ML021350067

Non-Proprietary Version of Topical Report WCAP-15753, Rev. 1, Structural Integrity Evaluation of Reactor Vessel Upper Head Penetrations to Support Continued Operation: D.C. Cook, Units 1 and 2

ML021350067
Person / Time
Site:	Cook
Issue date:	05/10/2002
From:	Greenlee S Indiana Michigan Power Co
To:	Document Control Desk, Office of Nuclear Reactor Regulation
References
AEP:NRC:2055-01, TAC MB3551, TAC MB3552 WCAP-15753, Rev. 1
Download: ML021350067 (92)

Text

Indiana Michigan Power Company Cock Nuclear Plant 5W0 Crce Drve Bud-e

, M1 49107 611465-&%l INDIANA MICHIGAN POWER May 10, 2002 AEP:NRC:2055-01 10 CFR 50.55a Docket Nos.

50-315 50-316 U. S. Nuclear Regulatory Commission ATTN: Document Control Desk Mail Stop O-P 1-17 Washington, DC 20555-0001 Donald C. Cook Nuclear Plant Units 1 and 2 PROPOSED ALTERNATIVES TO THE REQUIREMENTS OF SECTION XI OF THE AMERICAN SOCIETY OF MECHANICAL ENGINEERS CODE - TRANSMITTAL OF NON-PROPRIETARY REPORT (TAC Nos. MB3551 AND MB3552)

Reference:

Letter from S. A. Greenlee, Indiana Michigan Power Company, to Nuclear Regulatory Commission Document Control Desk, "Donald C. Cook Nuclear Plant Units 1 and 2, Proposed Alternatives to the Requirements of Section XI of the American Society of Mechanical Engineers Code - Request for Additional Information (TAC Nos. MB3551 and MB3552)," submittal AEP:NRC:2055, dated April 25, 2002.

In the referenced letter, Indiana Michigan Power Company (I&M), the licensee for Donald C. Cook Nuclear Plant Units 1 and 2, provided three copies of a proprietary report, WCAP-14118, Revision 5, "Structural Integrity Evaluation of Reactor Vessel Upper Head Penetrations to Support Continued Operation:

D. C. Cook Units 1 and 2," together with an affidavit from Westinghouse Electric Company, LLC (Westinghouse) that supported the withholding of the information in the report from public disclosure. At the time of the submittal, a non-proprietary version of the report was unavailable, and I&M committed to providing a non-proprietary version by May 31, 2002. The attachment to this letter contains three copies of non-proprietary report, WCAP-15753, Revision 1, "Structural Integrity Evaluation of Reactor Vessel Upper Head Penetrations to Support Continued Operation: D. C. Cook Units 1 and 2," together with a copy of the Westinghouse Proprietary Information Notice and Copyright Notice.

AIl': AI merica t" De/guy PWartner

U. S. Nuclear Regulatory Commission AEP:NRC:2055-01 Page 2 This letter contains no new commitments.

Should you have any questions, please contact Mr. Gordon P. Arent, Manager of Regulatory Affairs at (616) 697-5553.

Sincerely, A. Greenlee Director, Nuclear Technical Services RV/bjb Attachment c:

K. D. Curry, w/o attachments J. E. Dyer MDEQ - DW & RPD, w/o attachments NRC Resident Inspector R. Whale, w/o attachments

ATTACHMENT TO AEP:NRC:2055-01 WCAP -15753, Revision 1, "Structural Integrity Evaluation of Reactor Vessel Upper Head Penetrations to Support Continued Operation: D. C Cook Units 1 and 2" (Non-proprietary)

U. S. Nuclear Regulatory Commission PROPRIETARY INFORMATION NOTICE Transmitted herewith are proprietary and/or non-proprietary versions of documents furnished to the NRC in connection with requests for generic and/or plant-specific review and approval.

In order to conform to the requirements of 10 CFR 2.790 of the Commission's regulations concerning the protection of proprietary information so submitted to the NRC, the information which is proprietary in the proprietary versions is contained within brackets, and where the proprietary information has been deleted in the non-proprietary versions, only the brackets remain (the information that was contained within the brackets in the proprietary versions having been deleted). The justification for claiming the information so designated as proprietary is indicated in both versions by means of lower case letters (a) through (f) contained within parentheses located as a superscript immediately following the brackets enclosing each item of information being identified as proprietary or in the margin opposite such information. These lower case letters refer to the types of information Westinghouse customarily holds in confidence identified in Sections (4)(ii)(a) through (4)(ii)(f) of the affidavit accompanying this transmittal pursuant to 10 CFR 2.790(b)(1).

AEP:NRC:2055-01

U. S. Nuclear Regulatory Commission COPYRIGHT NOTICE The reports transmitted herewith each bear a Westinghouse copyright notice.

The NRC is permitted to make the number of copies of the information contained in these reports which are necessary for its internal use in connection with generic and plant-specific reviews and approvals as well as the issuance, denial, amendment, transfer, renewal, modification, suspension, revocation, or violation of a license, permit, order, or regulation subject to the requirements of 10 CFR 2.790 regarding restrictions on public disclosure to the extent such information has been identified as proprietary by Westinghouse, copyright protection notwithstanding. With respect to the non-proprietary versions of these reports, the NRC is permitted to make the number of copies beyond those necessary for its internal use which are necessary in order to have one copy available for public viewing in the appropriate docket files in the public document room in Washington, DC and in local public document rooms as may be required by NRC regulations if the number of copies submitted is insufficient for this purpose. Copies made by the NRC must include the copyright notice in all instances and the proprietary notice if the original was identified as proprietary.

AEP:NRC:2055-01

Westinghouse Non-Proprietary Class 3 Structural Integrity Evaluation of Reactor Vessel Upper Head Penetrations to Support Continued Operation:

D. C. Cook Units 1 and 2 Westinghouse Electric Company LLC

WESTINGHOUSE NON-PROPRIETARY CLASS 3 WCAP-15753 Revision 1 Structural Integrity Evaluation of Reactor Vessel Upper Head Penetrations to Support Continued Operation:

D. C. Cook Units 1 and 2 W. H. Bamford K. R. Hsu April 2002 Reviewer:

D. Tang J+/-*nTh Approved:.

-Z" S. A. S/wamy, Manage" Structural Mechanics Technology Westinghouse Electric Company LLC P.O. Box 355 Pittsburgh, PA 15230-0355

©2002 Westinghouse Electric Company LLC All Rights Reserved 5967.doc-041202

iii TABLE OF CONTENTS I IN T R O D U C T IO N.............................................................................................................................

1-1 2

HISTORY OF CRACKING IN HEAD PENETRATIONS...............................................................

2-1 3

OVERALL TECHNICAL APPROACH...........................................................................................

3-1 3.1 PENETRATION STRESS ANALYSIS.........................................................................................

3-1 3.2 FLAW TOLERANCE APPROACH..............................................................................................

3-1 4

MATERIAL PROPERTIES, FABRICATION HISTORY AND CRACK GROWTH P R E D IC T IO N...................................................................................................................................

