ML11109A084

Attachments B-D, to L-11-107, Reply to Request for Additional Information for the Review of the Davis-Besse Nuclear Power Station, Unit No. 1, License Renewal Application

ML11109A084
Person / Time
Site:	Davis Besse
Issue date:	04/15/2011
From:	FirstEnergy Nuclear Operating Co
To:	Office of Nuclear Reactor Regulation
References
L-11-107 86-9110440-000
Download: ML11109A084 (75)

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Enclosure B Davis-Besse Nuclear Power Station (DBNPS), Unit No. 1 Letter L-11-107 AREVA NP Document 86-910440-000 Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years 30 pages not including cover sheet

Controlled Document0402-01-FOl (20697) (Rev. 014, 04/13/2009)

NON-PROPRIETARY A CALCULATION

SUMMARY

SHEET (CSS)

AREVA Document No. 86 - 9110440 - 000 Safety Related: N Yes E]No Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for Title 60 Years PURPOSE AND

SUMMARY

OF RESULTS:

Intergranular separations in low alloy steel heat-affected zones under austenitic stainless steel weld claddings were detected in SA-508, Class 2 reactor vessel forgings manufactured to a coarse grain practice, and clad by high-heat-input submerged arc processes. BAW-10013 (issued in 1972) contains a fracture mechanics analysis that demonstrates the critical crack size required to initiate fast fracture is several orders of magnitude greater than the assumed maximum flaw size plus predicted flaw growth due to design fatigue cycles. The flaw growth analysis was performed for a 40 year cyclic loading, and an end-of-life assessment of radiation embrittlement (i.e., fluence at 32 EFPY) was used to determine fracture toughness properties. The report concluded that the intergranular separations found in B&W vessels would not lead to vessel failure. This conclusion was accepted by the AEC . To cover the period of extended operation, a fracture mechanics analysis was performed using current ASME Code requirements; that analysis is fully described inthis report.

In May 1973, the AEC issued Regulatory Guide 1.43, "Control of Stainless Steel Weld Cladding of Low-Alloy Steel Components". The guide states that intergranular separations "has been reported only inforgings and plate material of SA-508 Class 2 composition made to coarse grain practice when clad using high-deposition-rate welding processes identified as 'high-heat-input' processes such as the submerged-arc wide-strip and the submerged-arc 6-wire processes. Cracking was not observed in clad SA-508 Class 2 materials clad by 'low-heat-input' processes controlled to minimize heating of the base metal. Further, cracking was not observed in clad SA-533 Grade B Class 1 plate material, which is produced to fine grain practice. Characteristically, the cracking occurs only in the grain-coarsened region of the base-metal heat-affected zone at the weld bead overlap." The guide also notes that the maximum observed dimensions of these subsurface cracks is 0.5 inch x 0.165 inch.

An update of BAW-10013 was completed for the B&WOG as reported inAppendix C to BAW-2251A. While Davis-Besse was not included in this report, the methodology used to evaluate intergranular separations inthe DB-1 SA-508 Class 2 forgings is consistent with the methodology reported inAppendix C of BAW-2251A. The DB-1 specific analysis was performed for 60-years using the current fracture toughness information, applied stress intensity factor solutions, and fatigue crack growth correlations for SA-508 Class 2 material. The objective of the analysis is to determine the acceptability of the postulated underclad cracks for the period of extended operation using ASME Code,Section XI, IWB-3612 acceptance criteria.

The analysis was applied to two relevant regions of the RV: the beltline, and the nozzle belt. The DB-1 closure head/head flange will be replaced with a new closure head that was fabricated by AREVA using SA-508 Class 3 material, which is not susceptible to the subject intergranular separations. Replacement is scheduled prior to entering the period of extended operation so the closure head is not considered in this evaluation. Both axial and circumferential oriented flaws were considered in the evaluation; however, the detailed flaw evaluation was only performed for the bounding axially oriented flaws. All the significant normal/upset condition transients and emergency/faulted condition transients were evaluated inthe analysis. The fatigue crack growth analysis considered all the normal/upset condition transients with associated 60-year projected cycles for the period of extended operation.

An axially oriented, semi-elliptical surface flaw with an initial flaw size of 0.353-inch deep (approximately twice that which has been observed) and 2.12-inch long (approximately four times that which has been observed) with a 6:1 aspect ratio was conservatively assumed at each of the two regions. This is contrasted to the observed flaws which are subsurface with a maximum size of 0.165 inch deep by 0.5 inch long.

For an axially oriented flaw, the limiting location for satisfying the requirements of IWB-3612 is at the lower end of the nozzle belt forging where the thickness transitions from 8.438 to 12.0 inches. The maximum crack growth, considering normal/upset condition transients with associated 60-year projected cycles for the period of extended operation (Reference 4 and Reference 5 for 114 HU/CDs), was determined to be 0.043 inches, which results in a final flaw depth of 0.396 inches (Reference 6, Table 6-1). The maximum applied stress intensity factor for the normal and upset condition results in a fracture toughness margin of 3.67 which is greater than the IWB-3612 acceptance criterion of 4/10 (3.16). The maximum applied stress intensity factor for the emergency and faulted conditions results in a fracture toughness margin of 1.43, which is greater than the IWB-3612 acceptance criterion of ,/2 (1.41). It is therefore concluded that the postulated underclad cracks in the DB-1 reactor vessel are acceptable for continued safe operation through the period of extended operation.

THE DOCUMENT CONTAINS ASSUMPTIONS THAT SHALL BE THE FOLLOWING COMPUTER CODES HAVE BEEN USED IN THIS DOCUMENT: VERIFIED PRIOR TO USE CODENERSION/REV CODENERSION/REV YES

~NO AREVA NP Inc., an AREVA and Siemens company Page 1 of 30

1`10111ed Document A

AR EVA 0402-01-FOl (20697) (Rev. 014, 04/13/2009)

AR..EVAN , Document No. 86-9110440-000 AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years Review Method: ýq Design Review (Detailed Check)

F Alternate Calculation Signature Block P/R/A Name and Title and Pages/Sections (printed or typed) Signature LP/LR Date Prepared/Reviewed/Approved M.A. Rinckel P All Project Manager 41 146/_

-

A.D. Nana R All Supervisory Engineer .Q T.M. Wiger A 7/7/t P All Engineering Manager I Note: P/RJA designates Preparer (P), Reviewer (R), Approver (A);

LP/LR designates Lead Preparer (LP), Lead Reviewer (LR)

Project Manager Approval of Customer References (N/A if not applicable)

Name Title (printed or typed) (printed or typed) Signature Date N/A Mentoring Information (not required per 0402-01)

Page.2

Controlled Document A 0402-01-FOl (20697) (Rev. 014, 04/13/2009)

AR EVA AREVA NP Inc.,

Document No. 86-9110440-000 an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years Record of Revision Revision PageslSectionsl No. Date Paragraphs Changed Brief Description / Change Authorization 000 07/2010 All Original release

-I- t F +/- I

+ +

F + +

-I- I i + +

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Controlled Document A Document No. 86-9110440-000 AREVA AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years Table of Contents Page SIGNATURE BLOCK ................................................................................................................................ 2 RECORD OF REVISION .......................................................................................................................... 3 LIS T O F TA B LE S ..................................................................................................................................... 6 LIS T O F F IG U R E S ................................................................................................................................... 7 1 .0 INT R O D UC T IO N ........................................................................................................................... 8 1.1 Ba c k g ro u n d ...................................................................................................................................... 8 1.1.1 B&W Investigations in BAW-1 001 3-A .......................................................................... 8 1.1.2 Regulatory Guide 1.43 .................................................................................................. 8 1.2 P o stu la te d F law S ize ......................................................................................................................... 9 1.3 R e g io n s o f Inte re st ........................................................................................................................... 9 2.0 ANALYTICAL METHODOLOGY ............................................................................................ 11 2.1 Postulated Surface Flaw ........................................................................................................... 11 2.2 Stress Intensity Factor Solutions ................................................................................................ 12 2.3 Effect of Cladding Material on Stress Intensity Factor .............................................................. 13 2.4 Fatigue Crack Growth Model ................................................................................................... 14 2.5 Fracture Toughness Curves ............................................................ ... ..................................... 15 2.6 Flaw Acceptance Criteria ........................................................................................................... 16 3.0 GEOMETRY, VESSEL MATERIALS AND MATERIAL PROPERTIES ................................... 17 3 .1 G e o m e try ........................................................................................................................................ 17 3.2 Thermal and Mechanical Properties .......................................................................................... 18 3.3 Nil-Ductility Reference Temperature ........................................................................................ 18 3 .4 T ra n s ie nts ...................................................................................................................................... 19 3 .5 S tre s s e s ......................................................................................................................................... 20 3.5.1 Stresses for Operating Transients ............................................................................... 22 3.5.2 Discontinuity Factors ................................................................................................. 22 3.5.3 Stresses from Vessel Loads ........................................................................................ 23 3.5.4 Stresses from Nozzle Loads ...................................................................................... 25 3.5.5 Nozzle Interaction Stresses ........................................................................................ 25 4.0 SUM MARY OF RESULTS AND CONCLUSIONS ................................................................. 27 Page 4

Controlled Document A Document No. 86-9110440-000 AREVA NP Inc., "'on roor/,iy/,$

an AREVA and Siemens company PROPRIE A/

Fracture Mechanics Analysis of Postulated Underciad Cracks in the DB-1 Reactor Vessel for 60 Years Table of Contents (continued)

Page 4.1 S um m ary of R esults [7] ................................................................................................................. 27 4 .2 C o n clusio n s .................................................................................................................................... 27 5 .0 REF E RENCES ............................................................................................................................ 29 Page 5

Controlled Document A

AR VA Document No. 86-9110440-000 AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years List of Tables Page Table 3-1: Reactor Vessel Dim ensions [7] ........................................................................................ 17 Table 3-2: Reference Temperatures at 52 EFPY ............................................................................. 18 Table 3-3: Transient Events for Flaw Evaluation ............................................................................. 19 T a b le 4 -1: S um m a ry of R e sults ............................................................................................................. 28 Page 6

Controlled Document AD AR EVA Document No. 86-9110440-000 rnc.,

AREVA NP an AREVA and Siemens company PrOP 60 Years Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for List of Figures Page Figure 1-1: Davis Besse Unit 1 Reactor Vessel ............................................................................... 10 Figure 2-1: Postulated Surface Flaw ................................................................................................. 11 Figure 3-1: Reactor Vessel Locations ............................................................................................. 21 Figure 3-2: Reactor Vessel Locations for General Shell Stresses .................................................... 24 Page 7

Controlled Document A

AREVA Document No. 86-9110440-000 AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years

1.0 INTRODUCTION

Beginning in 1970, reactor vessel manufacturers have been detecting intergranular separations in the heat-affected zones of low alloy steel SA-508 Class 2 forgings melted to coarse grain practice and clad with austenitic stainless steel weld metal using high-heat-input submerged-arc welding processes. A fracture mechanics analysis, performed in 1971 to address this so-called underclad cracking phenomenon, demonstrated that the critical crack size required to initiate fast fracture was several orders of magnitude greater than the maximum postulated flaw size plus predicted flaw growth over a 40 year design life. This analysis utilized fracture toughness data based on radiation embrittlement at end-of-life conditions (i.e., fast neutron fluence at 32 EFPY). The final report, BAW-10013-A [1],

concluded that postulated intergranular separations in B&W-designed vessels would not lead to vessel failure. The present analysis updates the 1971 analysis for the Davis Besse Unit 1 (DB-1) reactor vessel using current ASME Code requirements and extends the operating period to 60 years. Inside surface flaws are postulated to be present at several locations on the inside surface of the reactor vessel (RV) that are susceptible to underclad cracking. The analysis is performed to determine the acceptability of the postulated flaws under design basis loading conditions for the period of extended operation using the acceptance criteria of the ASME Boiler and Pressure Vessel Code, Section. XI, Article IWB-3612 [2]. The DB-1 underclad cracking evaluation for 60 years summarized herein is based on the detailed calculation reported in Reference [7].

