis Besse, and Oconee 3), it is important to consider the possibility of crack extension in the circumferential direction. The first case was discovered as part of the destructive examination of the tube with the most extensive circumferential cracking at Bugey 3. The crack was found to have extended to a depth of 2.25 mm in a wall thickness of 16 mm. The flaw was found at the outside surface of the penetration (number 54) at the downhill side location, just above the weld.

The circumferential flaws in Oconee Unit 3 were discovered during the process of repairing a number of axial flaws, whereas the circumferential flaw in Oconee Unit 2 and Crystal River Unit 3 were discovered by UT. Experience gained from these findings has enabled the development of UT procedures capable of detecting circumferential flaws reliably.

To investigate this issue completely, a series of crack growth calculations were carried out for a postulated surface circumferential flaw located just above the head penetration weld, in a plane parallel to the weld itself. This is the only flaw plane that could result in a complete separation of the penetration nozzle, since all others would result in propagation below the weld, and therefore there is no chance of complete separation because the remaining weld would hold the penetration nozzle in place.

[

Iac.e Flaw Evaluation Charts August 20U3 Revision 0

6-5 Sacs

]¢ The results of this calculation are shown in Figure 6-21. From Figure 6-21, it can be seen that the time required for propagation of a circumferential flaw to a point where the integrity of the CEDM penetration nozzle would be affected (330-350 degrees [11]) would be about 25 years. From the same figure, the required time for propagation of a circumferential flaw to a point where the integrity of the ICI penetration nozzls would be affected is about 38 years.

Due to the conservatism in the calculations (the time period for a surface flaw to become a through-wall flaw was conservatively ignored) the service life is likely to be even longer. In addition, due to uncertainties in the exact composition of the chemical environment in contact with the nozzle OD, a multiplicative factor of 2.0 is used in the CGR for all circumferential surface flaws on the OD of the head penetration nozzles located above the elevation of the J-groove weld.

6.5 FLAW ACCEPTANCE CRITERIA Now that the projected crack growth curves have been developed, the question remains as to what flaw size would be acceptable for further service.

Acceptance criteria have been developed for indications found during inspection of reactor vessel upper head penetration as part of an industry program coordinated by NEI (formerly NUMARC).

Such criteria are normally found in Section XI of the ASME Code, but Section XI does not require in-service inspection of these regions and therefore acceptance criteria are not available.

In developing the enclosed acceptance criteria, the approach used was very similar to that used by Section XI, in that an industry consensus was reached using input from both operating utility technical staff and each of the three PWR vendors. The criteria developed are applicable to all PWR plant designs.

Since the discovery of the leaks at Oconee and ANO-l, the acceptance criteria have been revised slightly, to cover flaws on the outside diameter of the penetration below the attachment weld, and flaws in the attachment weld. These revised criteria are now formally endorsed by the NRC [12].

The criteria presented herein are limits on flaw sizes, which are acceptable. The criteria are to be applied to inspection results. It should be noted that determination of the future service during which the criteria are satisfied is plant-specific and dependent on flaw geometry and loading conditions.

It has been previously demonstrated by each of the owners groups that the penetration nozzles are very tolerant of flaws and there is only a small likelihood of flaw extensions to larger sizes.

Therefore, it was concluded that complete fracture of the penetration nozzle is highly unlikely.

The approach used here is more conservative than that used in Section X applications where the Flaw Evaluation Charts August 2003 Revision 0

6-6 acceptable flaw size is calculated by placing a margin on the critical flaw size. For the current application, the critical flaw size would be far too large to allow a practical application of the approach used in Section XI applications, so protection against leakage is the priority.

The acceptance criteria presented herein apply to all the flaw types regardless of orientation and shape. Similar to the approach used in Section XI, flaws are first characterized according to established rules and then compared with acceptance criteria.

Flaw Characterization Flaws detected must be characterized by the flaw length and preferably flaw depth. The proximity rules of Section XI for considering flaws as separate, may be used directly (Section XI, Figure IWA 3400- 1). This figure is reproduced here as Figure 6-22.

When a flaw is detected, its projections in both the axial and circumferential directions must be determined. Note that the axial direction is always the same for each penetration, but the circumferential direction will be different depending on the angle of intersection of the penetration nozzle with the vessel head. The "circumferential" direction of interest here is along the top of the attachment weld, as illustrated in Figure 6.23. It is this angle which will change for each penetration nozzle and the top of the attachment weld is also the plane which could cause separation of the penetration nozzle from the vessel head. The location of the flaw relative to both the top and bottom of the partial penetration attachment weld must also be determined since a potential leak path exists when a flaw propagates through the penetration nozzle wall and up the penetration nozzle past the attachment weld. Schematic of a typical weld geometry is shown in Figure 6-24.

Flaw Acceptance Criteria The maximum allowable depth (a) for axial flaws on the inside surface of the penetration nozzle, at or above the weld is 75 percent of the penetration wall thickness. The term af is defined as the maximum size to which the detected flaw is calculated to grow in a specified time period. This 75 percent limitation was selected to be consistent with the maximum acceptable flaw depth in Section XI and to provide an additional margin against through wall penetration. There is no concern about separation of the penetration nozzle from the vessel head, unless the flaw is above the attachment weld and oriented circumferentially. Calculations have been completed to show that the geometry of all penetrations can support a continuous circumferential flaw with a depth of 75 percent of the wall thickness.

Axial inside surface flaws found below the weld we acceptable regardless of depth as long as their upper extremity does not reach the bottom of the weld during the period of service until the next inspection. Axial flaws that extend above the weld are limited to 75 percent of the wall thickness.

Axial flaws on the outside surface of the penetration nozzle below the attachment weld are acceptable regardless of depth, as long as they do not extend into the attachment weld during the Flaw Evaluation Charts August 2003 Revision 0

6-7 period of service until next inspection. Outside surface flaws above the attachment weld must be evaluated on a case by case basis, and must be discussed with the regulatory authority.

Circumferential flaws located below the weld are acceptable regardless of their depth, provided the length is less than 75 percent of the penetration nozzle circumference for the period of service until the next inspection. Circumferential flaws detected in this area have no structural significance except that loose parts must be avoided. To this end, intersecting axial and circumferential flaws shall be removed or repaired. Circumferential flaws at and above the weld must be discussed with the regulatory authority on a case by case basis.