4 -1 4.1 MATERIALS AND FABRICATION............................................................................................

4-1 4.2 CRACK GROWTH PREDICTION..............................................................................................

4-1 5

ST R E SS A N A LY SIS.........................................................................................................................

5-1 5.1 OBJECTIVES OF THE ANALYSIS............................................................................................

5-1 5.2 M O D E L.........................................................................................................................................

5 -1 5.3 STRESS ANALYSIS RESULTS - OUTERMOST PENETRATION..........................................

5-1 5.4 STRESS ANALYSIS RESULTS-NEXT OUTERMOST PENETRATION..................................

5-2 5.5 STRESS ANALYSIS RESULTS-CENTER PENETRATION......................................................

5-2 5.6 STRESS ANALYSIS RESULTS: HEAD VENT..........................................................................

5-2 6

FLAW EVALUATION CHARTS......................................................................................................

6-1 6.1 IN T R O D U C T IO N.........................................................................................................................

6-1 6.2 O V ER A LL A PPR O A C H...............................................................................................................

6-1

6.3 RESULTS

AXIAL FLAWS.........................................................................................................

6-2 6.4 CIRCUMFERENTIAL CRACK PROPAGATION.......................................................................

6-3 6.5 FLAW ACCEPTANCE CRITERIA..............................................................................................

6-5 6.6 EXAMPLE CALCULATIONS.....................................................................................................

6-7 April 2002 5967.doc-041202 Revision 1

iv TABLE OF CONTENTS (cont'd) 7 SUM M ARY AND CONCLUSIONS.................................................................................................

7-1 8

REFERENCES..................................................................................................................................

8-1 APPEND IX A...........................................................................................................................................

A-1 April 2002 5967.doc-041202 Revision 1

1-1 1

INTRODUCTION In September of 1991, a leak was discovered in the reactor vessel control rod drive head penetration region of an operating plant. This has led to the question of whether such a case could occur at D. C. Cook Unit 1 or 2. The geometry of interest is shown in Figure 1-1.

The leak resulted from cracking which occurred in the outermost penetrations of a number of operating plants, as discussed in Section 2. This outermost location, as well as the center penetration, was chosen for fracture mechanics analyses to support continued safe operation of D. C. Cook Unit 1 or 2 if such cracking were to be found.

The basis of the analyses was a detailed three dimensional elastic-plastic finite element analysis of the two penetration locations, as described in detail in Section 5. The geometry of the hillside penetration analyzed is shown in Figure 1-2.

The fracture analyses were carried out using reference crack growth rates developed from the literature and from service experience. The results are presented in the form of flaw evaluation charts for both surface and through wall flaws, to determine the allowable time of safe operation if indications are found.

All the times calculated in this handbook are effective full power years.

Revision 1. This revision was prepared to provide the results obtained by using the Scott model [4A] for the crack growth predictions, and to add a detailed defense of that model. This revision is consistent with the proprietary report WCAP-14118, Rev. 5.

Introduction 5967.doc-042202 April 2002 Revision I

1-2 LOCATION OF AXIAL CRACKS PARTIAL PENETRATION MEL?

CRDM THERMAL SLEEVE Figure 1-1 Reactor Vessel Head Adapter Penetration Tube, Showing Locations of Axial Cracks Found in Some Plants Introduction 5967.doc-041202 April 2002 Revision I I

1-3 Radius 64.5 in.

from vessel centerline i*'-

Radius 59.8 in.

from vesse centerline Figure 1-2 Geometry of the Hillside Penetrations Analyzed Introduction 5967.doc-041202 April 2002 Revision 1

2-1 2

HISTORY OF CRACKING IN HEAD PENETRATIONS In September of 199 1, leakage was reported from the reactor vessel head penetration region of a French plant, Bugey Unit 3. Bugey 3 is a 920 megawatt three-loop PWR 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°F (90°C). The leak was detected by metal microphones located on the top and bottom heads, and 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. Further 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 be cracked, 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 I was cracked. The locations of all the cracked penetrations are shown in Figure 2-1. None of the 17 penetrations inspected at Paluel 3 showed indications of cracking, at the time, but further inspection of the French plants have confirmed at least one crack in each operating plant.

Thus far, the cracking in tubes not manufactured by Babcock and Wilcox Tubular Products 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.

History of Cracking in Head Penetrations April 2002 5967-doc-041202 Revision 1

2-2 Non-destructive examinations of the leaking CRDM nozzles showed that most of the cracks originated on the outside surface of the nozzles below the J-groove weld, were axially oriented, and propagated primarily in the nozzle base material to an elevation above the top of the J-groove weld where 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.

a,c.e a.c,e The cracking has now been confirmed to be primary water stress corrosion cracking. Relatively high residual stresses are produced in the outermost 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 the plants examined thus far are summarized in Table 2-1.

History of Cracking in Head Penetrations 5967 doc-041202 April 2002 Revision 1

2-3 Table 2-1 Operational Information and Inspection Results for Units Examined (Results to December 30, 2001)

Head Penetrations Plant Units Temp.

Total Penetrations With Country Type Inspected K Hours

('F)

Penetrations Inspected Indications France CPO 6

80-107 596-599 390 390 23 CPY 28 42-97 552 1820 1820 126 1300MW 20 32-51 558-597 1542 1542 95 Sweden 3 Loop 3

75-115 580-606 195 190 8

Switzerland 2 Loop 2

148-154 575 72 72 2

Japan 2 Loop 7

105-108 590-599 276 243 0

3 Loop 7

99 610 455 398 0

4 Loop 3

46 590 229 193 0

Belgium 2 Loop 2

115 588 98 98 0

3 Loop 5

60-120 554-603 337 337 6

Spain 3 Loop 5

65-70 610 325 102 0

Brazil 2 Loop 1

25 NA 40 40 0

South Africa 3 Loop I

NA NA 65 65 6

Slovenia 2 Loop I

NA NA 49 49 0

South Korea 2 Loop 3

NA NA 49 49 3

3 Loop 2

NA NA 130 130 2

US 2 Loop 2

170 590 98 98 0

3 Loop 1

NA NA 65 20 0

4 Loop 14 NA NA 899 287 35 TOTALS 113 7134 6123 306 History of Cracking in Head Penetrations 5967.doc-041202 April 2002 Revision 1

2-4 FRENCH R/V CLOSURE HEAD PENETRATION CRACKING EdF PLANTS - PENETRATIONS WITH CRACKING 270' 270" 0*

90"

• Cracked Pne'trotion BUGEY 3 90" 0 Cracked Ptr-~ot~on BUGEY 4 270" 080" QO" 0.4 0

0.0 0.30 0~

0.

0..

0.1 OZ 0.~ ?01 OZ@

0 C-4, 049 02 04 75 047 9*"

0 Cracked Peqntraton FESSENHEIM I

Figure 2-1 History of Cracking in Head Penetrations 5967.doc-041202 01 April 2002 Revision I

3-1 3

OVERALL TECHNICAL APPROACH The primary goal of this work is to provide technical justification for the continued safe operation of D. C. Cook Units One and Two in the event that cracking is discovered during inservice inspections of the Alloy 600 reactor vessel head penetrations.