1.1 Background Two key developments in the treatment of underclad cracking are the B&W investigations in BAW-10013-A and the issuance of Regulatory Guide 1.43 1.1.1 B&W Investigations in BAW-10013-A The results of the B&W investigations in BAW-1 0013-A confirmed that intergranular separations subject flaws only occur in SA-508 Class 2 forgings manufactured to a course grain practice and clad by high-heat-input submerged-arc welding processes such as the six-wire, strip, and the two-wire processes.

Furthermore, these defects have only been detected in the heat-affected zone immediately below the clad/base metal interface and the size of these defects did not exceed 0.100 inch in depth and 0.500 inch in length. A depth of 0.156 inch was used in the fracture mechanics analysis summarized in BAW-10013-A since it was the maximum discontinuity depth observed throughout the industry. The investigations also noted that no anomalies were observed in SA-533 Grade B, Class 1 plate materials clad by any of the high-heat-input welding processes.

1.1.2 Regulatory Guide 1.43 In May 1973, the Atomic Energy Commission issued Regulatory Guide 1.43 entitled "Control. of Stainless Steel Weld Cladding of Low-Alloy Steel Components," [3]. This guide states that "Underclad cracking has been reported only in forgings and plate material of SA-508 Class 2 composition made to coarse grain practice when clad using high-deposition-rate welding processes identified as "high-heat-input" processes such as the submerged-arc wide-strip and the submerged-arc 6-wire processes.

Cracking was not observed in SA-508 Class 2 materials clad by "low-heat-input" processes controlled.

to minimize heating of the base metal. Further, cracking was not observed in clad SA-533 Grade B Class 1 plate material, which is produced to fine grain practice. Characteristically, the cracking occurs Page 8

Controlled Document AR Document No. 86-9110440-000 AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years only in the grain-coarsened region of the base-metal heat-affected zone at the weld bead overlap." The guide also notes that the maximum size of these subsurface cracks is 0.5 inches x 0.165 inches.

1.2 Postulated Flaw Size The size of the postulated underclad crack to be considered in the present fracture. mechanics analysis is conservatively taken as the maximum size of subsurface cracks noted in Regulatory Guide 1.43. The subsurface dimensions of the postulated underclad crack are therefore taken to be 0.165 inch in depth and 0.500 inch in length.

1.3 Regions of Interest Based on the discussion above, the only regions of the reactor vessel that are susceptible to underclad cracking are those fabricated from SA-508 Class 2 forgings manufactured to a coarse grain practice and clad by a high-heat-input submerged arc welding process. SA-508 Class 2 material is used extensively in the DB-1 reactor vessel between the lower and upper heads, from the Dutchman forging to the reactor vessel flange. The lower head is made from SA-533 Grade B modified plate material which is not susceptible to underclad cracking. Furthermore, since the present fracture mechanics analysis for underclad cracking is intended to support licensing activities for plant operation beyond 2017, and the present reactor vessel closure head is scheduled to be replaced in 2014 with a one-piece SA-508 Class 3 forged head, the closure head region need not be considered further in the context of the present flaw evaluations for underclad cracking. Referring to Figure 1-1, the remaining regions of interest for the Davis Besse reactor vessel are:

a) Reactor vessel flange forging - SA-508 Class 2 b) Nozzle belt forging - SA-508 Class 2 c) Upper and lower shell forgings - SA-508 Class 2 d) Dutchman forging - SA-508 Class 2 Page 9

Controlled Document A Document No. 86-9110440-000 AREVA AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years Reactor Vessel Flange Nozzle Belt Forging ADB 203; 123Y317 Outlet Nozzle Forging Weld WF-232 Inside 9%

Weld WF-233 Outside 91%

Upper Shell Forging AKJ 233 - 123X244

• - Weld WF-182-1 4 Lower Shell Forging AKJ 233 - 123X244

. Weld WF-232 Inside 12%

Weld WF-233 Outside 88%

Dutchman Forging 122Y284VA1 Weld WF-1 82-1 Lower Head C6168-3 Figure 1-1: Davis Besse Unit I Reactor Vessel Page 10

Controlled Document A Document No. 86-9110440-000 AREV VA AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years 2.0 ANALYTICAL METHODOLOGY Linear elastic fracture mechanics analysis is used to determine the acceptability of intergranular separations in the heat-affected zone of coarse-grained SA-508 Class 2 forging material that has been clad with austenitic stainless steel weld metal using a high-heat-input submerged-arc welding process.

Although an underclad crack would be fully contained in the underlying base metal, it is conservatively postulated that such a crack would extend through the cladding material to create a surface breaking flaw. These flaws may be oriented in both the longitudinal direction (lying in a radial plane of the reactor vessel) and circumferential direction (in a horizontal plane). It is further postulated that these flaw are located at the most highly stressed locations in the susceptible regions of the reactor vessel. The applied stresses from internal pressure and thermal gradients, determined by finite element stress analysis using the ANSYS [4] and PCRIT [5] computer codes, and from external vessel and nozzle loads caused by deadweight, thermal expansion, seismic events, and loss of coolant accidents.

Closed-form solutions from the literature are used to characterize the crack tip stress intensity factors needed for ASME Code flaw evaluations. Additional computation is performed to capture the effect on stress intensity factors of differential thermal expansions between the cladding and base metal. After calculating fatigue crack growth under cyclic loads for the period of extended operation, the final flaw size is evaluated against the acceptance criteria of the ASME Code.

2.1 Postulated Surface Flaw For purposes of analysis, an underclad crack is conservatively represented by the 6:1 semi-elliptical surface crack illustrated in Figure 2-1. Combining the design underclad flaw depth of 0.165 inch (from Section 1.2) with a 0.188 inch nominal cladding thickness produces a 0.353 inch deep surface flaw.

The 6:1 aspect ratio then produces a postulated flaw length of 2.118 inches, which is significantly greater than the design length of 0.500 inch (from Section 1.2).

l 6:1 semi-elliptical surface flaw length Not to scale01 5 eaainrgo 0.165" separation region

// 7) 0.353" flaw depth 0.188" cladding thickness iF Figure 2-1: Postulated Surface Flaw Page 11

Controlled Document AR A VA Document No. 86-9110440-000 AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years 2.2 Stress Intensity Factor Solutions The Raju-Newman solution for an internal semi-elliptical axial surface flaw in a cylindrical vessel is used as a general form of stress intensity factor equation for calculating crack tip stress intensity factors for arbitrary through-wall stress distributions. For a third-order stress distribution, expressed as aY(x) = A0 + Aix + A 2x 2 + A 3X3 where "x" is the radial distance from the inside surface of the cylinder, the Raju-Newman [6] influence coefficient solution for the stress intensity factor at the maximum flaw depth is K, =_ i-.*GjAiaj Qj=0 where 1 65 Q = flaw shape factor = 1 + 1.464(a/c) a = flaw depth c = flaw half-length = 3a Gj = influence coefficient corresponding to Aj stress term The Gj influence coefficients for an inside surface axial flaw are provided in Table 1 of Reference [6] for t/Ri = 0.1 and several values of a/c and aft, where "t" is the wall thickness. Quadratic expressions have been developed [7] to accurately interpolate between the tabulated values.

Three variations of the above solution are used to calculate stress intensity factors, depending on the type of load and the order of the available through-wall stressdistribution, as explained below.

a) Through-wall thermal gradient stresses are available from finite element stress analysis (ANSYS [4] and PCRIT [5]).

2 3 (TT = Ao + Aix + A2x + A3x KIT G=-I[GoAoa05 +G 1 Ala 15

+G 2 A 2 a 2 .5 + G 3 A 3 a3.]

b) Through-wall pressure stresses are available from finite element stress analysis (ANSYS [4]).

2 3 ap = B0 + Bax + B 2 x + B 3 x Using an additional constant term to account for pressure acting on the crack face, 15 2 5 KIp 1P[GO(Bo + 1)a0.5 + G 1 Bla + G2 B2 a + G 3 B 3 a3"]

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Controlled Document A Document No. 86-9110440-000 AR EVA AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years c) Surface stresses are available from beam-element vessel loading analysis, nozzle loads, and nozzle/shell interaction analysis. For these linear gradient stresses, C'LG = CO + 0 1x K 1_LG=r- [GOC 0 a05+G1Clal5]

2.3 Effect of Cladding Material on Stress Intensity Factor Although the one-dimensional finite element code PCRIT models the cladding in the thermal solution to calculate temperatures, stresses are only calculated in the base metal, and not in the cladding itself.

Thus the code does not consider stresses in the cladding when calculating stress intensity factors due to thermal loading. To account for this cladding effect, an additional stress intensity factor, Kiclad, is calculated separately and added to the stress intensity factor computed by PCRIT.

To identify the portion of the total stress intensity factor that is attributable to the presence of cladding and which is not considered by PCRIT, stress intensity factors are determined for two through-wall thermal stress distributions, one with and one "without" the effects of cladding. The term "without cladding" is used to describe the stress intensify factors currently calculated by PCRIT even though the base metal stresses include the effects of differential thermal expansion between the base metal and cladding materials. At a given flaw depth, the stress intensity factor due to cladding, Klcjad, is the difference between the stress intensity factors calculated with and without cladding.

Thermal stresses are calculated from PCRIT temperature output using the following equations from Timoshenko [8] for stress in a hollow cylinder, Ror r ]

Hoop stress: Go= v0 (-'*ri) Ro2 "R-i JTrdr + JTr dr-Tr2 Ri [ Ref. [8], Eqn. (255) ]

Ri Axial stress: 2 fTr dr - T [ Ref. [8], Eqn. (256)]

I-R Ri where r = radial position Ri = inner radius R, = outer radius T = temperature Stress intensity factors are then calculated for each thermal stress distribution using the Raju-Newman

[6] solution for KIT from Section 2.2.

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Controlled Document A

AR EVA Document No. 86-9110440-000 AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years 2.4 Fatigue Crack Growth Model Flaw growth due to cyclic loading is calculated using the fatigue crack growth rate model from Article A-4300 of Section XI [2],

da_

-- C, (AK )n, dlN where AKI is the stress intensity factor range in ksilin and da/dN is in inches/cycle. For a surface flaw in a water environment, AKI = Klmax - Klmin R = Klmr,n/Klmax 0 _<IR *<0.25: AKI < 17.74, n = 5.95 C, = 1.02x10-12 xS S= 1.0 AKI _>17.74, n = 1.95 C. = 1.01 x 10-7 x S S= 1.0 0.25 < R < 0.65: AKI < 17.74 [ (3.75R + 0.06) / (26.9R - 5.725) ]0.25, n = 5.95 C. = 1.02x10-12 xS S = 26.9R - 5.725 AKI _>17.74 [ (3.75R + 0.06) / (26.9R - 5.725) ].25, n = 1.95 Co = 1.01 X 10-7X S S = 3.75R + 0.06 0.65*< R < 1.0: AKI < 12.04, n = 5.95 Co = 1.02x10-12 xS S= 11.76 AKI _>12.04, n= 1.95 Co= 1.01 X 10-7 x S S =2.5 Page 14

Controlled Document A Document No. 86-9110440-000 AREVA AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 YearsSection XI [2] also specifies that the following in-air rates must be used if it is greater than the in-water rates specified above.

n = 3.07 Co = 1.99x10"1 0 xS 0*_ R < 1: S = 25.72 x (2.88 - R)-3°7 AKI = Klmax - Klmin

<R<0: S 1 AKI = Klmax R<2: S= 1 AK= (1 - R) x Klmax / 3 The fatigue crack growth calculations contain an explicit check to ensure that the maximum crack growth rate is used on the present flaw evaluations.

2.5 Fracture Toughness Curves From Article A-4200 of Section XI [2], the lower bound Kia fracture toughness for crack arrest can be expressed as Kia = 26.8 + 12.445 exp [ 0.0145 (T - RTNDT)],

where T is the crack tip temperature, RTNDT is the nil-ductility reference temperature of the material, KIa is in units of ksi,,in, and T and RTNDT are in units of OF. In the present flaw evaluations, K15 is limited to a maximum value of 200 ks[i/in (upper-shelf fracture toughness).

A higher measure of fracture toughness is provided by the Kic fracture toughness for crack initiation, approximated in Article A-4200 of Section XI [2] by Kjc = 33.2 + 20.734 exp [ 0.02 (T - RTNDT)].