Surface flaws located in the attachment welds themselves are not acceptable regardless of their depth. This is because the crack growth rate is several times faster than that of the Alloy 600 material, and also because depth sizing capability does not yet exist for indications in the attachment weld.

The flaw acceptance criteria are summarized in Table 6-1. Flaws that exceed these criteria must be repaired unless analytically justified for further service. These criteria have been reviewed and endorsed by the NRC, as documented in references [7, 8, & 12]. These criteria are identical with the draft acceptance criteria now being considered for Section XI, for head penetrations.

It is expected that the use of these criteria and crack growth curves will provide conservative predictions of the allowable service time.

Flaw Evaluation Charts August 2003 Revision 0

6-8 Table 6-1 Summary of R.V. Head Penetration Acceptance Criteria (Limits for Future Growth)

Axial Circumferential Location af I a( I Below Weld (ID) t no limit t .75 circ.

At and Above Weld (ID) 0.75 t no limit *

• Below Weld (OD) t no limit t .75 circ.

Above Weld (OD)

Note: Surface flaws of any size in the attachment weld are not acceptable.

• Requires case-by-case evaluation and discussion with regulatory authority.

af = Flaw Depth I = Flaw Length t = Wall Thickness Table 6-2 Millstone Unit 2 Penetration Geometries [10A, 10CI Penetration Type Wall Thickness (in.) Penetration OD (in.)

CEDM 0.566 3.850 ICI 0.469 5.563 Head Vent 0.154 1.050 Flaw Evaluation Charts August 2003 Revision 0

( C( .9 70 60 9 50 I

-tI Cr 40 co 4.F

,he

%ft-104 30 e

20 10 0

0 50 100 150 200 Crack Half Angle Figure 6-1 Stress Intensity Factor for a Through-Wall Circumferential Flaw in a Head Penetration Nozzle Flaw Evaluation Charts August 2003 Revision 0

) ) )

6-10 0.9 _-79 -ND ----

0.8 0.7 Ul s0.

X 0.6 ==--__ _ _ _ _W Q l_ Z'e N=

__ l 1 I ~~~~~~~~~~~~~~~A

_ 24_

gZZRe /d_ Z =-

E.Z 0

0.3 0.2 0.1 Millstone Unit 2 0

0 1 2 3 4 5 6 7 time (year)

Figure 6-2 Inside,Axial Surface Flaws, .5" Below the Attachment Weld, Nozzle Uphill Side- Crack Growth Predictions clI Flaw~~~~~~~~~~~~~~~~~~~~ at o ~sA t ~2~

~g~u~~~ .vl ~~~~~~~~~~~~~~ ~

~~

~0_hr Flaw Evaluation Charts August 2003

) Al I

0.9 = == == = = = _ = = == =--- == = ==/IHzJ._-z I I I I

_=== === =_== = == = = = = = == = = T= gz _ F11171 0.8 - - A.-

_~~~~~~K 111 _________7_eI-,.

0.7 0

UZ

=== == = ===== = = = = j 6 6 41e = =o= = = e  :=/ 1 1111111 X 0.6 ~~~~~~~~~~~~~~~~~~~3

_~ ~ ~ ~ _. ygr _c _ -lZ__/

0 3 0.5 = == = == = == = == s ==1Z== 71T IIIA CL 0

_~~~~~~~~~~~~~~~ UV _s 3 0.4 0.3

==love-~t Millston ni 0.2 0.1 0

0 1 2 3 4 5 6 7 8 9 10 time (year)

Figure 6-3 Inside, Axial Surface Flaws, .5" Bebw the Attachment Weld, Nozzle Downhill Side - Crack Growth Predictions CA?-

Flaw Evaluation Charts August 2003 Revision 0

) ) )

6-12 1

I ---- -_ - - -- _-- /- F -- -i 0.9

_ _ =_

_- _ =_ _= 175 3 _ \ E- E E lozzle nge _

0.8

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0.7

-_- - e- = = 4d54. _eg _ __

to CA 0

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_ ___ ___< /z X+ Y.a~-- 4_A 0

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0 1 2 3 4 5 6 time (year)

Figure 6-4 Inside, Axial Surface Flaws, at the Attachment Weld, Nozzle Uphill Side - Crack Growth Predictions C.rt Evlato Charts Flaw Evaluation August. 2003..

Flaw August 2003

) h3 1

0.9

_ _ _ _ _ _ = =-_ __ ~~~~~~~~~~7.1 le '- _F___

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_ = _ _ _ =~~~~~~~~~-5.8 dc(ICI)__ _= _ ___

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______ __ _ _ W W-- / e~~~~~~~~~~~

ang -

0 s 0.6 Q

0.

04-

__ = jZ _ t~~~~~~~~~~ 74 t > ~ ~ ~

_ __=_- _ -7 / -e $_ _ _ >s _ ____

e- _-~~~~~~~~2TT 0.2 0.1 5 gaF _ __ _ _ __ _ __ _ _ _ _ j ~~~~~~~~~Millstone Unit 2t 0

0 I 2 3 4 5 6 time (year)

Figure 6-5 Inside, Axial Surface Flaws, at the Attachment Weld, Nozzle Downhill Side - Crack Growth Predictions C1 4 -

Flaw Evaluation Charts August 2003 Revision 0

)

6-14 1

Nazzzzz Anz11e 0.9 INe Angle: l An

_z le:

0.8 =~~~~~~~~~~~~~~ca ___Xi

_X n____I__I /1/ 111 0.7 w

Q 0.6

0.5 a.

0 _ __ _ _e An

____lle: _ _ __ _E _ C zf X C _ _ f I I

¢ 0.4 0.3 0.2 0.1 Millstone Unit 2 0

0 1 2 3 4 5 6 7 8 9 time (year)

Figure 6-6 Inside, Axial Surface Flaws, .5" Above the Attachment Weld, Nozzle Uphill Side - Crack Growth Predictions Flaw Evaluation Charts August 2003

) ) )15 1

0.9 _ _= = ~~~_= -N4

-42: ZW deg- A _=

0.8 NC~~~~~~~~~~~ \g/e: -leiwozz.eI 0.7 _____ _____49.1 deg n

0 Q0.6 0.

0 g 0.5 CU 1 0.4 0.3 0.2 0.1 Millstone Unit 2 0

0 1 2 3 4 5 6 7 8 time (year)

Figure 6-7 Inside, Axial Surface Flaws, .5" Above the Attachment Weld, Nozzle Downhill Side- Crack Growth Predictions gut2 0 Fla Evaluation Flaw ChrsA Evlato Charts August 2003 Revision 0

) ) Jk 6-16 0 .9 0

0 .7 _ _ _ _

18 l l l l l l 548 deg(ICl Ode~~~~~~~~ll 4

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0.