3.1 PENETRATION STRESS ANALYSIS Three dimensional elastic-plastic finite element stress analyses have been performed to determine the stresses in the head penetration region [6]. These analyses have considered the pressure and thermal transient loads associated with steady state operation, as well as the residual stresses which are produced by the fabrication process.

a,c-e 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 inservice inspection, its size can be compared with the flaw size which is considered allowable for continued service. This "allowable" flaw size is determined from the actual loadings (including mechanical, residual, and transient loads) on the head penetration for the plant of interest. Suitable margins to ensure the integrity of the reactor vessel as well as safety from unacceptable leakage rates, should also be considered. 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.

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

Overall Technical Approach 5967.doc-041202 April 2002 Revision 1

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

These two types of charts can be used to provide estimates of the time which remains 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 (tA) (eg a/t=.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 = tB Overall Technical Approach 5967.doc-041202 April 2002 Revision 1

3-3 Flaw Becomes Through - Wall

=.75 Detected Indication Allowable Time (tA ) Bef Allowable Time Before Wall Penetration, t p Time ( Months )

Figure 3-1 Schematic of a Head Penetration Flaw Growth Chart for Part Through Flaws Overall Technical Approach 5967.doc-041202 1.0 LL April 2002 Revision 1

Critical Length ( Excessive Leakage )

Time ( Months )

Figure 3-2 Schematic of a Head Penetration Flaw Tolerance Chart for Through-Wall Flaws Overall Technical Approach 5967.doc-041202 April 2002 Revision I 3-4 E

LL e-*

4-1 4

MATERIAL PROPERTIES, FABRICATION HISTORY AND CRACK GROWTH PREDICTION 4.1 MATERIALS AND FABRICATION The head adapters for D. C. Cook Units I and 2 were manufactured by Westinghouse from material produced by Huntington Alloys in the USA. The carbon content, mechanical properties and heat treatment of the Alloy 600 material used to fabricate the D. C. Cook vessels are provided in Tables 4-1 and 4-2. The material CMTRs were used to obtain the 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 1700 - 1800'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 D. C. Cook Units 1 and 2 vessels relative to a sample of the French head adapters which have experienced cracking. The general trend for the head adapter penetrations in the D. C. Cook vessels 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. 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. Upon study of his paper there were several ambiguities, and several phone conversations were held to clarify his conclusions. These discussions led to Scott's admission that reference 1 contains an error, in that no correction for cold work was applied to the McIllree/Smialowska data. The correct development is below.

Material Properties, Fabrication History and Crack Growth Prediction 5967.doc-041202 April 2002 Revision I

4-2 An equation was fitted to the data of reference [2] for the results obtained in water chemistries that fell in within the standard specification. Results for chemistries outside the specification were not used. The following equation was fitted to the data:

da.2.8x 10-(K _9)116 m/sec dt where K is in MPafm.

The next step described by Scott in his paper was to correct these results for the effects of cold work.

Based on work by Cassagne and Gelpi [3], he concluded that dividing the above equation by a factor of 10 would be appropriate to account for the effects of cold work. This step was inadvertently omitted from Scott's paper. even though it is discussed. The crack growth law for 330'C then becomes:

da-2.8 x 10-12 (K -9)116 m/sec dt This equation was verified by Scott in a phone call in July 1992.

Scott further corrected this law for the effects of temperature, but his correction was not used in the model employed here. Instead, an independent temperature correction was developed based on service experience, as will be discussed below.

The applicability of the Scott model to the head penetrations at the D. C. Cook Units I and 2 was recently confirmed by two independent approaches. The first was a collection of all available data from Huntington Alloys materials tested over the past ten years [4B]. The results are shown in Figure 4-3, along with the Scott model for the test temperature. It can clearly be seen that the Scott model is nearly an upper bound for the various Huntington heats studied.

A second independent set of data were used to validate the model, and these data were obtained form the two inspections carried out a penetration 75 of D. C. Cook Unit 2, which was first found to be cracked in 1994 [4C]. 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 4-3, and are consistent with the laboratory data.

Since both D. C. Cook Units operate at temperatures lower than 330'C in the head region, 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 4-4. 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].

Study of the data shown in Figure 4-4 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 Material Properties, Fabrication History and Crack Growth Prediction April 2002 5967.doc-041202 Revision I

4-3 accepted activation energy of 44-50 Kcallmole used to characterize the effect of temperature on crack initiation, but the trend of the actual data for many different sources is unmistakable.

lace Therefore the following growth rate models were used for the D. C. Cook head penetrations:

da =8.xl-13(K 9)116 M/sec (Unit 1) dt da =1. 4 8 x1 0 - 12 (K 9)1.16 m/sec (Unit2) dt where K = applied stress intensity factor, in MIPa-mi.

This equation implies a threshold for cracking susceptibility, Kiscc = 9 MPa'-mm.

Material Properties, Fabrication History and Crack Growth Prediction 5967.doc-041202 April 2002 Revision 1

4-4 Table 4-1 D. C. Cook Unit 1 R/V Head Adapter Material Information HT.No.

C Mn Fe S

Si Cu Ni Cr Co YS UTS Mtl.

Vendor Heat Treatment (ksi)

(ksi)

Spec NX-7926 Ladle 0.07 0.37 7.51 0.009 0.3 0.16 76.18 15.38 0,05 35.5 94.5 SB-167 Huntington 1725F 1.5 hr Air Cooled Check 0.072 0.37 7.47 0.009 0.38 0.15 74.83 15.7 0.03 NX-7280 Ladle 0.07 0.13 8.19 0.007 0.2 0.11 76.32 14.95 0.05 40.5 98.5 SB-167 Huntington 1725F 1.5 hr Air Cooled Check 0.08 0.14 8.26 0.006 0.26 0.11 75.15 15.1 0.03 NX-8069 Ladle 0.06 0.25 8.1 0.007 0.29 0.18 76.1 14.99 0.08 58.5 98 SB-167 Huntington 1725F 1.5 hr Air Cooled Check 0.061 0.25 8.21 0.004 0.32 0.15 74.14 14.9 0.08 NX-8251 Ladle 0.06 0.3 7.69 0.007 0.28 0.16 76.16 15.32 0.05 35 94.5 SB-167 Huntington 1725F 1.5 hr Air Cooled Check 0.056 0.29 7.73 0.007 0.3 0.15 74.89 15.2 0.04 NX-7760 Ladle 0.06 0.16 8.2 0.007 0.3 0.15 74.83 16.27 0.06 38 97.5 SB-167 Huntington 1725F 1.5 hr Air Cooled Check 0.062 0.18 8.01 0.003 0.33 0.14 74.86 16.32 0.05 Note: Chemistries are in wt. %.