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Controlled Document A

AR EVA Document No. 86-9110440-000 AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years 2.6 Flaw Acceptance Criteria Article IWB-3612 of Section XI [2] requires that the applied stress intensity factor at the final flaw size be less than the available fracture toughness at the crack tip temperature, with appropriate safety factors, as described below.

For normal and upset conditions:

K,(af) < Ka/ 10V where K1(af) = the maximum applied stress intensity factor for normal and upset conditions at the final flaw size af KIa = the available fracture toughness based on crack arrest for the corresponding crack tip temperature For emergency and faulted conditions:

Ki(af) < Kir 1-F where KI(af) = the maximum applied stress intensity factor for emergency and faulted conditions at the final flaw size af Kic = the available fracture toughness based on fracture initiation for the corresponding crack tip temperature Page 16

Controlled Document A Document No. 86-9110440-000 AREVA AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years 3.0 GEOMETRY, VESSEL MATERIALS AND MATERIAL PROPERTIES The input to the present flaw evaluations consists of the reactor vessel geometry, material properties, descriptions of transients, and relevant stress data.

3.1 Geometry The pertinent dimensions of the reactor vessel forgings are listed in Table 3-1. The nominal cladding thickness is 0.188 inch [7]

Table 3-1: Reactor Vessel Dimensions [7]

Nozzle Belt Forging Diametric Dimensions Radial Dimensions Cladding inside surface, in. 168.000 84.000 Minimum base metal thickness, in. 12.000 Upper and Lower Shell Forgings Diametric Dimensions Radial Dimensions Cladding inside surface, in. 171.000 85.500 Minimum base metal thickness, in. 8.438 Page 17

Controlled Document A Document No. 86-9110440-000 AREVA AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years 3.2 Thermal and Mechanical Properties The DB-1 reactor vessel forgings for the nozzle belt (Part No. 169), upper shell (Part No. 170), lower shell (Part No. 171), and the lower head (Dutchman forging - Part No. 181) are fabricated from SA-508 Class 21 material of nominal composition 3/4Ni- 1/2Mo- 1/ 3Cr-V. The austenitic cladding is assumed to be 18Cr-8Ni stainless steel material. The PCRIT computer code only accepts constant temperature thermal properties for calculating temperatures, but does admit temperature dependent mechanical properties for calculating stresses. Thermal and mechanical properties for the base metal and cladding materials were obtained from Reference [9].

3.3 Nil-Ductility Reference Temperature The reference temperature index for the nil-ductility transition region of the fracture toughness curves, RTNDT, is a function of the fast neutron fluence and the chemical composition of the material. Values of RTNDT are obtained from a recent determination of pressurized thermal shock temperatures for the DB-1 reactor vessel at projected fluence values for 52 EFPY of operation over 60 calendar years [10]. The limiting values of fluence and reference temperature are listed in Table 3-2 for the four regions of interest for underclad cracking. The unirradiated value of 60°F [11] is used for the reference temperature of the reactor vessel flange since the fluence at this location would be lower than the value of 2.57E+16 n/cm 2 at the top of the inlet nozzle forging [12]. This fluence value is less than the threshold value of 1.OE+17 n/cm 2 , below which a material surveillance program is not required to monitor changes to fracture toughness due to exposure to neutron irradiation [13].

Table 3-2: Reference Temperatures at 52 EFPY Forging Material Location Inside Surface Fluence* RTNDT RV Flange At Weld to Nozzle Belt <2.57E+16 n/cm 2 60.0 0 F Nozzle Belt At Weld WF-232 2.27E+18 n/cm 2 81.2 OF (Upper)

Lower Shell At Weld WF-182-1 1.68E+19 n/cm 2 95.7 OF Dutchman At Weld WF-232 2.28E+17 n/cm 2 80.8 OF (Lower)

• at the inside surface of the base metal material.

The current ASME Code designation for this material is SA-508 Grade 2 Class 1.

Page 18

Controlled Document A Document No. 86-9110440-000 AREVA AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years 3.4 Transients The inside surface of the reactor vessel is subjected to transient loads in the form of primary coolant cold leg temperatures and pressures as defined in the DB-1 reactor coolant system functional specification [14]. The normal, upset, emergency, and faulted operating transient events that are considered in the present flaw evaluations are listed in Table 3-3, along with the number of occurrences over 60 years of plant operation, as defined in References [15] and [16]. This listing includes all transients in Table 2 of the functional specification except for a few makeup, letdown, and steam generator transients that are considered to be insignificant in terms of loading the reactor vessel.

Table 3-3: Transient Events for Flaw Evaluation 60-Year Tranient ASME Cycle Number* Transient Description Category Count 1A Heatup from 0% to 8% Power Normal 114 1B Cooldown from 8% to 0% Power Normal 114 2A Power Change from 0% to 15% Normal 205 2B Power Change from 15% to 0% Normal 94 3 Power Loading from 8% to 100% Normal 1800 4 Power Unloading from 100% to 8% Normal 1800 5 10% Step Load Increase Normal 67 6 10% Step Load Decrease Normal 140 7 Step Load Reduction from 100% to 8% Power Upset 12 8A Reactor Trip - Type A (Loss of RC Flow) Upset 4 8B Reactor Trip - Type B (Control Sys Malfunction) Upset 47 8C Reactor Trip - Type C (Loss of MFW Flow) Upset 26 9 Rapid Depressurization Upset 2 10 Change of Flow Upset 10 11 Rod Withdrawal Accident Upset 40 12 Hydrotest Test 4 14 Control Rod Drop Upset 18 15 Loss of Station Power Upset 6 16 Steam Line Failure Faulted 1 17A Loss of Feedwater to One Steam Generator Upset 6 17B Stuck Open Turbine Bypass Valve Emergency 2 21 Loss of Coolant Faulted I

• Refers to the transient numbers used in the reactor coolant system functional specification [14].

The temperature and pressure time-history curves in the functional specification [14] are discretized for input to the PCRIT. Individual transient events are then combined into complete cycles and these combined transients are used to calculate the cyclic variations in stress intensity factor that produce fatigue crack growth over the life of the plant.

Page 19

Controlled Document A Document No. 86-9110440-000 ARE VA AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years 3.5 Stresses The stress intensity factor solution described in Section 2.2 for the postulated inside surface flaw requires that stresses be defined in the reactor vessel wall in order to develop polynomial stress distributions in the through-wall direction. The required stress profiles can be either first or third-order in the radial distance from the inside surface, depending on the source of stress. The various loadings that are considered for an evaluation of underclad cracking are transient stresses from internal pressures and reactor coolant temperatures, stresses from external loads on the vessel and nozzles, and discontinuity stresses (or adjustments to calculated stresses to account for local geometric discontinuities).

Several locations within the reactor vessel are considered for evaluation, as indicated in Figure 3-1.

The nozzle belt location at the reactor vessel flange is selected since stresses would be expected to be high since the thick flange forging serves as a stress riser for stresses in the thinner nozzle belt.

Similarly, the nozzle belt is evaluated at the transition from the thinner shell forging thickness (8.438") to the thicker nozzle belt thickness (12"). Although remote from any structural discontinuity, the lower shell forging will be evaluated at the mid-beltline elevation due to a high value of fast neutron fluence at this location. An earlier fracture mechanics analysis of potential underclad cracking in skirt-supported reactor vessels, performed in 1996 for several B&W-designed reactor vessels [17], showed that the lower shell at the weld to the support skirt was not a controlling location. Since stresses would be lower in the DB-1 nozzle-supported vessel at a comparable location, and the fluence level is lower at the Dutchman forging than at the mid-beltline elevation, the Dutchman forging location shown in Figure 3-1 need not be evaluated in the present fracture mechanics analysis.

Stresses are determined at each reactor vessel location identified where a flaw evaluation is to be performed to address the potential for underclad cracking; i.e., under the reactor vessel flange, at the transition in thickness, and at the mid-beltline elevation. The basic stress data is obtained from the PCRIT analysis for the transient conditions defined in Section 3.4. The remaining vessel and nozzle stresses, which are only available for selected operating conditions, are then scaled by the governing transient data and combined with the detailed stresses obtained from PCRIT.

Page 20

Controlled Document A Document No. 86-9110440-000 AR EVA AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years

-Nozzle Belt at RV Flange I Evaluation Point 3

- Nozzle Belt at Transition I EvaluationPoint Lower Shell Forging at Weld WF-1 82-1 Evaluation Point 1 Dutchman Forging at Weld WF-232 (Lower)

Figure 3-1: Reactor Vessel Locations Page 21

Controlled Document A

AREVA Document No. 86-9110440-000 AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years 3.5.1 Stresses for Operating Transients PCRIT transient thermal and stress analysis is performed for each transient combination. Input to PCRIT computer code consists of the data presented in Sections 3.1 through 3.4, as well as the convective heat transfer film coefficients developed for the beltline shell and nozzle belt forgings, respectively. In addition to calculating temperatures and stresses at selected times during the transient, PCRIT computes stress intensity factors at pre-determined flaw depths in the vessel wall, ranging from 1/40th to 32/ 4 0 th of the wall thickness. Critical time points during a transient are selected by reviewing the PCRIT output to determine the times when the maximum and minimum stress intensity factors occur at a representative flaw depth, say 2/4 0 th of the wall thickness (or 2/4 0t). This flaw depth will generally bound the actual flaw depth after fatigue crack growth. Once the critical transient time points are selected, the corresponding transient stresses are extracted from the transient results and used to compute stress intensity factors at each calculated increment of fatigue crack growth, as determined from the fatigue crack growth model described in Section 2.4. The required thermal stresses are extracted from supplemental PCRIT results files in the form of the Ai stress coefficients defined in Section 2.2 for calculating KIT.

Transient pressure stresses are expressed in terms of the Bi unit stress coefficients defined in Section 2.2 for calculating Kip. Pressure stresses have been determined for each evaluation point identified in Figure 3-1 from a generic axisymmetric finite element analysis of B&W-designed reactor vessels [18].

3.5.2 Discontinuity Factors The thermal stresses calculated by the PCRIT computer code are only applicable to cylindrical shell sections remote from structural discontinuities. However, the same finite element stress analysis used to derive stress coefficients for the closure flange preload and a unit pressure load can be used to develop scale factors, or discontinuity factors, to account for the presence of stress risers in the reactor vessel wall on thermal stresses calculated by PCRIT. In addition, pressure stresses from the axisymmetric finite element analysis can be used to develop discontinuity factors for the linear gradient stresses introduced in Section 2.2. An example of a linear gradient stress is the inside surface stress calculated from a vessel beam model.

While PCRIT thermal stresses are computed for all analyzed transients for remote shell locations, detailed finite element stresses are only available at structural discontinuities for preload, heatup, steady state, and cooldown conditions. The logic for developing thermal discontinuity factors is to consider that at any time during a thermal transient, the relative magnitude and shape of the thermal stress profiles at the structural discontinuity and the remote location remain nearly constant. Ratios of the four Ai thermal stress coefficients, defined in Section 2.2, between the two locations would then be similar for any "heating type" transient, and likewise for any "cooling type" transient. These ratios, or thermal gradient discontinuity factors, can then be developed for representative heating and cooling transients, such as reactor heatup and cooldown, where thermal stresses are available. Accordingly, discontinuity factors are calculated at 5 hours5.787037e-5 days <br />0.00139 hours <br />8.267196e-6 weeks <br />1.9025e-6 months <br /> into reactor heatup (at the end of the heatup ramp) and at 8 hours9.259259e-5 days <br />0.00222 hours <br />1.322751e-5 weeks <br />3.044e-6 months <br /> into reactor cooldown (at the end of the cooldown ramp). By classifying each thermal transient analyzed by PCRIT as being either a heating or cooling type transient (at the time points chosen for evaluation), the appropriate thermal gradient discontinuity factor can be selected for use in adjusting the Ai thermal stress coefficients from PCRIT. These adjusted stress coefficients may then be used to calculate the thermal stress intensity factor at the structural discontinuity.