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0 1 2 3 4 time (year)

Figure 6-8 Outside, Axial Surface Flaws, Below the Attachment Weld, Nozzle Uphill Side - Crack Growth Predictions C17 Flaw Evaluation Charts August 2003

) ~~~~~~~~~~~~~) ) 17 1 * ._ __ _ _ 11 / _I l 0.9 1 Nn771eAi ngleai t Z 0.8 = - = = de= t < e il;j

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0.3 _ _ _ _ __ __ __

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0.1 0_____ 1_____ _ _____ ______ ______ M illsto n e U n it 2 0

0 1 2 3 4 time (year)

Figure 6-9 Outside, Axial Surface Flaws, Below the Attachment Weld, Nozzle Downhill Side- Crack Growth Predictions FlawEvaluationCharts August 2003~~~~~~~~~~~~

Flaw Evaluation Charts August 2003 Revision 0

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0 - -=__=_l l 0 1 2 3 4 5 6 time (year)

Figure 6-10 Outside, Circumferential Surface Flaws, Along the Top of the Attachment Weld - Crack Growth Predictions (MRP Factor of 2.0 Included) cl9 A ug s 2003 Ev_.h at o Charts

_. a Evaluation Flaw art s August 2003

( ( ( 19 0.9 0.8 -

0.7 /

,0.6./

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0.3 0.2 0.1

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0 1 2 time (year)

Figure 6-11 Inside, Axial Surface Flaws, At the Attachment Weld, Head Vent, Nozzle Downhill Side- Crack Growth Predictions Flaw Evaluation Charts August 2003 Revision 0

( ( Q 6-20 4.5 o 4.0 E

0 0

o 3.5 z

E 0

01 1111111111111111111111111111111111111111 CED Ic

,C .5 053.0 a.

I.

U) 0.