Material Properties, Fabrication History and Crack Growth Prediction 5967.doc-041202 April 2002 Revision I

4-5 Table 4-2 D. C. Cook Unit 2 RNV Head Adapter Material Data HT.No.

C Mn Fe S

Si Cu Ni Cr Co YS UTS Mtl.

Vendor Heat Treatment (ksi)

(ksi)

Spec NX-0215 Ladle 0.07 0.22 8.64 0,007 0.25 0,22 75.21 15.36 0.07 51.0 103.0 SB-166 Westinghouse 1700 or 1800F I hr/Water Check 0.07 0.21 8.47 0.002 0.32 0.24 74.66 15.17 0.06 Quenched NX-0216 Ladle 0.09 0.24 8.53 0.007 0.21 0.22 75.11 15.57 0.06 57.0 107.0 SB-166 Westinghouse 1700 or 1800F I hr/Water Check 0.07 0.23 8.36 0.002 0.28 0.22 74.53 15.35 0.05 Quenched NX-0218 Ladle 0.08 0.28 9.02 0.008 0.19 0.28 74,31 15.81 0.06 51,0 102.0 SB-166 Westinghouse 1700 or 1800F I hr/Water Check 0.09 0.27 8.65 0.003 0.27 0.29 74.06 15.60 0.05 Quenched NX-0219 Ladle 0.06 0.24 8.76 0.007 0.18 0.22 75.0 15.51 0.07 41.0 100.0 SB-166 Westinghouse 1700 or 1800F I hr/Water Check 0.05 0.23 8.56 0.003 0.23 0.23 74.50 15,25 0.06 Quenched NX-0223 Ladle 0.07 0.31 8.5 0.007 0.29 0.17 75.29 15.34 0,07 63.0 104.0 SB-166 Westinghouse 1700 or 1800F

! hr/Water Check 0.08 0.31 8.37 0.002 0.36 0.19 74.31 15.33 0.06 Quenched NX-0230 Ladle 0.06 0.18 8.69 0,007 0.18 0.17 75,5 15.09 0.04 58/56 101.0/

SB-166 Westinghouse 1700 or 1800F 100.0 1 hr/Water Check 0.03 0.19 8,54 0.002 0.23 0.19 74.81 15.04 0.04 Quenched NX-0233 Ladle 0.06 0.18 7.93 0.007 0.26 0,14 76.17 15.23 0.05 58/44 101.0/

SB-166 Westinghouse 1700 or 1800F 100.0 1 hr/Water Check 0.04 0.18 7.85 0.003 0.30 0.16 75.44 15.09 0.04 Quenched Note: Chemistries are in wt. %.

Material Properties, Fabrication History and Crack Growth Prediction 5967.doc-041202 April 2002 Revision 1

D.C. Cook (12 Heats) 1 E~dF Vessels (11 Heats).

60 55 50 45 40 35 30 25 20 15 10 5

0 (gz

Yield Strength (ksi)

Figure 4-1 Yield Strength of the Various Heats of Alloy 600 Used in Fabricating the D. C. Cook Units 1 and 2 and French Head Adapter Penetrations Material Properties, Fabrication History and Crack Growth Prediction 5967.doc-041202 4-6 0

C')

a1) 0

-4 a,

C11 C-I cg

April 2002 Revision I cp

/,ý ý4

4-7 60 55 D.C, Cook (12 Heats)

SEdF Vessels (11 Heats))

5 0 - - - - - -.

.-, 4 5 - - - - - -.

1 -1 4 0 -- - - --

S 5.....................

Vs) 3 0..

0 5

-- 2 0 -.--

(~35 S-............

5 10 0

0" Carbon Content (Weight %)

Figure 4-2 Carbon Content of the Various Heats of Alloy 600 Used in Fabricating the D. C. Cook Units I and 2 and French Head Adapter Penetrations Material Properties, Fabrication History and Crack Growth Prediction April 2002 5967.doc-041202 Revision 1

4-8 1.E-09

-1.E-10 E

_-SCOT MODEL da/dt=2.23e-12(K-9)A1"16

-z 2

+ Test Temp.(302.6, 340.2, 0

.*-F 353.3, 343.0 corrected to o

"3250)

A COOK 2 0 1.E-11

+

+I 1.E-12 0

10 20 30 40 50 60 70 80 Stress Intensity Factor, K (MPa*sqrt(m))

Figure 4-3 Model for SCC Growth Rates in Alloy 600 in Primary Water Environments (325'C),

With Supporting Data from Huntington Materials Material Properties, Fabrication History and Crack Growth Prediction 5967.dcc-041202 April 2002 Revision 1

4-9 TEMPERATURE, DEG. C 372 352 333

000.0 100.0-315 298 282 0.00180 0.00155 0.00160 0.00165 0.00170 0.00175 RECIPROCAL TEMPERATURE, 1/DEG. K Key:

0 CRDM Field Data A CRDM Lab Tests All others are S/G Tube Lab & Field Datal Figure 4-4 Summary of Temperature Effects on SCC Growth Rates for Alloy 600 in Primary Water, Laboratory and Field Experience Material Properties, Fabrication History and Crack Growth Prediction 5967.doc-041202 April 2002 Revision 1 X

z 0

,.I)

-j IM 0

IC 0

5-1 5

STRESS ANALYSIS 5.1 OBJECTIVES OF THE ANALYSIS The objective of this analysis was to obtain accurate stresses in each CRDM housing and its immediate vicinity. To do so requires a three dimensional analysis which considers all the pertinent loadings on the penetration [6]. An investigation of deformations at the lower end of the housing was also performed using the same model. Three locations were considered: the outermost row, the next outermost row, and the center location.

The analyses were used to provide information for the flaw tolerance evaluation which follows 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 cracking which has been observed. The geometry of D.C. Cook Units 1 and 2 in the head penetration and head regions is identical, so one stress analysis covers both units.

5.2 MODEL A three dimensional finite element model comprised of isoparametric brick and wedge elements with midside nodes on each face was used to obtain the stresses and deflections. A view of the unstressed model is shown in Figure 5-1. Taking advantage of symmetry through the vessel and penetration centerlines only half of the penetration geometry plus the surrounding vessel were modeled for the outermost and next outermost penetrations. In the center penetration case, it was necessary to model only one-quarter of the penetration as opposed to one-half of the penetration. The difference between the hillside penetrations and the center penetration was that there was no differential height across the weld for the center penetration.

In the models, the lower portion of the Control Rod Drive Mechanism (CRDM) Adapter tube (i.e., penetration tube), the adjacent section of the vessel closure head, and the joining weld were modeled. The vessel to penetration tube weld was simulated with two layers of elements. The penetration tube, weld metal and cladding were modeled as Alloy 600 and the vessel head shell as carbon steel.

5.3 STRESS ANALYSIS RESULTS - OUTERMOST PENETRATION Figure 5-2 shows the outward displacement of the entire model for the steady state condition. For the steady state, the tube OD is pressing on the vessel (i.e. couple each tube node, except for the vertical direction, to its neighbor in the vessel). Figure 5-3 presents the hoop stresses for the steady state condition.