Page 22

Controlled Document AR EVA Document No. 86-9110440-000 AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years Discontinuity factors are developed for linear gradient stresses at the inside and outside surfaces using finite element stress results for steady state conditions, expressed in terms of pressure stresses. Since surface stresses are not available from the finite element results, nonlinear extrapolation is used to obtain the desired stresses. A linear gradient stress discontinuity factor is simply the ratio of the inside or outside surface pressure stress at a structural discontinuity to the comparable stress at a remote location in the vessel.

3.5.3 Stresses from Vessel Loads General reactor vessel shell stresses have been developed [19] for deadweight (DW), thermal expansion (TH), operating basis earthquake (OBE), safe shutdown earthquake (SSE), and loss of coolant accident (LOCA) loads from an isolated beam model of the reactor vessel and service support structure. These stresses, which are provided at several locations within the reactor vessel, as shown in Figure 3-2, were calculated at the inside and outside surfaces of the vessel for four conservative combinations of individual load cases,

1. (DW+TH) + OBE

2. OBE 2 2

3. (DW+TH) + [ (SSE) + (LOCA) ]1/2 2 2

4. [ (SSE) + (LOCA) ]1/2 according to the relationships, Longitudinal Stress = Axial Force / Area of Shell + Bending Moment x Radius I Moment of Inertia Circumferential Stress = Longitudinal Stress x Poisson's Ratio In order to mitigate conservatism in the present fracture mechanics analysis for underclad cracking, and to obtain the individual load case stresses needed for the fatigue crack growth and final flaw evaluation phases of the analysis, stresses are derived in this section for individual loads using the general methodology provided in the reactor vessel stress input document [19]. Stresses are calculated for flaw evaluation points 1, 2, and 3, respectively, using scale factors described in Section 3.5.2 to adjust for structural discontinuities, as applicable. The correlation between flaw evaluation points and locations in the vessel beam model is presented below.

Flaw Vessel Evaluation Description Model Point Location 1 Shell Forging at Mid-Beltline 6 2 Nozzle Belt at Transition 4 3 Nozzle Belt at RV Flange 2 Page 23

Controlled Document A Document No. 86-9110440-000 AR EVA AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years Figure 3-2: Reactor Vessel Locations for General Shell Stresses Page 24

Controlled Document A

AR EVA Document No. 86-9110440-000 AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years 3.5.4 Stresses from Nozzle Loads The reactor vessel stress input document [19] provides stresses in the vessel shell from external piping loads acting on the reactor vessel inlet and outlet nozzles. Inside and outside surface stresses were calculated at the locations shown in Figure 3-2 for four conservative combinations of individual load

cases,

1. (DW+TH) + OBE

2. OBE 2 2

3. (DW+TH) + [ (SSE) + (LOCA) ]11/2 2

4. [ (SSE) 2 + (LOCA) ]1/2 Stresses are calculated from nozzle loads in a two step process. First, surface stresses are calculated at the base of the nozzle using the Bijlaard method as documented in Welding Research Council Bulletin 107 [20]. These stresses are then attenuated to other locations in the vessel by treating the section of the vessel wall between the center of the nozzle and the point of interest as a semi-infinite beam on an elastic foundation.

In order to reduce conservatisms inherent in the reactor vessel stress input document [19] and to generate individual load case stresses for use in the fracture mechanics analysis, stresses in the reactor vessel shell are calculated in for individual loads using the general methodology provided in Reference [19].

3.5.5 Nozzle Interaction Stresses The nozzle interaction stresses are discontinuity stresses in the reactor vessel shell at the outlet nozzle-to-shell juncture. These stresses, which are the difference between the actual stresses at the structural discontinuity and the stresses in a remote region of the nozzle belt shell, are available from the fracture mechanics input document [19] for several loading conditions. Nozzle interaction stresses are attenuated to other locations in the reactor vessel in the same manner as the stresses at the base of the nozzle (Bijlaard stresses) that were calculated in Section 3.5.4 from external nozzle loads.

First introduced in 3.5.2 to adjust thermal gradient stresses calculated at remote locations for use at structural discontinuities, reference heating and cooling load cases are selected from the documented loading conditions for use in determining nozzle interaction stresses for other transient conditions.

Stresses at the end of the reactor heatup to 100% power are used for reference "heating type" transient stresses, while stresses at the end of the cooldown ramp are selected for use as the reference "cooling type" transient stresses.

The reference nozzle interaction stress coefficients for heating and cooling transients are calculated for locations where flaw evaluations will be performed, using the appropriate surface stress discontinuity factors.

The local reference nozzle interaction stress coefficients are used to calculate nozzle interaction stress intensity factors for a particular thermal transient by first calculating a set of linear gradient stress coefficients to characterize the through-wall thermal stress profile. The reference nozzle interaction stress coefficients at a particular location will then be multiplied by the ratio of transient thermal stress coefficients to similar thermal stress coefficients derived for the reference heating or cooling thermal Page 25

Controlled Document A Document No. 86-9110440-000 AREVA AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years stress profile, as applicable. Thermal stress coefficients are obtained from PCRIT thermal stress results in the same spreadsheets used to calculate Kiclad.

Page 26

Controlled Document A

AR VA Document No. 86-9110440-000 AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years 4.0

SUMMARY

OF RESULTS AND CONCLUSIONS Fracture mechanics analysis has been used to perform crack growth and flaw tolerance analyses to evaluate conservatively postulated 6:1 semi-elliptical underclad flaws extending 0.165 inches into the low alloy steel forging materials of the reactor vessel for normal, upset, emergency, and faulted design conditions. The total initial flaw depth, including cladding, was 0.353 inch, corresponding to a length of 2.12 inches along the inside surface of the cladding for the postulated 6:1 flaw aspect ratio. Based on a combination of stress levels and predicted values of fast neutron fluence at 52 EPFY, flaw evaluations were performed at three locations in the reactor vessel; the lower shell forging at the mid-beltline elevation, the nozzle belt at the change in thickness, and the nozzle belt at the reactor vessel flange.

The reactor vessel closure head was not included in the flaw evaluations for underclad cracking since a one-piece SA-508 Class 3 replacement head is planned to be in place by 2014, prior to the period of extended operation under license renewal, which begins on April 22, 2017.

Furthermore, since the present fracture mechanics analysis for underclad cracking is intended to support licensing activities for plant operation beyond 2017, and the present reactor vessel closure head is scheduled to be replaced in 2014 with a one-piece SA-508 Class 3 forged head, the closure head region need not be considered further in the context of the present flaw evaluations for underclad cracking 4.1 Summary of Results [7]

The final fracture toughness margin for a postulated axial flaw in the beltline region is 4.88 for normal and upset conditions and 2.59 for emergency and faulted conditions.

A postulated axial flaw at the transition region of the nozzle belt would produce margins of 3.67 for upset conditions and 1.43 for emergency and faulted conditions. For a circumferential flaw at this location, these margins are 4.39 for upset conditions and 1.78 for emergency and faulted conditions.

At the reactor vessel flange region of the nozzle belt, an axial flaw would result in fracture toughness margins of 4.36 for upset conditions and 1.75 for emergency and faulted conditions. A circumferential flaw at this location would produce margins of 7.40 for upset conditions and 2.94 for emergency and faulted conditions.

These results show that fracture toughness margins are higher for flaws oriented in a circumferential plane. Table 4-1 summarizes the results of underclad cracking flaw evaluations for postulated flaws in the controlling axial orientation.

4.2 Conclusions The results presented in Table 4-1 demonstrate that the potential for underclad cracking in the DB-1 reactor vessel are within the acceptable margins of the ASME Code for fracture toughness of the susceptible SA-508 Class 2 forging material for the period of extended operation. For an axially oriented flaw, the limiting -normal and upset condition fracture toughness margin is 3.67, which is greater than the value of 3.16 required by Section X1, Article IWB-3612 [2]. Likewise, the calculated margin for emergency and faulted conditions of 1.43 is greater than the minimum value of 1.41 required Page 27

Controlled Document A Document No. 86-9110440-000 AREVA AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years by the Code. It has also been demonstrated by analysis that the fracture toughness margins calculated for circumferentially oriented flaws are consistently higher than those for axial flaws.

It is further noted that the size of the postulated underclad cracks considered in the flaw evaluations are at least as large as the 0.165 inch deep by 0.500 inch long subsurface crack mentioned in Regulatory Guide 1.43 [3] as the maximum extent of underclad cracking that has been observed in the industry.

Therefore, the postulated underclad cracks in the DB-1 reactor vessel are acceptable for continued safe operation through the period of extended operation.

Table 4-1: Summary of Results Initial Flaw Depth = 0.353 inch Location in Final flaw Loading Controlling Calculated Required Reactor Vessel depth, af Condition (1) KI(af) (2) Fracture Fracture Toughness Toughness (inches) (ksiqin) Margin Margin Shell Forging N/U 41.01 4.88 /10 and Mid-Beltline 0.380 Elevation ElF 77.31 2.59 /2 Nozzle Belt at N/U 54.44 3.67 */10 Thickness 0.396 Transition ElF 139.6 1.43 /2 Nozzle Belt at N/U 45.88 4.36 /10 Reactor Vessel 0.390 Flange ElF 114.4 1.75 N12 Notes: (1) N/U = Normal and upset conditions E/F = Emergency and faulted conditions (2) Controlling stress intensity factor occurs at the final flaw depth for an axially oriented flaw Page 28

Controlled Document AR A VA Document No. 86-91 10440-000 AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years

5.0 REFERENCES

1. BAW-1001 3-A, Study of Intergranular Separations in Low-Alloy Steel Heat-Affected Zones Under Austenitic Stainless Steel Weld Cladding," October 1972.

2. ASME Boiler and Pressure Vessel Code,Section XI, Rules for Inservice Inspection of Nuclear Power Plant Components, 1995 Edition with Addenda through 1996.

3. U.S. Atomic Energy Commission Regulatory Guide 1.43, "Control of Stainless Steel Weld Cladding of Low-Alloy Steel Components," May 1973.

4. ANSYS Finite Element Computer Code, Version 10.0, ANSYS Inc., Canonsburg, PA.

5. AREVA NP Document 32-1174278-007, "Verification of PCRIT 6.3 & User's Manual," May 2010.

6. I.S. Raju and J.C. Newman, Jr., "Stress-Intensity Factors for Internal and External Surface Cracks in Cylindrical Vessels," Journal of Pressure Vessel Technology, American Society of Mechanical Engineers, Vol. 104, pp 293-298, November 1982.

7. AREVA NP Document 32-9110439-000, "DB-1 FM Analysis for 60 Years," June 2010.

8. S.P. Timoshenko and J.N. Goodier, Theory of Elasticity, Third Edition, McGraw-Hill Book Co.,

New York..

9. ASME Boiler and Pressure Vessel Code, Section II-Materials, Part D - Properties, 1995 Edition with Addenda though 1996.

10. AREVA NP Document 32-9123247-000, "RTPTS Values of Davis-Besse Unit 1 for 52 EFPY, Including Extended Beltline," November 2009.

11. BAW-1 0046A, Rev. 2, "Methods of Compliance with Fracture Toughness and Operational Requirements of 10 CFR 50, Appendix G," B&W Owners Group Materials Committee Topical Report, June 1986.

12. AREVA NP Document 86-9025792-001, 'Davis Besse Fluence Analysis - Cycles 13-14 Summary Report," October 2009.

13. U.S. Code of Federal Regulations, Title 10, "Domestic Licensing of Protection and Utilization Facilities," Appendix H to Part 50, "Reactor Vessel Material Surveillance Program Requirements," Federal Register, April 12, 2010.

14. AREVA NP Document 18-1149327-003, "Functional Specification for Reactor Coolant System for Davis-Besse," July 2008.

Page 29

Controlled Document AR A VA Document No. 86-9110440-000 AREVA NP Inc.,

an AREVA and Siemens company NON-PROPRIETARY Fracture Mechanics Analysis of Postulated Underclad Cracks in the DB-1 Reactor Vessel for 60 Years

15. AREVA NP Document 38-9121554-000, "Davis-Besse 60-Year Projected Transient Cycles,"

September 2009.

16. AREVA NP Document 38-9128551-000, "Additional Transient Cycle Data Applicable to the Davis-Besse Pressurizer Surge Line," December 2009.

17. AREVA NP Document 32-1245893-00, "FM Analysis of Postulated Underclad Cracks in B&W Designed RV for 48 EFPY," July 1996.