CL2.5

--- Unit 2

~~~~~~~~~~~~~~~~~~~~~~~Millstone 2.0 0 1 2 3 4 5 6 7 8 9 10 11 12 perlod(EFPY)

Figure 6-12 Through-Wall Axial Flaws Located in the Center CEDM (0.0 Degrees) Penetration, Uphill and Downhill Side - Crack Growth Predictions August 2003 Evaluation Charts Flaw Evaluation August 2003

( ( 1 21 4.5 a 4.0 0

E0 km 0

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.65 03.

0.

D25 2.0 0 1 2 3 4 5 6 7 8 9 10 11 12 period(EFPY)

Figure 6-13 Through-Wall Axial Flaws Located in the 29.1 Degree Row of Penetrations, Downhill Side- Crack Growth Predictions Flaw Evaluation Charts August 2003 Revision 0

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0.

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4.5 12 0.

0.

4.0 4.5 4.0 0 1 2 3 4 5 6 7 8 9 10 11 perlod(EFPY)

Figure 6-14 Through-Wall Axial Flaws Located in the 29.1 Degree Row of Penetrations, Uphill Side- Crack Growth Predictions gut2 0 Fla Evaluation Flaw Ch_ sA Evlato Charts August 2003

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2.0 1.5 0 1 2 3 4 5 6 7 8 9 perlod( F Y Figure 6-15 Through-Wall Axial Flaws Located in the 37.1 Degree Row of Penetrations, Downhill Side- Crack Growth Predictions Flaw Evaluation Charts August 2003 Revision 0

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0 1 2 3 4 5 6 7 8 9 10 11 12 perlod(EF1PY)

Figure 6-16 Through-Wall Axial Flaws Located in the 37.1 Degree Row of Penetrations, Uphill Side- Crack Growth Predictions August 2003 Flaw Evaluation Charts Flaw August 2003

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o 4.0 0

m 2

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0 0

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a.

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Figure 6-17 Through-Wall Axial Flaws Located in the 42.5 Degree Row of Penetrations, Downhill Side- Crack Growth Predictions Flaw Evaluation Charts August 2003 Revision 0

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Figure 6-18 Through-Wall Axial Flaws Locatedin the 42.5 Degree Row of Penetrations, Uphill Side- Crack Growth Predictions Fla w Ev al u ati o n Ch a rts August 2003~~ ~~ ~~ ~

Flaw Evaluation Charts August 2003

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C 0

E 0

m 0

z E

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c C

0 a.

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0.

0 1 2 3 4 5 6 7 8 9 10 11 perlod(EFPY)

Figure 6-19 Through-Wall Axial Flaws Located in the 54.8 Degree Row of Penetrations, Downhill Side- Crack Growth Predictions Charts August 2003 Evaluation Charts Flaw Evaluation August 2003 Revision 0

( C (

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Figure 6-20 Through-Wall Axial Flaws Located in the 54.8 Degree Row of Penetrations, Uphill Side- Crack Growth Predictions August 2003 Charts Flaw Evaluation Charts August 2003

) )

180.000 150.000 o, 120.000 Im 0

C

'4-2 90.000 3

4-C

= 60.000 30.000

- -~~~~~~~~~~~~~Millstone Unit 2-0.000 0 5 10 15 20 25 30 35 40 period (year)

Figure 6-21 Through-Wall Circumferential Flaws Near the Top of the Attachment Weld for CEDM and ICI Nozzles- Crack Growth Predictions (MRP Factor of 2.0 Included)

CZ-O Auut20

_la Evlato Flaw Evaluation ht

. Charts August 2003 Revision 0

6-30 la) Single Linear Flaw (b) Single Curvilinear Flaw S4112in I 2- in. S S e2 £2>t 1-Q - R S l/112 in. r - Q -1 S 4 1/2 in.

_- I; -

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hel Overlappirig Parallbl Flows (f) Overlapping Flaws S C1/2 in.

lSC 1/2 in.

[4* 2 1 l -02 L t (g) Nonaligned Parallel Flaws Ihl Multiple Parallel Flaws R1 > 22 Figure 6-22 Section XI Flaw Proximity Rules for Surface Flaws (Figure TWA-3400-1) u ut2 0 Eato Charts Fl wEvaluation hat Flaw August 2003 Revision 0

6-31 I

Figure 6-23 Definition of "Circumferential" Flaw Evaluation Charts August 2003 Revision 0

6-32 NOZZLE TOP OF WELD I

BOTTOM OF WELD Figure 6-24 Schematic of Head Penetration Geometry Flaw Evaluation Charts August 2003 Revision 0

7-1 7

SUMMARY

AND EXAMPLE PROBLEMS An extensive evaluation has been carried out to characterize the loadings and stresses which exist in the head penetrations at Millstone Unit 2. Three-dimensional finite element models were constructed, and all pertinent loadings on the penetrations were analyzed [6A, 6B, 6C]. These loadings included internal pressure and thermal expansion effects typical of steady state operation. In addition, residual stresses due to the welding of the penetrations to the vessel head were considered.

Results of the analyses reported here are consistent with the axial orientation and location of flaws which have been found in service in a number of plants, in that the largest stress component is the hoop stress, and the maximum stresses were found to exist in the locations nearest and farthest away from the center of the vessel. The most important loading conditions were found to be those which exist on the penetration for the majority of the time, which are the steady state loading and the residual stresses.

These stresses are important because the cracking which has been observed to date in operating plants has been determined to result from primary water stress corrosion cracking (PWSCC).

These stresses were used in fracture calculations to predict the future growth of flaws postulated to exist in the head penetrations. A crack growth law was developed specifically for the operating temperature of the head at Millstone Unit 2, based on the EPRI recommendation, which is consistent with laboratory data as well as crack growth results for operating plants.

The crack growth predictions contained in Section 6 show that the future growth of cracks which might be found in the penetrations will be very slow, and that a number of effective full power years will be required for any significant extensions. An exception is the vent line, where the time is less because of the smaller thickness.

7.1 SAFETY ASSESSMENT It is appropriate to examine the safety consequences of an indication which might be found. The indication, even if it were to propagate through the penetration wall, would have only minor consequences, since the pressure boundary would not be broken, unless it were to propagate above the weld.

Further propagation of the indication would not change its orientation, since the hoop stresses in the penetration are much larger than the axial stresses. Therefore, it is extremely unlikely that the head penetration would be severed.

If the indication were to propagate to a position above the weld, a leak could result, but the magnitude of such a leak would be very small, because the crack could not open significantly due to the tight fit between the penetration and the vessel head. Such a leak would have no immediate impact on the structural integrity of the system, but could lead to minor wastage in the ferritic steel of the vessel head, as the borated primary water concentrates due to evaporation. Davis Besse has demonstrated the consequence of ignoring such leaks.

August 2003 Summary and Suxnrnaiy Example Problems and Example Problems August 2003 Revision 0

7-2 Any indication is unlikely to propagate very far up the penetration nozzle above the weld since the hoop stresses decrease in this direction, causing the indication to slow down and stop before it reaches the outside surface of the head.