Stress Analysis 5967.doc-041202 April 2002 Revision 1

5-2 a]ac~e 5.4 STRESS ANALYSIS RESULTS-NEXT OUTERMOST PENETRATION a.c e 5.5 STRESS ANALYSIS RESULTS-CENTER PENETRATION I

ace 5.6 STRESS ANALYSIS RESULTS: HEAD VENT The head vent is a smaller penetration than the CRDM head penetrations, 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 7.8 inches from the centerline of the head dome, and its dimensions are shown in Figure 5-7.

The head vent was evaluated using a three dimensional finite element model, as shown in Figure 5-8.

1 Stress Analysis 5967 doc-041202 April 2002 Revision 1

5-3

]a,c,e

]ac~e Stress Analysis 5967.doc-041202 April 2002 Revision I I

S© c

N.

• o 0

L*0

5-5 0.006 Inch No.78 STEADY STATE( ); CYC"L ANNULUS OPENS

'WRP o

7 M.

Figure 5-2 Steady State Displacement of R/V Closure Head and Outermost Penetration Stress Analysis 5967.doc-041202 April 2002 Revision I

.'\\

\\

Displ. Masif. = 20.

Load ste 12 tEýraUio = 5 Time Variable = 22 Max. = 191186.

Mi.. = -176799.

1500D0.

120000.

90000.

60000.

30000.

0.

HOOP 3 16:15:50 WOG78

.en920731 Dispf. Magif. - 20.

L*ad ste a=12Lteration

= 5 Time= I Neutral ile =NRfW78D5 RES Variable = S33 Max. = 164860.

Mi.. = -189078.

150000a I120(K0.

900C0.

60000, 30000, 0.

316:172:I3 W0078 Vs,920721 Figure 5-3 Stress Distribution at Steady State Conditions: Outermost Penetration Stress Analysis April 2002 Revision I 56

5-7

'2 Noytr; 

- NRrfl

Figure 5-4 Stress Distribution at Steady State for the Outermost Penetration, Along a Plane Oriented Parallel to, and Just Above, the Attachmlent Weld Stress Analysrs April 2002 5967 doe-041 202 Revision I C-c() t-

5-8 flispi. Magnit = 20.

Load st

= 12 [terano" = 12 Time j0.13 eNira

• ie = W 507SK1 RES Variable = 822 Ma,. =

1343 min..

-129698, 90(X)0.

-250000.

HOOP Dis'l, M~g.,EL = 20.

Load

ýnt e -

12 I tenratco i -

12 T im e=0.

Neu nal bil t,

= W 6507SK IE MS Variable = $333 M.

= 273656.

Miý = -148810 15D000.

900DI0.

600W0.

0.

-250000.

5WOO AXIAL Figure 5-5 Stress Distribution at Silli y

% ttI Con dkinotl for th Next O uter nmost Penetration Stress Analysis 5967doc 041202 April 2002 Revision I 0-03

DaL LL Ma Variable = 322 Max. = 100976.

Mitt. = 636.86 00 4 50M, 300M

-25600, LTE; CYCL; KINE HARD; P (j

l

-]*fOuf S II Moe, Dec 7 1992 21:33:23 WOGO Ve9S20731

>feutrai tile = WOOC Voriabre = S33 Mný.

77877.5 Mm.

46381.

60 S45000.

150W0.

0.

"Do000 Asi Ls~ea (j) IGW

]

Or Dec 7 1992 21:36:56 WOGO V1:920731 Figure 5-6 Stress Distribution at Steady State Condition for the Center Penetration Stress Analysis 5967 dc, 041202 59 HWoo stress HOOP AXIAL April 2002 Revision I Coh-

5-10 Figure 5-7 Vent Pipe Dimensions (Inches) [9]

Stress Analysis 5967.doc-041202 April 2002 Revision I

5-11 ANSYS 5-4 DEC 28 1998 10:44:53 PLOT NO.

4 NODAL SOLUTION STEP=3 SUB =1 TIME=3 SINT (AVG)

DMX =.231174 SMN =2028 SMNB=566.623 SMX =88876 SMXB=93298 El Figure 5-8 Vent Pipe Finite Element Model Stress Analysis 5967.doc-041202 11677 21327 30977 40627 50277 59926 69576 79226 88876 April 2002 Revision I I

5-12 ANSYS 5.4 DEC 28 1998 11:07:51 PLOT NO.

I NODAL SOLUTION STEP=I SUB -I TIME= 1 SY (AVG)

RSYS=4 DMX =.523E-03 SMN =-119.681 SMNB=-2697 SMX -7758 SMDXB=7865 Figure 5-9 Hoop Stress in the Head Vent As A Result Of Design Pressure of 2500 psi Stress Analysis 5967.doc-041202

-119.68 755.622 1631 2506 3382 4257 5132 6007 6883 7758

-119.68 755. 622 1631 2506 5132 6007 6883 7758 mm~

FIZi April 2002 Revision 1

5-13 8.466--

• I

-1dVntPp R.V. Closure Head 2) 7.0(UNIT 1)/ 6.5(UNIT 2)

R = 86(UNIT 1)/ 83.69(UNIT 2)

R = 0.5(UNIT 1)/ 0.313(UNIT 2)

Weld Buttering 30' Figure 5-10 Stress Intensity Results in the Head Vent for the Governing Upset Condition Stress Analysis 5967.doc-041202 April 2002 Revision 1

-Head Vent Pipe

Penetration Tube cut5 Weld Buttering Figure 5-11 Various Cuts Taken for Analysis of the Head Vent Note: Cut 1 Was the Governing Location Stress Analysis 5967 doc-041202 April 2002 Revision I 5-14 cutI Penetratl!

Weld

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 growth law developed for D. C. Cook in Section 4.2 was used for each case, and two flaw tolerance charts were developed for each penetration location. The first chart characterizes the growth of a part through flaw, and the second chart characterizes the growth of a through-wall flaw in the length direction. The allowable remaining life of the penetration may then be directly determined, using the combined results of the two charts. All times resulting from these calculations are effective full power years.

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 maximum stress is the hoop stress, and the flaws which have been found inservice are all longitudinally oriented, so the hoop stress component was used.

The crack growth evaluation for the part-through flaws was based on the stress distribution through the penetration wall at the location which corresponds to the highest stress along the inner surface 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:

t/\\2

/XN3 cF(x) = A0 + A1 -- + A2,-

+ A3 -I t

It) yt where x is the coordinate distance into the wall t

=

wall thickness o

=

stress perpendicular to the plane of the crack Ai coefficients of the cubic fit For the surface flaw with a length six times its depth, the stress intensity factor expression of McGowan and Raymund [5A] was used. The stress intensity factor K, ((p) can be calculated anywhere along the crack front. The point of maximum crack depth is represented by y = 0. The following expression is used for calculating K, (f0), where (p is the angular location around the crack.