18. AREVA NP Document 32-1128225-00, "Stress Analysis of Reactor Vessel for L.E.F.M. Use,"

October 1981.

19. AREVA NP Document 32-9120525-000, "Davis Besse Unit 1 Reactor Vessel Stress Input for Fracture Mechanics Analysis," February 2010.

20. Welding Research Council Bulletin 107, "Local Stresses in Spherical and Cylindrical Shells Due to External Loadings", by K.R. Wichman, A.G. Hopper, and J.L. Mershon, August 1965, Welding Research Council, Inc., New York, NY.

Page 30

Enclosure C Davis-Besse Nuclear Power Station (DBNPS), Unit No. I Letter L-11-107 Centerior Energy Letter, "High Pressure Iniection/Makeup Nozzle and Thermal Sleeve Program Davis-Besse Nuclear Power Station Unit I (Serial No. 1968)."

(ADAMS Accession Number ML9109030090), August 23, 1991 4 pages not including cover sheet

CENTERIOR ENERGY Dondl C. Shelton 300.Madism Avenue Vice PuWsent. Nudar Toledo, OH 43652-0001 DavI&s 2492300

• (419)

Docket Number 50-346 License Number NPF-3 Serial Number 1968 August 23, 1991 United States Nuclear Regulatory Commission Document Control Desk Washington, D. C. 20555

Subject:

High Pressure Injection/Makeup Nozzle and Thermal Sleeve Program Davis-Besse Nuclear Power Station Unit 1 Gentlemen:

The purpose of this letter is to inform the Nuclear Regulatory Commission (NRC) of Toledo Edison's (TE) evaluation of High Pressure Injection (HPI) Makeup Nozzle Thermal Sleeve reliability. By letter dated May 3, 1990 (Serial Number 1802), TE summarized the actions taken during cycle 6 and the sixth refueling outage (6RFO) resulting from the discovery of the failed HPI/Makeup nozzle thermal sleeve during the fifth refueling outage (5RFO) at the Davis-Besse Nuclear Power Station (DBNPS) Unit 1. The major actions taken through the end of the 6RFO were focused on the assessment and preservation of the structural integrity of the nozzle which had experienced thermal cycling due to the thermal sleeve failure. Analysis demonstrated that an intact thermal sleeve effectively protects the HPI/Makeup nozzle from the effects of thermal cycling fatigue. Since the thermal sleeve failure discovered during the 5RFO was attributed to thermal cycling fatigue, induced by makeup flow cycling, control of makeup flow was improved as described in TE's letters to the NRC dated September 14, 1988 (Serial Number 1580), May 3, 1990 (Serial Number 1802) and May 25, 1990 (Serial Number 1808). Recognizing the importance of long term thermal sleeve reliability, TE committed to investigate the mechanisms which affect thermal sleeve life, and evaluate alternatives which might be pursued to ensure long term reliability. Toledo Edison's letter to the NRC dated December 3, 1990 (Serial Number 1871) provided the details of TE's plans and schedule to address thermal sleeve reliability.

Operating Companies

• Cleveland Electrc Illuminating 7oledo Edison

Docket Number 50-346 License Number NPF-3 Serial Number 1968 Page 2 Toledo Edison's plans focused on a fracture mechanics based prediction of thermal sleeve life under various makeup flow cycling conditions.

As contingencies, TE planned to investigate the impact of premature thermal sleeve failure on nozzle structural integrity, and to conduct a preliminary assessment of the feasibility of on-line monitoring for thermal sleeve presence. The favorable results of the fracture mechanics thermal sleeve lifetime prediction has obviated the need for further consideration of the impact of premature thermal sleeve failure and on-line monitoring for thermal sleeve presence. The fracture mechanics analysis predicts a lifetime exceeding 20 eighteen-month operating cycles in makeup service under current makeup flow control conditions. Cycle 7 is the first operating cycle in makeup service for the current HPI/makeup nozzle thermal-sleeve.

Babcock and Wilcox Nuclear Services (BWNS) and Structural Integrity Associates (SIA) performed the evaluation of thermal sleeve reliability for TE. Babcock and Wilcox Nuclear *Services, using the proprietary BWNS FLOIJTRAN computer code, provided fluid temperature distributions corresponding to the actual range of makeup flow rates. SIA used this information to develop stress distributions within the thermal sleeve for the range of makeup flow rates using the ANSYS computer code:

These results were used as input to a fracture mechanics analysis of the thermal sleeve in conjunction with makeup flow cycling conditions determined from actual recorded operating data. SIA's pc-CRACK computer code was used for the fracture mechanics analysis.

Recorded makeup flow data for several representative days in fuel cycles 5, 6 and 7 were analyzed to establish the makeup flow cycling conditions used in the fracture mechanics analysis. The data included normal power operation days during all three fuel cycles as well as some off-normal events such as trips and shutdowns.

The fracture mechanics model used to analyze the sleeve was chosen to be representative of the fracture behavior of the sleeve observed from metallurgical analysis of the broken sleeve pieces recovered from the reactor vessel during the 5RFO. A fracture mechanics model for an axial, semi-elliptical surface flaw of varying aspect ratio, in a cylinder, was selected from the pc-CRACK fracture mechanics model library. This model is consistent with the findings of the metallurgical analysis that fracture initiated predominantly on the inside surface of the sleeve, at or near the outlet end of the sleeve, and that the primary flaw propagation direction observed was outward from the inside surface with the flaw oriented axially on the thermal sleeve, indicating that axial extension in length was a consequence of the changing aspect ratio of the semi-elliptical flaw once it propagated through-wall. For this analysis, the pc-CRACK model was adjusted to calculate the stress intensity factor at the inside surface of the thermal sleeve. The stress intensity factor at this location is more important to flaw propagation in the axial direction than the stress intensity factor at the deepest point of flaw penetration, which is normally calculated by pc-CRACK.

Docket Number 50-346 License Number NPF-3 Serial Number 1968 Page 3 The stress intensity factor was calculated by pc-CRACK using this model and the ANSYS thermal sleeve stress distributions for the range of makeup flow rates. These pc-CRACK stress intensity factor results and the American Society of Mechanical Engineers (ASME) Boiler and Pressure Vessel Code,Section XI fatigue crack growth correlation for stainless steel were then used to predict flaw growth under the various makeup cycling conditions.

Using this approach, the calculated flaw growth rate for the makeup flow cycling conditions which were prevalent prior to the 5RFO compares favorably with the metallurgical analysis of the failed thermal sleeve.

The fracture mechanics anplysis yields predicted flaw growth rates on the order of 2.5 to 3.5 x 10- inches per makeup flow cycle.- The metallurgical analysis reported striation spacing of less than one micron whiSh corresponds to an approximate crack growth rate of less than 3.9 x 10- inches per stress cycle. Therefore, the fracture mechanics model predictions correlate well with the observed thermal sleeve failure.

Under the makeup flow cycling conditions which were prevalent prior to the 5RFO, the fracture mechanics analysis predicts that an initiated flaw would propagate sufficiently to result in thermal sleeve failure-in approximately eighteen months of operation. Since the thermal sleeve failure occurred about halfway through fuel cycle 5, approximately the first three fuel cycles were required to initiate the flaw under the severe makeup flow cycling conditions which dominated that period of operation.

After the thermal sleeve failure was discovered during the 5RFO, makeup flow control was improved to reduce both the frequency and amplitude of makeup flow cycling. Under the improved makeup flow cycling conditions typical for fuel cycles 6 and 7, the fracture mechanics analysis predicts a thermal sleeve life exceeding 20 eighteen-month fuel cycles. This prediction is based on a larger assumed initial flaw size because with the improved makeup flow control conditions, a much larger initial flaw size must be assumed for any significant flaw propagation to occur.

Considering that approximately three fuel cycles were required to initiate a flaw that would propagate under the severe makeup flow cycling conditions which existed prior to the 5RFO, it is reasonable to conclude that at least as long a period will be required to initiate a flaw which would propagate under the current improved makeup flow control conditions. The time required for flaw initiation provides added conservatism to the predicted thermal sleeve lifetime of 20 fuel cycles.

The thermal sleeve currently in makeup service was installed in the 5RFO, but has only one fuel cycle of makeup service. The makeup flow path was re-routed to the current nozzle during the 6RFO as one of the actions in response to the discovery of the failed thermal sleeve during the 5RFO.

With the current improved makeup flow cycling conditions, TE plans no further actions to address HPI/Makeup nozzle thermal sleeve reliability and considers the matter to be closed.

Docket Number 50-346 License Number NPF-3 Serial Number 1968 Page 4 If you have any questions regarding the information provided by this letter, please call Mr. R. W. Schrauder, Manager - Nuclear Licensing at (419) 249-2366.

Very truly yours, cc: P. M. Byron, NRC Region III, DB-1 Senior Resident Inspector A. B. Davis, Regional Administrator, NRC Region III J. B. Hopkins, NRC/NRR DB-1 Senior Project Manager Utility Radiological Safety Board

Enclosure D Davis-Besse Nuclear Power Station (DBNPS), Unit No. 1 Letter L-11-107 Babcock and Wilcox document 32-1172294-00, "Davis Besse 1 SG Flaw Evaluation," dated 6/9/1988 Babcock and Wilcox document 32-1172294-01, "Davis Besse 1 SG Flaw Evaluation," dated 7/18/1988 Babcock and Wilcox document 32-1172523-00, "DB-1 SG Flaw Evaluation," dated 7/18/1988 38 pages not including cover sheet (12 pages, 4 pages, 22 pages)

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GENERAL CALCULATIONIS, 32-1172294-00 Pmuchar Pt*.v DIMISrn

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&Jxvým&m" - "mmu& M& w" K a x S6/88 00 original Release -.

PRIPARID NY K. K. Yoon I~IZ DATE NuIS11WD by DATE P01 NO,

?

,aK*A awMX CAENERAL CALCULATIONS

..........

Doc, I,' . ....

Nuclear Pow er Divislon Davis Sense I Steam UClewato: flaw 3y'eluation -.. .

1.0 INTRODUCTION

DurinA May 1988 nipoe tlon, thare were a number of flaw indications.detacted in the steam generator shell near the steam outlet nesule (Ref, I and 2).

It was found that only ýtoindieatinor were needed for evaluation (Indcaton.000 and. #01)__, Theremaining indication. are bound.d by these two, Using the atranmeas iro RaeS, theoe M tndications were evaluated acoording to the ASHI Boiler and Pressure Vessel Code,Section XI (Ref.5),

.... . .-...

• PRPARD 1Y 00A04 AT REVIEWD .y OAT[. PAOE NO,

Muabca4 Maw

• MoOefiwm CO0PflV GENERAL CALCULATIONS.

NuchBar Power DIvivion DOC, I,.'5 " I"t¶A4 -OO.

2,0 STRESSES _

.T.h pupose of the section Is to s.ummrlz, the stres em In the vi'init, of fl aw.

Per Section 1.0# the flaw Is in the steam generator hhell near the steam outlet

.. The_.thern~ressure stresses for thtlu repion are contained tn Referencm Ma). Report 8. The stresses due to external1 loads a r given in Reference Ms~ Reoport 13. Asiwryof these stresses is given b'low.

Therml + Pressurs Stresses (KS .,_Rs 3_R.ep1ot.._.g..u-JZ) -

Pressa. Longitudinal+lc:p....Rd~

297 HU 900. -9 13 -4 15 -1 0 3,'O3Ju 9o0 -,7 -2 -1 0 7P & 90--'-----

0 --- t - 5 io - - 0

.737.... CO _ 9

-900.

.-. .- A 906 CD so 4 MI 7 -4 0 0 External Load Strsses (KS!, Ref. [33P Report 13, Page-D.5.4)

Longitudinal Hoop Radial 4.S 5.3 .8.2 83 0.0 0.0 PRIPARSIi OoAOI-AT iiYi

36Cock A WX S McD, m*n~ compw*

GENEPAL CALCULATIONS Nucear Power Division Doc, I.o.lk. 012. "00.

2., MAXIMUM TENSILE STRSS.ES The maximum tensile strosses due to therml + presoure are conserativtely conbined with the stressn due to .x.o iial loads to obtain the maximum tensile stress in the flaw regi.. The tonsile stresses are summarized below.