The high likelihood that the indication will not propagate up the penetration nozzle beyond the vessel head ensures that no catastrophic failure of the head penetration will occur. The indication will be enveloped in the vessel head itself, which precludes the opening of the crack and limits leakage.

7.2 EXAMPLE PROBLEMS The crack growth prediction curves in Figures 6-2 through 6-21 can be used with the acceptance criteria of Section 6.5 to determine the available service time. In this section, a few examples will be presented to illustrate the use of these figures. The example cases are listed in Table 7-1.

Example 1. - Determine the service life of an axially oriented inside surface flaw whose upper extremity is located 1.3" below the weld on the uphill side of penetration no. 30, the penetration locality angle must first obtained from Table 1-1. In this case, the locality angle is 29.1 degrees and the initial flaw depth is 0.0425" (an.iJ and the initial flaw length is 0.085" (2cinkiia).

Assuming that the initial aspect ratio of 2:1 is maintained throughout the time that the inside surface flaw becomes a through-wall flaw, the final length of the flaw (2cfi,.,) will be 1.132". The upper extremity of the flaw is now located 0.7765" below the weld and validates the use of a single crack growth curve. The crack growth curve for the 29.1 degrees nozzle angle of Figure 6-2 is applicable and Figure 6-2 has been reproduced as Figure 7-1. The flaw is initially 7.5 percent of the wall thickness, and a straight line is drawn horizontally at a/t = 0.075 that intersects the crack growth curve. Using the acceptance criteria in Table 6.1, the service life can then be determined as the remaining time for this flaw to grow to the limit of 100 percent of the wail thickness or approximately 4.4 years (labeled as Service Life in Figure 7-1).

Example 2. In this case, the flaw is identical in size to that used in Example 1, but located on the outside surface and on the downhill side of penetration no. 50. This flaw, just as the flaw in Example 1, will not cross into the weld region. The applicable curve to use is Figure 6-9. The ratio a/t and initial reference time are likewise found using the same approach as used in Example

1. Using the acceptance criteria in Table 6.1, the determination of service life is illustrated in Figure 7-2, where we can see that the result is approximately 2.1 years.

Example 3. An axial inside surface flaw is located at the weld and on the uphill side of penetration no. 70. The initial length of the flaw is 0.250" and the initial depth is 0.05". From Table l-1, the angle of this penetration nozzle is 54.8 degrees. The applicable curve is Figure 6-4 and is reproduced here as Figure 7-3. In this case, the initial flaw depth is 10.7 percent of the wall thickness. The initial reference time can be found by drawing a horizontal line at a/t = 0.107.

Using the acceptance criteria in Table 6-1, the allowable service life can then be determined as the time for the flaw to reach a depth of 75 percent of the wall thickness. The final reference time is found through a horizontal line drawn at alt = 0.75. The service life can be determined through the intersection points of these lines and the crack growth curve. The resulting service life is approximately 2.85 years, as shown in Figure 7-3.

Summary and Example Problems August 2003 Revision 0

7-3 Example 4. In this case, we have postulated an axial inside surface flaw that will require the use of two flaw charts for its evaluation because the upper extremity of this flaw will reach within 0.5 inch below the attachment weld as it propagates into the nozzle wall. The flaw has an initial depth of 0.079 inch (2mm) and is located on the downhill side of penetration no. 1, which is the center nozzle at 0.0°. The length of the initial flaw is 0.395 inches and its upper extremity is located 1.0 inch below the attachment weld. Assuming that the initial aspect ratio of 5:1 is maintained as the flaw propagates into the nozzle wall, the final length of a through-wall flaw would be 2.83 inches long. The location of the upper extremity of this flaw would have reached within 0.5 inch below the weld as it propagates into the nozzle wall.

For this postulated flaw, the first step is to estimate the time required for the initial flaw to grow to within 0.5 inch of the weld. This can be accomplished with the use of Figure 62 or 6-3 (Figure 6.3 is used) because CGR is the same for both uphill and downhill side of the center penetration and is reproduced here as Figure 74A. The crack tip is I inch below the weld and is assumed to grow until the tip is 0.5 inches below. The final half-length of the flaw when it reaches 0.5 inches below the weld is the sum of the initial half-length and the 0.5 inches it has grown (0.5 +0.395/2). Then by dividing the aspect ratio and multiplying by two, the flaw depth when the crack tip is 0.5 inches below the weld is 0.279 in. The required tirne to reach such a flaw size (a/t=0.49) is estimated as 3.35 years in Figure 7-4A. Using the flaw depth calculated previously (49 percent) as the initial flaw depth, the curves in Figure 64 or 6.5 can be used, because CGR is the same for both uphill and downhill side of the center penetration, for inside surface flaws near the weld can be used to determine the service life before the flaw reaches the allowable flaw size. Using the acceptance criteria in Table 61 and Figure 65 as the source figure, Figure 7-4B shows an additional 0.75 years of service life for a total of 4.1 years. (See Guideline 5 below, for a quick and conservative approach to evaluate flaws that are more than 0.5"below the weld and will extend to within 0.5" of the weld).

Example 5. This case is an axial through-wall flaw whose upper-most end is 0.40 inches below the weld region on the uphill side of penetration 65, which is in the 42.5 degree row of penetrations, as seen in Table I-1. From Figure 618 we obtain the appropriate curve for the crack growth prediction, and this is reproduced as Figure 75. This figure gives a service life estimate of approximately 2.8 years to grow to the bottom of the weld as illustrated in Figure 7-5.

It is clear from these examples that the most important figures for use in evaluating flaws in head penetrations are the surface flaw Figures 6-2 through 6-9 for axial flaws and 6-10 for circumferential flaws. The figures which project the growth of through-wall flaws are valuable, but may be of limited practical use with the acceptance criteria. There is an important safety aspect to the through-wall flaw charts, however, in that they demonstrate that flaw propagation above the weld will be very limited.

Several guidelines are important to understand when using these charts.

1. If a flaw is found in a penetration nozzle for which no specific analysis was performed and there is a uniform trend as a function of penetration nozzle angle, interpolation between penetration nozzles is the best approach.

Summary and Example Problems August 2003 Revision 0

7-4

2. If a flaw is found in a penetration nozzle for which no specific analysis was performed and there is no apparent trend as a function of penetration nozzle angle, the result for the penetration nozzle with the closest angle should be used.

3. If a flaw is found which has a depth smaller than any depth shown for the penetration nozzle angle of interest, the initial flaw depth should be assumed to be the same as the smallest depth analyzed for that particular penetration nozzle.