Sp20.5(

a 2

1/4(

2a I

a2 4a3

)

Qi) 2 COS

+-s A

0 t

2 A

2 3ntA 3H The magnification factors Ho((p), H1(p), H2(0p) and H3(q) are obtained by the procedure outlined in reference [5A]. The parameter C is the flaw half-length.

Flaw Evaluation Charts April 2002 5967.doc-041202 Revision I

6-2

] a C

6.3 RESULTS

AXIAL FLAWS CRDM Surface Flaws The results of the calculated growth through the wall for inside surface axial flaws postulated in the penetrations are summarized in Figures 6-1 a and 6-1b for Unit 1, and Figures 6-2a and 6-2b for Unit 2.

Figures 6-1 and 6-2 apply to surface crack locations anywhere in the weld region of any of the penetrations, since the stress results were taken at the highest stressed location, which is in the outermost penetration. The "a" figure in each case is a prediction of crack growth at and below the attachment weld region, while the "b" figure covers crack growth above the weld. Figures 6-1c and 6-2c apply to crack growth for outside surface axial flaws, regardless of location, for the two units. Note that the predicted extension through the penetration thickness requires many years at the operating temperature for either D. C. Cook Unit 1 or 2, regardless of the location.

Head Vent The only flaw evaluation chart necessary for the head vent region is for flaws at and above the weld, since there is no portion of the head vent which projects below the weld. Figure 6-1d and 6-2d provide the projected growth of a part through flaw in the head vent just above the attachment weld (cut 1 in Figure 5 11). The growth through the wall is relatively rapid, because the thickness of the head vent is small.

CRDM Through-Wall Flaws Figures 6-3 (a and b) and 6-4 (a and b) present the predicted crack growth for a through-wall flaw postulated to exist below the weld region in the outermost row of penetrations. These results are for the Flaw Evaluation Charts 5967.doc-041202 April 2002 Revision I

6-3 lower hillside and centerside locations respectively. Note that separate figures are provided for crack growth vs. time for the two different Units. The growth for Unit 1 is slower, because it operates at a lower temperature. Although there are different levels of ovality (and therefore residual stress) in the different penetrations, it is clear that in the vicinity of the weld and below it, the total stresses approach the yield stress of the material, which was set at 378.6 MPa (55 ksi) for this calculation. Figures 6-5 (a and b) and 6-6 (a and b) provide similar results for the next outermost row of penetrations.

Figures 6-7 (a and b) provide projections of growth above the weld region for the center penetration.

Note that for some of the penetrations crack extension actually stops, as the stress intensity factor decreases with the lower stresses, to a value below the threshold cracking susceptibility value of 9 MPa-Vm.

6.4 CIRCUMFERENTIAL CRACK PROPAGATION Since circumferentially oriented flaws have been found at four plants (Bugey3, Oconee 2, Crystal River, 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 longitudinal cracking at Bugey 3, and the crack was found to have extended to a depth of 2.25 mm in a wall thickness of 16 ram. The flaw was found at the outside surface of the penetration (number

54) at the lower hillside location, just above the weld.

The circumferential flaws in Oconee Unit 3 were discovered during the process of repairing a number of axial flaws, while the circumferential flaw in Oconee Unit 2 and Crystal River were discovered by UT.

Experience gained from these findings has enabled the development of UT procedures capable of detecting circumferential flaws reliably.

It is important to realize that a flaw would have to propagate through the penetration or the attachment weld, and result in a leak, before the outer surface of the penetration would be exposed to the water.

Cracking could then begin for an outside surface flaw. (This is believed to have been the case at all three plants in which circumferential flaws were found). This time period was conservatively ignored in the calculations to be discussed.

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 which could result in a complete separation of the penetration, since all others would result in propagation below the weld, and therefore no chance of complete separation because the remaining weld would hold the penetration in place.

axce Flaw Evaluation Charts April 2002 5967.doc-04 1202 Revision I

6-4 a.c,e Sa.c.e Flaw Evaluation Charts April 2002 5967.doc-041202 Revision 1

6-5 Therefore we see that the time required for propagation of a circumferential flaw to a point where the integrity of the penetration would be affected would be at least 38 years. Because of the conservatisms in the calculations, as discussed above, it is likely to be even longer.

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

Acceptance criteria have been developed for indications found during inspection of reactor vessel upper head penetrations. These criteria were developed as part of an industry program coordinated by NUMARC (now NEI). Such criteria are normally found in Section XI of the ASME Code, but Section XI does not require inservice 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-1, 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 in draft form, but they are expected to be acceptable to the NRC, and will be used in these evaluations. The draft portions of the acceptance criteria will be noted below.

The criteria which are 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 penetrations are very tolerant of flaws and there is only a small likelihood of flaw extension to large sizes. Therefore, it was concluded that complete fracture of the penetration is highly unlikely and, therefore, protection against leakage during service is the priority.

The approach used here is more conservative than that used in Section XI applications where the acceptable flaw size is calculated by putting a margin on the critical flaw size. In this case, the critical flaw size is far too large to allow a practical application of this approach so protection against leakage is the key element.

The acceptance criteria apply to all flaw types regardless of orientation and shape. The same approach is used by Section XI, where flaws are characterized according to established rules and then compared with acceptance criteria.

Flaw Characterization Flaws detected must be characterized by length and preferably 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 6-10.

Flaw Evaluation Charts April 2002 5967.doc-041202 Revision 1

6-6 When a flaw is found, 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 with the head. The "circumferential" direction of interest here is along the top of the attachment weld, as illustrated in Figure 6-11. It is this angle which will change for each penetration and which is also the plane which could cause separation of the penetration tube from the head. The location of the flaw relative to both the top and bottom of the partial penetration attachment weld must be determined since a potential leak path exists when a flaw progresses through the wall and up the penetration past this weld. A schematic of a typical weld geometry is shown in Figure 6-12.

Flaw Acceptance Criteria The maximum allowable depth (af) for flaws on the inside surface of the penetration, at or above the weld is 75 percent of the penetration wall thickness regardless of the flaw orientation. 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 head penetration from the head, unless the flaw is above the attachment weld and oriented circumferentially. Calculations have been completed to show that all penetration geometries can support a continuous circumferential flaw with a depth of 75 percent of the wall.

Axial inside surface flaws found below the weld are 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 which extend above the weld are limited to 75 percent of the wall.

Axial flaws on the OD of the penetration below the attachment weld are acceptable regardless of depth, as long as they do not extend into the attachment weld during the period of service until next inspection.

Axial OD 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 circumference for the period of service until the next inspection. Flaws in this area have no structural significance but 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.

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

These criteria are summarized in Table 6-1. Flaws which exceed these criteria must be repaired unless analytically justified for further service. These criteria have been reviewed and approved by the NRC, as documented in references 7 and 8, with the exception of the draft criteria discussed above, for OD flaws and flaws in the attachment weld.