Maximum Tensile Stresses (KSIP TH + PR + External Load)

I ngatu44Ml Hoop Instde__ Outild Inf4d Ou4tsde

• nd ujt d

..............-.-...--. #4-----s,3 ...... 16,2 .... 23,3...... -- 0rO'- -'OO--"

2.2 MAXIJM STRESS PMANE:_

The stress range due to themal.+ prssu_re loads Ils conservatlvely__o.n.tnod with the._external load stre.es. .to obtain the maximum stress rangei,__T.he

__max t~mum__u.tr~.es~J'an*OaI~r.,ijumma ri: fzd_,bei.L.. .. . .

Mo~~l~~luinl Sti i~Il*F +~MaiRunWV + ernail -Mad)

Longitudinal Hoop Radial 17.5 19.3 20.2 27.3 1 0 2.3 PRESSURE VERSES TEDPERAKPRE-Figure 1.0 of this section shows the pressure and temperature conditlons of the .1 OTSG secondary side (shell) during the cooldown transient, This figure is taken from Reference 143. _

PVIPF0.010 NY DATS _.A lVllNWt3 11Y . . .,A OAT dl FAOI NO,. 5_+ .. ..

FT-"' 1.0 Transient N~o. IB (Cooldoim from 8% FP) _

Steam Temperature aid-Pressure During Cooldowm from 8%'FP or Reactor Trip A'ý X-:

tie 11 ... . .o.... .. . .C -..

0 Tz

-- -- - -- -- -3

, e c o r p. . .. ... - - -

-~

---

.o

..

.C

.

(t---

. v 100 LJLL I d

• 5 I Z 4 tsers 0141id late Uea 14 480 3" W Dry~am

4 Oft&. 0 1A 1 I. nA % "

Babe"* a Me. GENERAL CALCULATIONS aMc~ro, 1

• e 32-1172294-00 Nucler Pow.? D.Molon DOC. 1.1).

!

3,0 MAW EVALUATION OF INDICATION .1O, I'"D 3,1 STRESS INTEM11TY FACTQM From Ref. 1, the folloving shall and flaw geometry data are available:

. . .. ...... .... ......... .. .. .. - p

-s-_ _ _,

thickness t - 7.1 in.

flaw location - 1.5 in.,

flaw size 2a- 1_0 in.

flaw length A -' 1.0 in, a/t 0.75/7.10

__________

• I 4 4--

I

(

t°+.

eccentricity a = 3,55-1.5-s

- 1.3 in,

. Pron.Fig, A-330U...1.. of Re f,, 5j..the aspect ratio.for this flaw war. found to be outside of the range of this figure, thereafors,..a conservative value of

_Q-2.2 was jeleoted.

.__..ig._A.ý3300-2 ofRefg.j5y4ildq __

• m - 1,04 for.2*/tm, 0,2113 Prom Section 2.0 of this raport themaxi membrane teu 23._SmZ3J

_

will cover all cane.

M.F, xoon SPOI 2EvtIwEo AY i . .... . .

__ ___ _ __ _ __ _ _ __ __,_ __ _ -

, ~POS.ilOSG.3 (*4e4)

__ _ _ I  : " '" " -lt at l *. *a J Bauock t Wilcox a MtDo.tm-A company GENERAL 32-1172294-C0 CALCULATIONS Nuclear Power Divhdori Doc. 1.o...--.5.-

Sinc. " ii -Ms,*.  :

-23 x 1.04 (3.14 x0.75/ 2.2)

- 24.7 kal n.

The stree*se are for MU. and CD..

3.2 FATIGUE CRACK ORO='flAMLYSI9 .-..-.

.Snc, there, are 240 heatup and cooldo*wn ycles hiav,, s.ignificaut-:

stresses in the region under consideration, the final crack iizq af will be

• calculated by the Appendixi A of Ref. 5 procedure.

..... ..- crack The fatigue .* ..growth  ; c rate*equation . . for

. . .a .subsurface

.. .. . . ..Zl~at'

. .&v. .. . . ... -

d/N Co (A~iC)'

where da/dN

• growth rate in micro inche./cycle Co -0.0267 x 10' n -3.726

.. KI streas intensity factor range .j_ ,

For a siamplifi.d evaluation, a constant AIS in asaumed and will be ohecIe ..

after the final crack a.iLxin determined,. Also MKin_ assumed to be 24 7 kai Fn which in.larger than any atress range show.n in Section 2,0.

da/dN ,

- c(; 24 7 )n 4.128 nicro-in,/eyle..

-4.128 x 10' in. / ycis . -

,a - da/d?

• 240

- 0.00099 in.

___Since Aa is loss than one-xil,_ 1 .there in almost-no crack growth. -Therefore,.

the ftna*7 crAck sixe Le the saoe as tho initial,.rack asi, and the constant.....

P.IPARIDY~.w.. ----.-

AIVIEWID mY 7 DATE y-y-,,

DATE O.

PAO I NO. I-I

I .. PDO1103.3 -(90.4)

&a"-com* aVOICOXGENERAL 32-117229"-00 CALCULATIONSý-

Nuclear Power Division ,oD. .o. ...

J.

AXI assumption In valid. Finally, X1. -24.7 kal 3.3 ACCEPTANCE CRITERIA Figure 1 shown that the operating temperature neve'r cornes down bolcw 280F

.*less the pressure drops below 100 psi level from 900 psi. Therefore:, th-minimum operating tenrep'ature can be set at 280F.

For tho SG shell there in no appreciable fluence ou$,:lation, hence, RT - iitial RT

............ . .....- P(O_I. af.*)...... . .. _ _ .

... . .. . . .. . T . *. ' - , 2 8O - .. 0 _- . 2 2 0 T . F.. . . .. .. . .

.From Fig, A-4200-1. of sfL_3.__KI 20_-L2 0.-. . ...

.... ]*a/ ICI.-.200/-2A.._-m.8.+/- > 3.l6.m*/10- ..........................

. "Therefore, thI.hAflA4aeeaptabhl,-by -IW3*4B.-612.-

K. K. 7O0f.

PROPARED Sy OATS WiyiiID my DAE__ Q H.

S ~-.-.-" - . - -

'P01.21036.3 (p.44) "

aM &QfaMt Co* nXftCOM* a" GENERAL CALCULATIONS 12 -1,17 2 2 9 4-0 0 Nucd r Power Division uoC. 1.0.

4.0 FLAW EVALUATION OF INDICATION NO. 401 4.1 STR9S83 NTERKS1Y FACTOR From Ref. 2, the following.shell and flaw geometry data are &NAilable:

thickness t - 7.1 in.

flaw location a- 1.2 in.

flaw size 2a- 1.66 i.n.

flaw length A- 2.2 in.

a/'0.38

-a/tm 0.12 eccentricity em 3.505-1.2-a

- 1.52 in.

From Fig. A-3300-1 of Ref. 5, Fig. A-.3300-2.of Ref._ 5.yield. .....................

Mm- 1.08 for 2a/t- 0.234 From Section 2.0 of this rbporp;, the maximu membrane stress Sm -23 k ... .

will Cover all came,.

__Since~o NY m ....

t)ATI PU! Y~Ot  !AF

'IIeP~wlO IV. .. *DIE *_ OI

__ _ _ __ __ _ _ __ _ _ _

l"

/

• VliWIO .... ly . ATI* PAD N@.. . ..

4 4 PDI.11064* (9.14)

Babcock A WVlcox aMcDrmot!compeny GENERAL CALCULATIONS 32-1172294-00 MJuclear Power Division . DOC. I

• DA Id.

-_23 z1.08(3.14 x 0.63/ 1.7 *

. 30.8 kii, in. ............ ...

The strouses are for HUs sa4 CDs.

4.2 FATIGUE CRACK GROT ANALYSS Following the -m- e equation given in Section 3.2, da/dH = Co (AKI )n where da/dN* .rowth rate in micro inch.s/cycle - .

co 0,0267 x....

n - 3.726

_I - stress intensity factor range For a &imp ified evnluation a EonstantLA .is assumed and vwil be checked after the final crack size is determined. AlsoAK i asaumed to be 30.8 I k'i r"i which in larger than any stress range shown in Section 2.0, dA/d.= Co ( 3 0 ,8 )n

-- 9,394 micro-4 .n./cyrcle

.. 9..3.94. x..

-..- . in... /cycle-- - ----.. -4 As _da/dN 240 .... .......

- 0.00225 in.

af 0..083 t.0,00225 - 0.83225 . . . .........

_Since KI-ft Mm ~__ 5

-23 x 1.08 _(3.14_x 0.83225/1 .. 7.

....- 30.8 ksi Even though the crack. suie -ncreased slightly the. stress intensity. f ntor ........

_value did not, change,.__The conetant stressintensityfactor-.-anga- assumption.........

K. K. 7oon 1 PR~~r Is.4 1,i - - Ir A

" .... .. ,. " "y .. -

i, ,

... .S: .

• j,

____________PMS21O*8(b0-84) .

aMeom.. d C* gaf, m'32 -11.72294-00 McockAAlMcax GENERAL CALCULATIONS ' I.

Nuclear Powv.r Ivislen Doc. ,.6.

• __-__ _ _ _.

wan valid.

4.3 ACCEPTANCE CRITE*IA Following the seom procedure described in Section 3.3 T - RT*DT- 280 -60 - 220 F

From FPL. A-4200-1 of Ref. 5, KI, - 200 kiJ in-.

- . C KI-/ KI - 200/ 30.6 6.49 > 3,16 -. _ _

Therefore, this ilaw in acceptable by IWB-3612.

5.0 CONJCLUSION The two soele_.*d indicationa are found tn be aocn table IVy_ ;WB.3612 _,_

criteria.

6.0 FERENC*KS-

1. BW&Volumetric _Exmwination Evaluation Rp ort 88-026 TED '!ay 12. 19_8.'

2. *WVolumetric Examination Evaluation Rep ort,: TED 88-029, May 21, 1988,

3. MW Doc..ument. 'Des£in poqrt," Toledo Edison Company & C.E.I., B*

Contract No. 620-0014-55, RC Microfilm Rpý' g._ -_____ a____8_O1IG,

4. MW Documnt 018-1149327-ID "Functional GpecifIoation forR'aotor Coolant Syatow. 7 " m1traot No. 582- 751.onC.,_S,.,_.

5. ASHE Boilez and Pressure Vessel CodeSectionXI F_..Ed .tL*_'

I

L.

B01, P 20032-15 (7/Uc'i;

"* ~I e*s.e 7-,. ~a,,

FbccNk & Wilcox -,

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DOtIUMEr1T RELEA!RE PO"TICE (DRN) I crd MImm/64

.HARUL 18o__ PV B iES I J/o NO.

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SAFETY 3 DIGIT TASK NO., CLASS PUI.1W PART NO., OR B&W DOCUMENT NO, DOCUMENT TITI,[ (OMIT FOR SlAT ATVI TASK GROUPSEO. (30 Chart irs Miximum) NSSIB, (09 KY/N)t(l 9

16 &23) 32-. Ulntq*,4 -.-f, I-. m~- 4.

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kCQUIRED OISTR'BUTION , C*ORMAT I NO, COPIE'M (ORGANIZATION OR TE) DRN DO CM (INDIVIDUAL'S NAME) iRN DOC C

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(SiGNATURAT

-- I ~9-~ I -

I _ I_ I- 1Pr1NTED NAMEL27 Z6,0 4 gm tax--

WINJASEO BY (P#PARE* AR'IIEWED RY:

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..wNr ,6mp t 1(-

1 ab'Ack & Wilrox

• h t*,frtot! o* ny_

I I L* Ill

• DOCUMENT

SUMMARY

SHEET i I mI I is I I pI.f I DOCUMENT IDENTIFIER " "ZI.I 7z2,9./- o /

PREPARED BY: HEVIEWED BY:

NAtLJjJ-- HAME_._ 3-46.A"- .