4. The flaw evaluation charts are applicable for aspect ratio of 6 or less. Consult with Westinghouse if the as-found flaw has an aspect ratio larger than 6.0.

5. As shown in Example 4, flaws whose upper extremities grow within 0.5" below the weld require the use of both the 0.5" below the weld and "at the weld" flaw tolerance charts. To avoid the use of these two charts, the "at the weld" charts may solely be used in determining the service life.

This shall provide a conservative estimate of the crack growth due to a larger stress field.

6. Results are only provided for the uphill and downhill sides of the selected penetration nozzles. If flaws are found in locations between the uphill and downhill side, use the results for either the uphill or downhill location, whichever is closer K-,

Summary and Example Problems August 2003 Revision 0

7-5 Table 7-1 Example Problem Inputs: Initial Flaw Sizes and Locations Wall Vertical Circum. Penetration Length Depth Thick. Pen. Source No. Orientation Location Location Angle (2c) (a) a/t (t) No. Figure Axial - Inside 1.3"

_ SSurface Below Uphill 29.10 0.085" 0.0425" 0.075 0.566" 30 6-2 Weld Axial - 1.3" 2 Outside Below Downhill 37.10 0.085" 0.0425" 0.075 0.566" 50 6-9 Surface Weld 3 Axial - Inside At Weld Uphill ICI (54.8°) 0.250" 0.05" 0.107 0.469" 70 6-4 Surface 4 AxSurfacnside Below -- 0.00 0°395" 0.079" 0.14 0.566" 1 6-3, 6-5 SurfaegW ~Weld Axial~ 0.4" Axirug-all Below Uphill 42.5 -- - - 0.566" 65 6-18 Throgh-all Weld K>

Summary and Example Problems August 2003 Revision 0

) )

7-6 Acceptance Criteria (Table & 1) 1~~~~~~~~~~~~~~~~~~

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~~~~~~~~~~~~

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_- _ _-m __ __ _

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Figure 7-1 Example Problem 1 Problems Example and Summary August2003C Summary and Example Problems August 2003

) ) A7 Acceptance Criteria (Table 6-1) 1 0.9 I II L I I I~~~~~~~~-O711 I I 111 4 I/A 0.8 Locality Angles (Table 1-1) 0.7 T

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Figure 7-2 Example Problem 2 Summary and Example Problems August 2003 Revision 0

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7-8 Acceptance Criteria (Table & 1)

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< 1__

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0 2 3 4 5 6 time (year)

Figure 7-3 Example Problem 3 Auus 200 Summary Example Prblm and Exml Summary an Problems August 2003

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~~~~~=====_=_============st 0.8 =r _

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0 1 2 3 4 5 6 7 8 9 10 JC time (year)

Figure 7-4a Example Problem 4 (also see Figure 7-4b)

S y x P ms Summary and Example Problems August 2003 Revision 0

) ) )

7-10 Acceptance Criteria (Table 6-1) 1 I I I I i !Millstone Unit 2 0.9 -

l l l l I 1 L.~~I I 0.8 0.7 Locality Angles (Table 1-1)

/ I Nozzle U o. .6 No. T Angle s 0.2 5

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0

0.

0.

0. .2- __- __ - -< - - _-__- L if e

0. _ _e_ Ic Ife 0 1 2 3 4 5 6 time (year)

Figure 7-4b Example Problem 4 (also see Figure 7-4a)

CS25 Summary and Example Problems August 2003

r ( (

Acceptance Criteria (Table 6-1) 9.0 Qa Locality Angles (Table 1-1) 8.0 E

Nozzle No. Type IAngle 0 65 ICI 42.5 Z 7.0 E0

' 6.5 i:U 6.0 5.5 5.0 0 1 2 3 4 5 6 7 8 9 10 11 12 period(yrs)

Figure 7-5 Example Problem 5 Summary and Example Problems August 2003 Revision 0

8-a 8 REFERENCES

1. Scott, P. M., "An Analysis of Primary Water Stress Corrosion Cracking in PWR Steam Generators," in Proceedings, Specialists Meeting on Operating Experience With Steam Generators, Brussels Belgium, Sept. 1991, pages 5, 6.

2. Mcllree, A. R., Rebak, R B., Smialowska, S., "Relationship of Stress Intensity to Crack Growth Rate of Alloy 600 in Primary Water," Proceedings Internalional Symposium Fontevraud II, Vol, 1,

p. 258-267, September 10-14, 1990.

3. Cassagne, T., Gelpi, A., "Measurements of Crack Propagation Rates on Alloy 600 Tubes in PWR Primary Water, in Proceedings of the 5h International Symposium on Environmental Degradation of Material in Nuclear Power System-Water Reactor," August 25-29, 1991, Monterey, California.

4A. Crack Growth and MicrostructuralCharacterizationofAlloy 600 PWR Vessel Head Penetration Materials, EPRI, Palo Alto, CA. 1997. TR-109136.

4B. Vaillant, F. and C. Amzallag. "Crack Growth Rates of Alloy 600 in Primary Water," Presentation to the EPRI-MRP Crack Growth Rate (CGR) Review Team, Lake Tahoe, NV, August 10, 2001.

4C. Vaillant, F. and S. Le Hong. Crack Growth Rate Measurements in Primary Water ofPressure Vessel Penetrationsin Alloy 600 and Weld Metal 182, EDF, April 1997. HT-44/96/024/A.

4D. Framatome laboratory data provided by C. Amzallag (EDF) to MRP Crack Growth Rate Review Team, October 4, 2001 (Proprietary to EDF).

4E. Cassagne, T., D. Caron, J. Daret, and Y. Lefevre. "Stress Corrosion Crack Growth Rate Measurements in Alloys 600 and 182 in Primary Water Loops Under Constant Load," Ninth InternationalSymposium on Environmental Degradationof Materialsin Nuclear Power Systems-Water Reactors (Newport Beach, CA, August 1-5, 1999), Edited by F. P. Ford, S. M. Bruemner, and G. S. Was, The Minerals, Metals & Materials Society (TMS), Warrendale, PA, 1999.

4F. Studsvik laboratory data provided by Anders Jenssen (Studsvik) to MRP Crack Growth Rate Review Team, October 3, 2001 (Proprietary to Studsvik).

4Q1 4H "Materials Reliability Program (MRP) Crack Growth Rates for Evaluating Primary Water Stress Corrosion Cracking (PWSCC) of Thick Wall Alloy 600 Material (MRP-55) Revision 1," EPRI, Palo Alto, CA:, November 2002. 1006695.

4I. "Crack Growth Rate Tests of Alloy 600 in Primary PWR Conditions," Communication from Mvt L. Castafto (CIEMAT) to J. Hickling (EPRI), March 25,2002.

4J. Gomez-BriceZo, D., J. LapeZa, and F. Bkzquez "Crack Growth Rates in Vessel Head Penetration Materials," Proceedingsof the InternationalSymposium FontevraudIII: Contribution of Materials Investigation to the Resolution of Problems Encountered in Pressurized Water Reactors (Chinton, France, September 12-16, 1994), French Nuclear Energy Society, Paris, 1994, pp. 209-214.

References August 2003 Revision 0

8-2 4K. Gomez-BriceZo, D. and J. LapeZa. "Crack Growth Rates in Vessel Head Penetration Materials,"

Proceedings: 1994 EPRI Workshop on PWSCC of Alloy 600 in PWRs (Tampa, FL, November 15-17, 1994), EPRI, Palo Alto, CA, TR-105406, August 1995, pp. E4-1 through E4-15.

4L. Gomez-BriceZo, D., et al. "Crack Propagation in Inconel 600 Vessel Head Penetrations,"

Eurocorr 96, Nice, France, September 24-26, 1996.

4M. CastaZo, M. L., D. G.-.mez-BriceZo, M. Alvarez-de-Lara, F. Bkzquez, M. S. Garcia, F. Hern<ndez, and A. Largares. "Effect of Cationic Resin Intrusions on IGA/SCC of Alloy 600 Under Primary Water Conditions," Proceedings ofthe InternationalSymposium FontevraudIV:

Contributionof MaterialsInvestigation to the Resolution of Problems Encountered in Pressurized Water Reactors (France, September 14-18, 1998), French Nuclear Energy Society, Paris, 1998, Volume 2, pp. 925-937.

5A. Newman, J. C. and Raju, I. S., "Stress Intensity Factor Influence Coefficients for Internal and External Surface Cracks in Cylindrical Vessels," in Aspects of Fracture Mechanics in Pressure Vessels and Pirina, PVP Vol. 58, ASME, 1982, pp. 37-48.

5B. Hiser, Allen, "Determinists and Probabilistic Assessments," presentation at NRC/Industry/ACRS meeting, November 8, 2001.

6A. [ ]ae 6B.