Flaw Evaluation Charts April 2002 5967.doc-041202 Revision 1

6-7 It is expected that the use of these criteria and crack growth curves will provide conservative predictions of the allowable time of service. Similar criteria have been proposed in Sweden and France, and are under discussion in other countries.

6.6 EXAMPLE CALCULATIONS The crack growth prediction curves in Figures 6-1 through 6-9 can be used with the acceptance criteria of Section 6.5 to determine the available service time for either unit. In this section, a few examples will be presented to illustrate the use of these figures. Although this handbook allows calculations to be done for either unit, the examples presented here have used Unit 2. The example cases are listed in Table 6-2.

Example 1. For an axially oriented surface flaw, the crack growth curves of Figure 6-2 are appropriate.

Since the flaw is located below the weld, Figure 6-2a is appropriate, and has been reproduced as Figure 6-13. Figures 6-2a and 6-2b here both use the same crack growth curve, but illustrate two different scenarios. Figure 6-2a shows the result if the flaw is close to the weld, or is projected to grow to the bottom of the weld during service. In this case the flaw initial depth is 25 percent of the wall thickness, so project a line horizontally at alt = 0.25, intersecting the crack growth curve. The service life is then determined as the time for this flaw to grow to the limit of 75 percent of the wall thickness, or approximately 5.5 years (labelled Service Life 1 in Figure 6-13).

The other case, also illustrated in figure 6-2a, is that the flaw remains below the bottom of the weld. In this case, the criteria allow the flaw to extend through the wall, which results in a longer service life, approximately 7.0 years (labelled Service Life 2 in Figure 6-13). If the flaw were sufficiently far below the weld, we could take advantage of the additional time for a through wall flaw to grow up to the weld.

This case will be illustrated in example 6.

Example 2. In this case the flaw is identical in size to example 1, but located at the weld, and at a location 180' away from the flaw in example 1. The curve to use is in Figure 6-2a. The circumferential location is not important for surface flaws, only for through-wall flaws. The determination of service life is illustrated in Figure 6-14, where we see the result is approximately 5.5 years.

Example 3. The flaw is at the weld, and twice as deep as the flaw considered in example 2. It is oriented at 0'. The curve from Figure 6-2a is again used to determine the service life. The flaw depth is 50 percent of the wall thickness, so project horizontally at this value to intersect the crack growth curve.

The allowable service life is then determined as the time for the flaw to reach a depth of 75 percent of the wall. As shown in Figure 6-15, this time is approximately 1.8 years.

Example 4. This case is for a circumferential flaw which has been discovered above the weld. The appropriate figure for this type flaw is Figure 6-8, which has been reproduced as Figure 6-16, where the flaw size has been plotted. The additional service life is obtained by plotting the flaw depth (a/t = 0.25) on the vertical axis and projecting horizontally to the crack growth curve. The service life is the time for the flaw to reach 75 percent of the vessel wall, which is approximately 10.7 years, as seen in Figure 6-16.

Example 5. This case considers a shallow surface flaw at the weld, which again requires use of Figure 6-2a, reproduced here as Figure 6-17. The flaw is 2 mm deep, or 12.5 percent of the wall thickness. Note that this value falls on the crack growth curve in Figure 6-17. In this case the flaw would Flaw Evaluation Charts April 2002 5967.doc-041202 Revision I

6-8 be predicted to follow the curve during future service, because the crack growth curve has been based on the smallest flaw size which would be predicted to grow. Therefore, the true service life would be over 9.2 years.

Example 6. This case is an axial surface flaw well below the weld region. From Figure 6-2a we obtain the appropriate curve for the crack growth prediction through the wall, and this is reproduced as the upper figure of Figure 6-18. This figure gives a service life estimate of approximately 5.2 years to through wall penetration. Additional life can still be added by considering the growth of the flaw up the tube to the bottom of the weld. This is illustrated in the bottom figure of Figure 6-18.

The bottom figure is taken from Figure 6-3. When the surface flaw grows through the wall, it will have increased in size by a factor of three, so its length will be 30 mm. If the flaw is centered at one inch below the weld, its new length after growth through the thickness is 15 mm (0.6 inches) above and 15 mm below its center point. This makes its upper extent at 2.6 inches. The additional service life for propagation to the bottom of the weld is approximately 0.8 years, making a total service life of approximately 6 years before the flaw would be predicted to violate the acceptance criteria.

It is clear from these examples that the most important figures for use in evaluating flaws in head penetrations are the surface flaw Figures 6-1 and 6-2 for axial flaws and 6-8 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.

Flaw Evaluation Charts 5967.doc-041202 April 2002 Revision I

6-9 Table 6-1 Summary of R.V. Head Penetration Acceptance Criteria Axial Circ Location ar f

af t

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: Flaws of any size in the attachment weld are not acceptable.

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

af

=

Flaw Depth as defined in JWB 3600 P

=

Flaw Length t

=

Wall Thickness Table 6-2 Example Problem Inputs Example Vertical Radial*

Penetration No.

Orientation Location Location Row Length Depth (t) 1 Axial Below Weld 00 Outer 10 mm.

4 mm.

2 Axial At Weld 180' Outer 10 mm.

4 mm.

3 Axial At Weld 00 Outer 10 mm.

8 mm.

4 Circumferential Above Weld 180' Outer 8 mm.

4 mm.

5 Axial At Weld 0o Outer 10 mm.

2 mm.

6 Axial 1" Below 0o Outer 10 5.3 mm.

Weld

• Note:

Centerside = 00 Lower Hillside = 180' Flaw Evaluation Charts 5967.doc-041202 April 2002 Revision I

6-10 1.0 0.9 0.8 0.7

) 0.6 2

.05 0.4

02.

0.1 0.0 0

Figure 6-1 a 3

6 9

12 15 18 Time(year)

Crack Growth Predictions for Longitudinal Inside Surface Flaws in the Head Penetrations in the D. C. Cook Unit 1 at and Below the Attachment Weld Flaw Evaluation Charts 5967.doc-041202 21 April 2002 Revision 1

6-11 1.0..

-T 0.8 Ai e reteia t e nd Aoe Ne Id

=0 75 a A-0 MO am a

MU e

W a

a a

M a

a 0

Rw M

0.7 O0.6

0.

ztý 0.5 0.3 0.2 4000 0.1 0.0 0

10 20 30 40 50 60 70 80 Time(year)

Figure 6-1b Crack Growth Predictions for Longitudinal Inside Surface Flaws in the Head Penetrations at the D. C. Cook Unit 1 Above the Attachment Weld Flaw Evaluation Charts 5967.doc-04] 202 April 2002 Revision I

2 4

6 8

10 12 14 16 18 Time(year)

Crack Growth Predictions for Longitudinal Outside Surface Flaws in the Head Penetrations at the D. C. Cook Unit I Flaw Evaluation Charts 5967.doc-041202 April 2002 Revision I 6-12 1.0 0.9 0.8 0.7 S0.6 0.5 a5 S0.4 0.3 0.2 0.1 0.0 0

Figure 6-ic

aft (flaw depth/wall thickness)

• ri P

9 P

0 S

0~00 M

I II

• I i

I i

i i

'I s

I

I ii I

• m

_==

I I

I U

II I

I i0

]

o II.i W11

6-14 1 0 F

7-T i 09 I

I I

0.9I I

I ii

/

0.8 c,pta" cecriteiia: 3elo

=l1., Atar An Weld t=.75 08.