SSIONATURLU!. SS TITLENAT _L'*1

.*4* -.-.. DATE___-_ TITLE S V" T'M STATEMENT:

COST CENTER-..3/Q, .REF, PAGE s.J_/. REVIEWER INDEPENDENCE PURPOSE AND

SUMMARY

OF RESULTS:

'he purpone of this calculation in to evaluate Davis Bon*s Steam Oenarator n1nw1o)1AtionN according to the ASHfC code pronedure.

The flaw indications found to be acceptAble by the ASML Section XI, 114B.3612 xtandard, Revision 1: Rpfarenoe 5 was changed to 1977 edition, Instead of 1986.

No oth,"r changes are resulted from thin correction, THE FOI LOWING COMPUTER CODES HAVE BEEN USED INTHIS DOCUMENT-Cr" /VTMION R/V m 1 0N Kv PACTF nnr

~Dt.3IO36S (944) 1 Babcock

&McD~emttw A Ww~cam Conilniy

.GENERAL CALCULATIONS' 32 -1172294 -0/

~NuclearPower Divkision DOC, 1. 0,

!C04 = 5F10~IN a11 I IUU fl~ ~!hnnwa flAk t-00 Original Release 6/20/88 01 Reference 5.revized. Lve*g 10A 6/30/88.........

!

I K.ZK. yXoI*

~1 6/3u/88 PRIPAME by-DATI 91IPWIED NY PAOF NO.I

Aabcock & Wieox a hMCOWtMal compef'y GENEARAL CALCULATIONS 32-1172294 -Of Nuclear Power Division DOC. i1o0in

__..3 ACCEPTANCE RTU .......

Volloviztg the same procedure.B described £x..eotLotn .3.3. ..-...... --

.-.-. T Ufft.-280_ý..6Q....u 220 T._........

From Fig. A-4200-1. of Ref.. 5*-.XI& 20g. kni in......... .. .

KIa./.IC. -. 200/ 30.8--. 6.49 > 3.1.6 - ~ -.

-Therefore, this flo L acceptable by ITh-36l24..'-..-........

im ... ....

- 5. 0 CONCLUS ION ... *. ... .. . .......

The two -solooted indications are found to be aeo~ptable by IW -3612

6. 0 REFERMCKS . .-.. .

1~ 34 Valwmetric H Nxaition tvyaluntiont Repoft, E 88-026, haty 12, 1988--

2.. S4W Volum.trio Examination Ilyfa uation Rport-,-TftD -6629,H-ay-21-98~

- , - MM j)ument, "Des ign~-!fteoft j0 Toletdo--Edison--Coupmny *-&.-0 R -1 0-0 8W 6 ----

.-

- - CttraeL~-No. 620-00Q1*455 --NPOD, Hicrof ilit Roll f c;- p 5f7ndtO-W-MM84 DooLment 010141*9327-00; "'Fuatioinal spn'ification .for t.

- O oolwft Rytu---3etatto 82-7151, Tolvda-Rdivan -a. 4 ~ -

PEEPARID IV,-~- DATE

BWNP 20697 (6-85)

DOCUMENT

SUMMARY

SHEET I~ a McDermott company Babcock & Wilcox DOCUMENT IDENTIFIER 3'. 1T 757-1, "bO TITLE 'E~LNW aAt&,Tt,,j PREPARED BY: REVIEWED BY:

.Dr-a Nlaur0M &P S-A NAME V7F SWeP'>

SIGNATURR. . .

TITLEDATE 40 SINTUR PrVSM4A IF 5J-' l TM STATEMENT: ..A, COST CENTER &8 REF. PAGE(S) Z2- REVIEWER INDEPENDENCE& A, ,

PURPOSE AND

SUMMARY

OF RESULTS:

Purpose The purpose of this calculation is to evaluate Davis-Besse Steam Generator flaw indications according to the ASME Code Section XI procedure.

t p, Results The flaw indications were found to be acceptable by the ASKE Section XI, IWB-3612 standard.

THE FOLLOWING COMPUTER CODES HAVE BEEN USED IN THIS DOCUMENT:

CODE / VERSION / REV CODE / VERSION / REV PAGE I O 'FZ7

BWNP-20687 (6-84)

Babcock & Wilcox

. a McDerm-ott company DOC. L.D, /?

RECORD OF REVISIONS

n. + -

00 Original Release 7/88 PREPARED BY, R. L. Carey DATE_________

REVIEWED BY DATE PAGE NO.___________

BWNP-20667 (8-.4

.Babcock..& Wilcox.

aiMcDermott company.~ 5Z 3O.10 TABLE OF CONTENTS Page 1.0 Introduction ............................ ....... .

2.0 Stresses ............ I

....... ........... . . . ..... . ...

.

Evaluation of Indication No. 200 ..............

3.0 Flwe 4.0 Flaw Evaluation of Indication No. 206 .................... IZ 5.0 Flaw Evaluation of Indication No. 208............ ....... /4, 6.0 Flaw Evaluation of Indication No. 210 .................. 17 7.0 Flaw Evaluation of Indication No. 212 ................

8.0 Conclusion .................. ................

9.0 References ............... ......... ................ ....

. Z L.-. Carey 7/8/88 PREPARED BY *L. DATE_________

REVIEWVED BY '*!L DATE 4'. PAGE NO.___________

BWNP-20687 (6-84)

-abcock & Wilcox a McDermott company ~~~~~DOC '"o . I.D

" . .-. ____

. ..... . -_______*_""-__

. . . . .- .--. _ "

1.0 Introduction During the June 1983 inspection, there were a number of flaw indications detected in the steam generator shell welds near the lower tubesheet to shell juncture (References I and 2).

It was found that five indications necessitated an evaluation (Indication #200 from Ref. 1;' Indications #206, #208, #210, #212 from Ref. 2). The remaining indications are bounded by these five.

Using the stressis from Reference 3, these five indications were evaluated according to the ASHE Boiler and Pressure Vessel Code, Section XI (Reference 5).

PREPARED BY, R.L. Carey DATE 7/8/88 REVIEWED BY ~7L)~DATE.94J..E PAGE NO. 4'

-~ BWNP-20667 (6-84) abcoCk & Wilcox.

~~*e ,a McDermott company DOC. I.0. .- /1 5Z.3 -00 2.0 Stresses The purpose of this section is to summarize the stresses in thevicinity of flaws.

Per Section 1.0, the flaws are in the steam generator shell near the lover tubesheet to shell juncture. The thermal pressure stresses for this region are contained in Reference [3], Report 4. Per page A-2 of Reference [33 the stresses in the shell near the tubesheet interface are defined by junctions 1 and 2. For the purpose of this analysis the maximum stresses from either junction will used. A summary of the stresses (P + S + free thermal) for the two junctions is given in the following tables.

JUNCTION I Thermal

• Pressure Stresses (KSI, Ref. (32, Report 4, Pages C-16-2 & C-16-10)

Transient Longitudinal Hoop Radial Iter Condition Inside Outside Inside Outside Inside Outside 801 HU0% 5 4 12 11 -1 0 803 HU 8% 3 6 13 13 -1 0 805 HU 15% 3 6 14 14 -l 0 846 ss 15% 3 6 10 10 -l 0 849 CD 8% 4 5 10 9 -l 0 851 CD 0% 4 5 9 9 -l 0 1299 DK HEAT 4 5 2 3 -1 0 868 HU 100% 5 4 9 8 -j 0 924 SS 100% 4 5 8 8 -l 0 959 CD 15% 3 6 10 10 -l 0 The transient condition comes from Reference (3], Report 4, page C-7-1.

PREPARED BY DATE REVIEWED BY DATE PAGE NO.____________

BWNP-20667 (6-84)

-Babcock & Wilcox

'*"

• * * =*-*

• a McDermott  ::

company DOC. I.D. i "'.Z5 i 3-/ . -00

.

JUNCTION 2 Thermal + Pressure Stresses (XSI, Ref. [3], Report 4, Pages C-16-4 & C-16-12)

Transient Longitudinal Hoop Radial Lter Condition Inside Outside Inside Outside Inside Outside 801 HU 0% -7 10 2 19 -1 0.

803 HU 8% -8 12 4 21 -1 0 805 -9 4 22 -1 0 HU 15% 13 846 1 12 10 -1 0 ss 15% 6 849 5 11 9 -1 0 CD 8% 5 851 CD 0% 6 5 11 9 -1 0 DK HEAT 16 -2 12 -5 -1 0 1299 868 HU 100% 5 4 9 8 -1 0 SS 100% 4 5 9 8 -1 0 924 CD 15% 3 7 11 10 -1 0 959 The transient condition comes.from Reference 133, Report 4, page C-7-1.

2.1 Maximum Tensile Stresses The maximum tensile stresses due to thermal

• pressure loads for both junction I and junction 2 are summarized belov.

Maximum Tensile Stresses.(KSI, TH + PR)

Longitudinal Hoop Radial Inside Outside Ingide Outside Inside Outside Junction I 5 6 14 14 0 0 Junction 2 16 13 12 22 0 0 PREPARED BYý& = DATE_______________

REVIEWED B'Y W DATE________ PAGE NO._________

BWNP-20667 (6-84)

Babcock & Wilcox

-,,a McDermott company. .- //7Z-"5a*-60 2.2 Maximum Stress Range The maximum stress ranges due to thermal

• pressure loads for both junctions I and 2 are summarized below.

Maximum Stress Ranges (KSI, TH + PR Range)

Longitudinal Hoop Radial Inside Outside Inside Outside Inside Outside Junction I 6 14 14 I 0 Junction 2 12 27 1 0 2.3 Pressure vs. Temperature

(

Figure 1 of this section shovs the pressure and temperature conditions of the OTSU secondary side (shell) during the cooldown transient. This figure is taken from Reference (4).

PREPARED BY rDATE*

REVIEWED BY DATE -( PAGE NO.

m FIGURE I H

Transient No. 18 (Cooldown from 8% FP)

Steam Temperature and-Pressure During Cooldown from 8%"FP or Reactor Trip Reactor Trip .0

,0-1050Pala, 5h6OF ].0 IOO Se MormalL 8% FP. Cooldovn 0 00

900 Pala, 532*F  :

>~J a3 see All "IT "l111-0

- 0 IL se48 , -

l~le0 mperaturt (&hillside later L Steam) (

• .- . ". Dry Layup 411 ~ 303 i Wet. LAyup (temp. foliova HC Temp.

0 200 90 ------- on Figure I81 diliI 1AleafJ-100 P 8 50til1 0 Z 4 6 8 I0 12 14 16 lime. Meetr

BWNP-20667 (6-84)

A Babcock & Wilcox 8McDermott company DOQ. I.O. 3-T ~ ~

3.0 Flaw Evaluation of Indication No. 200 3.1 Stress Intensity Factor From Ref. I, the following shell and flaw geometry data are available:

thickness t=6.8 in.

flaw location S=0.3 in.

flaw size a 0.25 in.

flaw length 1 1.0 in.

• From IWA-3310-1 the flaw is to be treated as a surface flaw 2ass in depth, since s is less than 2a. Instituting this change lends:

f B = 0 in. and a = 0.8 in.

a/l f 0.8 a/t 0.12 The surface flaw is shown below:

0.

R. L. Carey PREPARED RY g L'( DATE REVIEWED BY V

-DATE..L f1&*:-* PAGE NO.__ _ _

BWNP-206M1 (6-84) .

Ba.bcock A Wilcox.

aMcDermott

- company 0.17 -.

ZCC Z-' -0 From Section 2.0 of this report, the highest longitudinal stress is that of the Decay Heat transient, and the lowest longitudinal stress is g$5ok*ed +0 6e-  ;&o- The stress distribution for the Oere--. '4-q.-e4& isj5d &s 411 S.ý -

'7jl

-M5PEoI* I D For the calculation of Ki.,. these stress distributions indicate a membrane stress, a., of 7 ksl and bending stress, ab of 9 ksi. For K.,1 . the membrane stress is C ksi and the bending stress is 0 ksi.

From Figs. A-3300-1, A-3300-3, and A-3300-5 of Ref. 5, the aspect ratio and ratio of flaw depth to thickness for this flaw were found to be outside of the range of the figures, therefore, the conservative values of Q = 2.2 , H. = 1.1, and Hb=0.85 were selected.