6C. Fleming, M.A., "Palo Verde I and 2 CEDM and Head Vent Stress Analysis", Dominion Engineering Calculation No. C-7736-00-6, February 2002.

7. USNRC Letter, W. T. Russell to W. Raisin, NUMARC, "Safety Evaluation for Potential Reactor Vessel Head Adapter Tube Cracking," November 19, 1993.

8. USNRC Letter, A. G. Hansen to R. E. Link, "Acceptance Criteria for Control Rod Drive Mechanism Penetrations at Point Beach Nuclear Plant, Unit 1," March 9, 1994.

9. Millstone Unit 2, NRC Bulletin 2001-01 Response (Docket Numbers 50-336 and 50-423, B18466) August 31, 2001.

10A. CE Drawing No. E-233-537, " Closure Head Penetrations", Revision 5, October 1969.

10B. CE Drawing No. E-233-534, " Closure Head Forming & Welding", Revision 8, October 1969.

10C. CE Drawing No. E-233-536, " Closure Head Nozzle Details", Revision 8, December 1969.

11. CEOG Report # CEN-614, "Safety Evaluation of the Potential for and Consequences of Reactor Vessel Head Penetration Alloy 600 OD Initiated Nozzle Cracking:' December 1993.

12. USNRC Letter, R Barrett to A. Marion, "Flaw Evaluation Guidelines," April 11, 2003.

References August 2003 Revision 0

A-i APPENDIX A ALLOWABLE AREAS OF LACK OF FUSION: WELD FUSION ZONES There are two fusion zones of interest for the head penetration nozzle attachment welds, the penetration itself (Alloy 600) and the reactor vessel head material (A533B ferritic steel). The operating temperature of the upper head region of the Millstone Unit 2 is 312'C (5940 F), so the materials will be very ductile.

The toughness of both materials is quite high, so any flaw propagation along either of the fusion zones will be totally ductile.

Two generic calculations were completed for the fusion zones, one for the critical flaw size, and the second for the allowable flaw size, which includes the margins required in the ASME code. The simpler case is the Alloy 600 fusion zone, where the potentia l failure will be a pure shearing of the penetration as the pressurized penetration tube is forced outward from the vessel head, as shown in Figure A-1.

The failure criterion will be that the average shear stress along the fusion line exceeds the limit shear stress. For the critical flaw size, the limiting shear stress is the shear flow stress, which is equal to half the tensile flow stress, according to the Tresca criterion. The tensile flow stress is the average of the yield stress and ultimate tensile stress of the material. The criterion for Alloy 600 at 318'C (604'F) is:

Average shear stress < shear flow stress = 26.85 ksi This value was taken from the ASME Code,Section III, Appendix I, at 6000 F.

For each penetration, the axial force which produces this shear stress results from the internal pressure.

Since each penetration has the same outer diameter, the axial force is the same. The average shear stress increases as the load carrying area decreases (the area of lack of fusion increases). When this increasing lack of fusion area increases the stress to the point at which it equals the flow stress, failure occurs. This point may be termed the critical flaw size. This criterion is actually somewhat conservative.

Alternatively, use of the Von Mises failure criterion would have set the shear flow stress equal to 60 percent of the axial flow stress, and would therefore have resulted in larger critical flaw sizes.

The allowable flaw size, as opposed to the critical flaw size discussed above, was calculated using the allowable limit of Section III of the ASME Code, paragraph NB 3227.2. The criterion for allowable shear stress then becomes:

Average shear stress < 0.6 Sm = 13.98 ksi where S. = the ASME Code limiting design stress from Section III, Appendix I.

The above approach was used to calculate the allowable flaw size and critical flaw size for the outermost and center penetrations. The results show that a very large area of lack of fusion can be tolerated by the head penetrations, regardless of their orientation. These results can be illustrated for the outermost CEDM penetration.

Appendix A August 2003 Revision 0

A-2 The total surface contact area for the fusion zone on the outermost head penetration is 17.4 in2 . The calculations above result in a required area to avoid failure of only 1.45 in 2 , and using the ASME Code criteria, the area required is 2.79 in 2 . These calculations show that as much as 83.9 percent of the weld may be unfused, and the code acceptance criteria can still be met.

To envision the extent of lack of fusion which is allowable, Figure A-2 was prepared. In this figure, the weld fusion region for the outermost penetration has been shown in an unwrapped, or developed view.

The figure shows the extent of lack of fusion which is allowed, in terms of limiting lengths for a range of circumferential lack of fusion. This figure shows that the allowable vertical length of lack of fusion for a full circumferential unfused region is 84 percent of the weld length. Conversely, for a region of lack of fusion which extends the full vertical length of the weld, the circumferential extent is limited to 302 degrees. The extent of lack of fusion which would cause failure is labelled "critical" on this figure, and is even larger. The dimensions shown on this figure are based on an assumed rectangular area of lack of fusion.

The full extent of this allowable lack of fusion is shown in Figure A-3, where the axes have been expanded to show the full extent of the tube-weld fusion line. This figure shows that a very large area of lack of fusion is allowable for the outer most penetration. Similar results were found for the center penetration, where the weld fusion area is somewhat smaller at 16.1 in2 .

A similar calculation was also carried out for the fusion zone between the weld and the head, and the result is shown in Figure A-4. The allowable area of unfused weld for this location is 84.8 percent of the total area. This approach to the fusion zone with the carbon steel head is only approximate, but may provide a realistic estimate of the allowable. Note that even a complete lack of fusion in this region would not result in rod ejection, because the weld to the tube would prevent the tube from moving up through the vessel head.

The allowable lack of fusion for the weld fusion zone to the head may be somewhat in doubt, because of the different geometry, where one cannot ensure that the failure would be due to pure shear. To investigate this concern, additional finite element models were constructed with various degrees of lack of fusion discretely modeled, ranging from 30 to 65 percent. The stress intensities around the circumference of the penetration were calculated, to provide for the effects of all stresses, as opposed to the shear stress only, as used above. When the average stress intensity reaches the flow stress (53.7 ksi), failure is expected to occur. The code allowable stress intensity is 1.5 Sm, or 35 ksi, using the lower of the Alloy 600 and ferritic allowables at 316'C (6000 F).

The results of this series of analyses are shown in Figure A-5, where it is clear that large areas of lack of fusion are allowable. As the area of lack of fusion increases, the stresses redistribute themselves, and the stress intensity does not increase in proportion to the area lost. These results seem to confirm that the shear stress is the only important stress governing the critical flaw size for the head fusion zone as well.

Appendix A August 2003 Revision 0

A-3 Siide L 1 LI

~~nbili