A bo 07 1

( 0.6

.2 0.5 0&3 0.2//

I I,

0.1, 0.0 0

1 2

3 4

5 6

7 8

9 10 11 12 Time(year)

Figure 6-2a Crack Growth Predictions for Longitudinal Inside Surface Flaws in the Head Penetrations at the D. C. Cook Unit 2 at and Below the Attachment Weld Flaw Evaluation Charts 5967.doc-041 2.02 April 2002 Revision 1

1.0 0.9 0.8 0.7 0 0.6 0.5 0.4 0.3 t 0.2 0.1 f I a

[ ceane(

0 iter a: tar

-II

-ýa d

10 OVE

/

We d a/

/

=0.

/

5

/

20

-M ra

/

t/

/-

30 I

40 Time(year)

Figure 6-2b Crack Growth Predictions for Longitudinal Inside Surface Flaws in the Head Penetrations at the D. C. Cook Unit 2 Above the Attachment Weld Flaw Evaluation Charts 5967.doc-041202 6-15 0.0 50 April 2002 Revision 1 C

I

Figure 6-2c Acceptance ?riteria:i n --

Be.lenuI 1.0 0.9 0.8 0-7 0 0.6 f-0 0.4 0.3 0.2 0.1 1.0t below wel(

lili..~

wuiB m

2.0 4.0 6.0 8.0 Time(year)

Crack Growth Predictions for Longitudinal Outside Surface Flaws in the Head Penetrations at the D. C. Cook Unit 2 Flaw Evaluation Charts 5967.doc-041202 6-16 7/i 0.75t

& above weld; ni ii.

0.0 0.0 10.0 April 2002 Revision I al0.i

1.0 0.9 0.8 0.7 o 0.6

.5Q5

-0.5 0.4 0.3 0.2 0.1 0.0 0

3 6

9 12 TIME(YEARS)

Figure 6-2d Crack Growth Prediction for Longitudinal Surface Flaws in the Head Vent Near the Attachment Weld Flaw Evaluation Charts 5967.doc-041202 April 2002 Revision 1 6-17 15

l "t

j h 0

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10 Trmc Figure 6-7a Crack Growth Predictions for Growth of Through-Wall Flaws in the Center Penetration at D. C. Cook Unit I Flaw Evaluation Charts 5967.doc-041202 6-26 6

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7-1 7

SUMMARY

AND CONCLUSIONS An extensive evaluation has been carried out to characterize the loadings and stresses which exist in the head penetrations at D. C. Cook Units 1 and 2. Three-dimensional finite element models were constructed, and all pertinent loadings on the penetrations were analyzed [6]. 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 circumferential 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 D. C. Cook Units 1 and 2, based on information from the literature as well as a compilation of 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.

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 as a result of any indications.

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 wastage in the ferritic steel of the vessel head, as the borated primary water concentrates due to evaporation.

Any indication is unlikely to propagate very far up the penetration above the weld, because the hoop stresses decrease in this direction, and this will cause it to slow down, and to stop before it reaches the outside surface of the head. This result supports the conclusion that it is extremely unlikely that leakage of any magnitude will occur.

Summary and Conclusions April 2002 5967.doc-041202 Revision 1

7-2 The high likelihood that the indication will not propagate up the tube beyond the vessel head ensures that no catastrophic failure of the head penetration will occur, since the indication will be enveloped in the head itself. which precludes the opening of the crack and limits leakage.

Summary and Conclusions 5967.doc-041202 April 2002 Revision 1

8-1 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.

McIlree, A. R., Rebak, R. B., Smialowska, S., "Relationship of Stress Intensity to Crack Growth Rate of Alloy 600 in Primary Water," Proceedings International Symposium Fontevraud 11, 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 5"' International Symposium on Environmental Degradation of Materials in Nuclear Power Systems-Water Reactors," August 25-29, 1991, Monterey, California.

4A.

Crack Growth and Microstructural Characterization ofAlloy 600 PWR Vessel Head Penetration Materials, EPRI, Palo Alto, CA. 1997. TR-109136.

4B.

Ia,c,e 4C.

Bamford, W. H., "D. C. Cook Unit 2 Upper Head Penetration Crack Growth Determined from Inspection Data," Westinghouse Electric Report LTR-SMT-0 1-72, November 2001.

5A.

McGowan, J. J. and Raymund, M., "Stress Intensity Factor Solutions for Internal Longitudinal Semi-elliptic Surface Flaw in a Cylinder Under Arbitrary Loading," ASTM STP 677, 1979, pp. 365-380.

5B.

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 Piping, PVP Vol. 58, ASME, 1982, pp. 3748.

6.

a,c,e

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.

Letter from V. Vanderburg to B. Mickatavage, American Electric Power, "Best Estimate Hot Leg Temperature for D.C. Cook Unit 2," July 21, 1994.

References April 2002 5967.doc-041202 Revision 1

A-2

10.

Letter from V. Vanderburg to W. Bamford, American Electric Power, "Best Estimate Hot Leg Temperature for D. C. Cook Unit 1," October 14. 1994.

11.

Donald C. Cook Nuclear Plant Units I and 2. NRC Bulletin 2001-01 Response (TAC Numbers MB2624 and MB2625), September 4, 2001.

References 5967 doc-041202 April 2002 Revision I

A-1 APPENDIX A ALLOWABLE AREAS OF LACK OF FUSION: WELD FUSION ZONES There are two fusion zones of interest for the head penetration 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 D. C. Cook Unit 1 is 303°C (578OF) Unit 2 is 316'C (601 F), so both 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 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 potential 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 OF) is:

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

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 Sm = 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 presentation.

Appendix A April 2002 5967.doc-041202 Revision 1

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 in2, and using the ASME Code criteria, the area required is 2.79 in2. 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 (600°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 April 2002 5967.doc-041202 Revision t

Location of Axial Croc CRDM Thermal Sleeve

[al Penetration Weld Figure A-i Typical Head Penetration Appendix A 5967.doc-041202 A-3 April 2002 Revision 1

310 320 330 340 350 Circumferential Extent (Degrees)

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100 Circumferential 200 300 Extent (Degrees)

Allowable Regions of Lack of Fusion for the Outermost Penetration Tube to Weld Fusion Zone April 2002 Revision 1

Allowable Critical 100 90 80 70 60 50 40 30 20 10 0

Circumferential Extent (Degrees)

Figure A-4 Allowable Regions of Lack of Fusion for all Penetrations: Weld to Vessel Fusion Zone Appendix A 5967.doc-041202 A-6 a-a1)

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100 200 April 2002 Revision 1

A-7 Lack of Fusion Critical

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