~~~ 1 Ai\ Weld Cladding (CP))

Head Penetration Nozzle Figure A- Typical Head Penetration Appendix A August 2003 Revision 0

A-4 100 95 L.0 41) 0~ Critical L- 90 0) cm)

-. 1 85 Allowable 80 300 310 320 330 340 350 360 Circumferential Extent (Degrees)

Figure A-2 Allowable Regions of Lack of Fusion for the Outermost Penetration Tube to Weld Fusion Zone: Detailed View Appendix A August 2003 Revision 0

A-5 Alowa~L Critical 100 90 80 70 C-)

a) 60 1-c 50 a- 40 30 20 10 0

0 100 200 300 Circumferential Extent (Degrees)

Figure A-3 Allowable Regions of Lack of Fusion for the Outermost Penetration Tube to Weld Fusion Zone Appendix A August 2003 Revision 0

A-6 Agmoable Cnticai 100

_ __ ___ _ _ _ _ _ _ _ _ _ _ _ _ _ _ ~~~~~~~~~~~~~~~~~~

90 80 70 U1) 60 50 CL 40 0)

-J 30 20 10 0

0 100 200 300 Circumferential Extent (Degrees)

Figure A-4 Allowable Regions of Lack of Fusion for all Penetrations: Weld to Vessel Fusion Zone ut2 0 Ape dx AAu August 2003 Appendix A Revision 0

• 7 ( ( -7 60 Critical 50 40 Code Allowable II 30 20 10 0 . I ~~~~i

-- i I I I _ _

0 5 10 15 20 25 30 35 40 45 50 Percentage Area Reduction Figure A-5 Allowable Regions of Lack of Fusion for the Weld to Vessel Fusion Zone Appendix A August 2003 Revision 0

B-I Appendix B CEDM AND ICI NOZZLE HOOP STRESS VS DISTANCE FROM BOTTOM OF WELD PLOTS Appendix B

C. C B-2 Figure B-I Hoop Stress Distribution Below the Weld Downhill and Uphill Side (00 CEDM Penetration Nozzle) 70,000 60,000 50,000 40,000

,In in 30,000 C,, 20,000 0

0 10,000 0

-10,000

-20,000 CD.0 0.5 1.0) 1.5i I.~~~~~J j~~~

2.0 2.5 3.0 Distance from Bottom of Weld (In)

-- *Inside -- Outsidel II*

Appendix B August 2003

C ( ( B-3 Figure B-2 Hoop Stress Distribution Below the Weld Downhill Side (29.10 CEDM Penetration Nozzle) 90,000 80,000 70,000 60,000 50,000 0.

a)

In 40,000 E

co c.

30,000 8

20,000 10,000 0

-10,000

.. I

<~~~~~~~~~~---

-20,000 o.o 0.5 1.0 1.5 2.0 2.5 3.0 Distance from Bottom of Weld (in)

I *Inside -U- Outside I Appendix B August ----

2U03 Revision 0

( C B-4 Figure B-3 Hoop Stress Distribution Below the Weld Uphill Side (29.10 CEDM Penetration Nozzle) 70,000 60,000 50,000 40,000 0.1

\ X

0) 30,000 Co 20,000 10,000 0 AS~~~~~~~~~~~~~~~~~~~~- Id_

-10,000

-20,000 ' _. . . . . ..... ~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~

0.0 1.0 2.0 3.0 4.0 5.0 6.0 Distance from Bottom of Weld (in)

I-$- Inside -- Outside I August 2003 Appendix B Appendix B August 2003

(A (( B5 Figure B-4 Hoop Stress Distribution Below the Weld Downhill Side (37.10 CEDM Penetration Nozzle) 90,000 80,000 70,000 60,000 50,000 -

0.

140,000 230,000 20,000 10,000 0

-10,000 0.0 0.5 1.0 1.5 2.0 2.5 3.0 Distance from Bottom of Weld (In)

-+ Inside -- Outside Appendix B August 2003 Revision 0

(4 (C (

B-6 Figure B-5 Hoop Stress Distribution Below the Weld Uphill Side (37.10 CEDM Penetration Nozzle) 60,000 I

50,000 40,000 30,000 t I 1 ¶ 0.

CL w 20,000 I. - I. t I I a

i.*

Co 0-0 0

o 10,000 N t N11 1* t I I

---

a 0I

-10,000 .

-20,000 i i i

-30,000 - - -

11 - - - - -t -. - -

0.iD 1.0 2.0) 3.0 I4.0 5.0 6.0 Distance from Bottom of Weld (in) l--Inside - Outside August 2003 Appendix B August 2003

C

-t a'

Figure B-6 Hoop Stress Distribution Below the Weld Downhill Side (42.50 CEDM Penetration Nozzle) 100,000 90,000 80,000 70,000 60,000

0. 50,000 06 (to 40,000 4..

to, 30,000 20,000 10,000 0

-10,000

-20,000

-30,000 0.0 0.5 1.0 1.5 2.0 2.5 3.0 Distance from Bottom of Weld (in) 14Inside -'- Outside Appendix B August 2003 Revision 0

(-( X B-8 Figure B.7 Hoop Stress Distribution Below the Weld Uphill Side (42.50 CEDM Penetration Nozzle) 50,000 6a0,000 40,000 - > i

-20,000 Co 0.

0

-10,000 10.000 - v

-20,000 -

0.0 1.0 2.0 3.0 4.0 5.0 6.0 7.0 Distance from Bottom of Weld (In)

-l--Inside -*-Outside I August 2003 B

Appendix B August 2003

(

C B-9 Figure B-B Hoop Stress Distribution Below the Weld Downhill Side (54.80 ICI Penetration Nozzle) 80,000 Y T 1 1~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~

70,000 ~ ~ ~ 4 1'Rx v~ .l4 l l 4l 60,000 50,000

t. 40,000 CL 0a 30,000 0

• S 20,000 M:

CD 10,000 W-0

-10,000

-20,000 l l l l.l.

-30,000 . _ - - - 4-

- - -- 4 - I - - - 9-- - - - -

0.0 0.2 0.4 0.6 0.8 1.0 1.2 Distance from Bottom of Weld (in) 1--Inside - Outside Appendix B August 2003 Revision 0

(B-j C, Figure B-9 Hoop Stress Distribution Below the Weld Uphill Side (54.80 ICI Penetration Nozzle) 60,000 50,000 40,000 30,000 i, 20,000 0.

O 10,000 P

C% 0 0L 0

0 X -10,000

-20,000

-30,000

-40,000

-50,000 0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 2.0 Distance from Bottom of Weld (in)

-+- Inside -- Outside]

August 2003 Appendix B August 2003