Since K3 6 61 H, (na/0)0 + O Jib (no/Q)O'o K3 S = 7 (1.1) In (0.8)/2.21'- + 9 (0.85) [n (0.8)/2.2J105

= 16.4 kseiin.

and = 0 I1/2L47~'

K,= 16.4 -. 0 IG~.j ksi jin.

The stresses are for HUs and CDs.

3.2 Fatigue Crack Growth Analysis PREPARED BY R.L. Carey DATE 7/8/88 REVIEWED BY DATE PAGE NO._ _ _

BWNP-20667 (6-84)

Babcock & Wilcox - -

'a McDermott company .. "" -" , - -

Since there are 240 heatup and cooldown cycles which have significant stresses in the region under consideration, the final crack size, a,, will be calculated by tfie Appendix A of Ref. 5 procedure.

The fatigue crack growth rate equation for a surface flaw is:

da/dN = C. (AK),

where da/dN = growth rate in micro-inches/cycle C, = 0.3795 x l0-3 n = 3.726 K, = stress intensity factor range, ksifn.

For a simplified evaluation, a constand K, is assumed and will be checked after the final crack size is determined.

da/dH = C. (16.4)n

= 12.756 micro-in./cycle

= 12.756 x 10-6 in./cycle a a da/dN (240)

= 0.00306 in.

a, = 0.8 + 0.00306 0.80306 in.

Since K, =.q, M. (H a/9)0'0 + ab Kl (na/Q)0'=

K. 8., = 7(1.1) In (0.80306)/2.2)0-3 + 9(0.85) (n(0.80306)/2.2]0.a

= 16.4 ksi in.

Even though the flaw size increased slightly, the stress intensity factor value did not change. The constant stress intensity factor range assumption was valid.

3.3 Acceptance Criteria PREPARED BY R.L. CareyDATE 7/8/88 REVIEWED BY DATE `2  ! ~" PAGE NO. '

BWNP-20667 (6-84)

.-abcock & Wilcox.-:.

a.McDermott company ~ 17Z 3 -6o~

Figure 1 shows that operating tempprature never comes down below 280°F unless the pressure drops below the 100 psi level from 900 psi. Therefore, the minimum operating temperature can be set at 280*F.

For the SG shell there is no appreciable fluence cummulation, hence, RTDy = initial RTT

= 606F (Ref. 4)

T - RTNDT = 280-60 = 2206F From Fig. A-4200-1 of Ref. 5, KX 1 200 ksi4 in.

K:./K,..= 200/16.4 = 12.2 > 3.16 =f-Therefore, this flay is acceptable by IWB-3612 standards.

4.0 Flay Evaluation of Indication No. 206 4.1 Stress Intensity Factor From Ref. 2, the following shLell and flaw geometry data-are available:

thickness t = 6.5 in flaw location s = 0.0 in flaw size a = 0.4 in flaw length 1 = 0.9 in a/l = 0.44 a/t = 0.06 From Fig. A-3300-1 of Ref. 5, 0 = 24!

7/8/88 PREPARED BY R.L. Carey DATE REViEWED AY DATE PAGE NO.

IIJ

BWNP-20667 (6-84)

-;Babc'ock & Wilcox

.a McDermott company OC .. -

Figs. A-3300-3 and A-3300-5 of Ref. 5 yield H. : 1.1 M,= 0.95 Using the same stresses as those used in Section 3.0 of this report:

Since K, = a. K. (na/Q)O'o + a, Mi (na/9)6'=

K,.,. = 7(1.1)[ (0.4)/2.1111-5 +' 9(0.,9)i[n (0.4)/1.110-'1 X 12.6 ksii.

and Kiag,.i = Y-1L Ti-N.

Ks 12.6 0 2-12.10 ksi4ii The stresses are for HUe and CDs.

4.2 Fatigue Crack Growth Analysis Following the same equation given in Section 3.2, da/dH C. (AK,)

where da/dN = growth rate in micro-inches/cycle C, = 0.3795 x 10-n = 3.726 K, = stress intensity factor range, ksi rin.

For a simplified evaluation, a constand K, is assumed and will be checked after the final crack size is determined.

A da/dH C, (i2,.L)a

= 4.*~;micro-in.

7 /cycle

=.777 x 10o- in./cycie PREPARED BY

• R.L. Carey 7/8/88 REVIEWED BY ATE'* /-1 PAGE NO._ _ _

BWNP-20667 (6-84)

& Wilcox...-... :Babcock

ýa.McDermott -company D OC._________________________

= da/dH (240)

= 0.o00,111- in.

= 0.4 *0.001(5 0.40316, in.

5'si Since =, r. (H a/Q)0

• 5

+ Ob Mlb (ni a/g)*'5

=7(i1l)[n (0.40M:,.)/2.1)0- + 9(0.95)ln (0.40,11-55)/2.130-

= 12.6 ksifin Even though the flaw size increased slightly, the stress intensity factor value did not change. The constant stress intensity factor range assumption was valid.

4.3 Acceptance Criteria Following the same procedure described in Section 3.3, T - RTmIT = 280-60 220oF From Fig. A-4200-1 of Ref. 5, K,, = 200 ksi ýin.

K,,/K=.,. = 200/12.6 m 15.9 > 3.16 =

Therefore, this flaw is acceptable by IWB-3612 standards.

5.0 Flaw Evaluation of Indication No. 208 5.1 Stress Intensity Factor From Ref. 2, the following shell and flaw geometry data are available:

thickness t = 6.5 in.

flaw location s = 0.0 in.

R. L. CareyDATE 7/8/88 PREPARED BY REVIEWED BY -DATE '7/ glef PAGE .NO._____________

BWNP-20667 (6-84)

Babcock & Wlicox a8McDermott company.DC .. 3 flaw size a a 0.4 in.

flaw length I 1.0 1 in.

a/l = 0.4 a/t = 0.06 From Fig. A-3300-1 of Ref. 5, 0 = 1.9 Figs. A-3300-3 and A-3300-5 of Ref. 5 yield M. 1.1 Mb 0.95 Using the same stresses as those used in Section 3.0 of this report:

Since Ki = a. H, (na/Q)6's

• 06 fib (na/Q)"'

K,.,. 7(1.1)[n (0.4)/1.910-3 + 9(0.95)[E (0.4)/1.910-3 13.2 ksnii-and K,.,. r 0 K, - 13.2- 2 ISfA ksei$j The stresses are for HUs and CDs.

5.2 Fatigue Crack Growth Analysis.

Following the same equation given in Section 3.2, da/dN = C. (AK,)=

where da/dN = growth rate in micro-inches/cycle.

C. = 0.3795 x 10-3 n 3. 726

....... R.L. Carey 7/8/88 PREPARED BY 0 PAGDATE__

BWNP-20667 (6-84)

• Babcock&'Wilcox _____

a McDermott company.; - oszcu..`a o

K, - stress intensity factor range, ksi .ýJn.

For a simplified evaluation, a constand K, is assumed and will be checked after the final crack size is determined.

C. (,Z No, da/dN 5.§ micro-in. /cycle 5.61;'&x 10-' in./cycle a = da/dN (240)

= 0.0031(p in.

a, = 0.4

• 0.00 .1ý = 0.401iA in.

Since K, a. M. (n a/0)05 + a, Mi. (n a/0)0.3 KI a a 7(1.1)(n (0.40453)/1.9)0

• 9(0.95)En (0.4013(0)/1.930-5 13.2 ksi sfii Even though the flaw size increased slightly, the stress intensity factor value did not change. The constant stress intensity factor range assumption was valid.

5.3 Acceptance Criteria Following the same procedure described in Section 3.3, T - RTwr = 280-60 = 220*F From Fig. A-4200-i of Ref. 5, KI= 200 ks n Kj,/Kj,,, 200/13.2 15.2 > 3.16 =4Ti Therefore, this flaw is acceptable by IWB-3612 standards.

R.L. Carey 7/8/88 PREPARED BY )C., DATE REVIEWED BY DATE .. PAGE NO._ _ _

BWNP-20667 (6-84)

'B~abcock

fi*'j* .. &company a McDermott Wilcox.- ~ " ~ DOC

• .::*::.*..:",...: . o .D..  ;:.. ..... 7-. ......

Z 3 --C'O- -'

6.0 Flaw Evaluation of Indication Ho. 210

6. 1 Stress Intensity Factor From Ref. 2, the following shell and flaw geometry data are available:

thickness t = 6.5 in.

flaw location B = 0.0 in.

flaw size a = 0.6 in.

flaw length 1 = 0.9 in.

a/I = 0.67 a/t = 0.09 From Fig. A-3300-1 of Ref. 5, Q c-2.2 Figs. A-3300-3 and A-3300-5 of Ref. 5 yield H. 1.1 hb 0.88 Using the same stresses as those used in Section 3.0 of this report:

Since Kt t 0. H, (Ka/Q.' + a, H, (na/Q)0 K=,, 7(l.1)[nt (0.6)/2.220--l + 9(0.88)(n (0.6)/2.210-5

= 14.5 ksi and K1 ,.

K, = 14.5 - 0 = -.S ksili The stresses are for HUe and CDs.

6.2 Fatigue Crack Growth Analysis R.L. Carey 7/8/88

[L PPD PREPARED BYDAE________

REVIEWED BY DATE PAGE NO.

BWNP-20667 (684)I

!*!ii~z*); Babcock a;:

& Wilcox_:___________;_______,__

McDermott company Doc D. 3 Z- ,

Following the same equation given in Section 3.2, da/dH = C. (A)X:)

where da/dH = growth rate in micro-inches/cycle C, = 0.3795 x 10-3 n = 3.726 6 K, = stress intensity factor range, ksi in.

For a simplified evaluation, a constand K, is assumed and will be checked after the final crack size is determined.

\da/dN C. (14. 5).

S.*

= micro-in. /oycle

=z. x 10-t in./cyole a da/dK (240)

=0.0011'. in.

a, = 0.6 + 0.00101p 0.60193 in.

Since K, (I. M. (n a/Q)0'3 + o, No (i a/Q)0'3 Ki a 77(1.,1)[n (0. 60H10/2. 2)*.- + 9(0. 88)[n (0. 601ý1 )/ 2. 2]0-5 14.5 ksi Even though the f a w size increased slightly, the stress intensity factor value did not change. The constant stress intensity factor range assumption was valid.

6.3 Acceptance Criteria Following the same procedure described in Section 3.3, PREPARED BY R.L. CareyDAT 7/8/88 REVIEWED BY,' DATE 7 '  ?'PAGE NO._________

BWNP-20667 (6-84)

".. -McDermott company . 0.D . ".0. 2- - ?3 ...

T - RTm, 7 280-60 = 220"F From Fig. A-4200-1 of Ref. 5, K1 . 200 ksefi~

200/14.5 = 13.9 > 3.16 , F1 Therefore, this flaw is acceptable by IWB-3612 standards.

.kO Flaw Evaluation .of Indication No. 212

7. 1 Stress Intensity Factor I 2 om Ref. 2, P*

thickness the following shell and flaw geometry data are availablei t = 6.5 in.

flaw location s r 0.0 in.

flaw size a = 0.5 in.

flaw length I = 1.4 in.

a/1 = 0.36 a/t = 0.08 Prom Fig. A-3300-1 of Ref. 5, 1.8 pigs. A-3300-3 and A-3300-5 of Ref. 5 yield t(. = 1.1L

- Il 0.95 Using the same stresses as those used in Section 3.0 of. this report:

Since K, = M. (na/Q)- + a, Mt (na/T)W S 1.2 . . + (.95)n (0.5)/1.83

=15. 2 ksiin 1ý'R"ýPAR:ED BY SR.L. Carey 7/8/88

_DATE

____ AE O_____ -.

K!*3 DATE*

OWNP-20687 (864)

/0 a cem omPaY. -OOID and K 3,. 0 15.2 - (1 = .r ksi Win.

A X, The stregoeas are for HUa and CJs, 7.2 Fatigue Crack Growth Analysis' Following the same equation given in Section 3.2, da/dN n C, (,&X )1 where da/dH

• growth rate in micro-inohes/oyole C,

• 0.3795 x 10"3

= 3.726 Y1'f

• stress intensity factor range, knl in, For a simplified evaluation, a conatand K, is assumed and- will be checked after the final crack size is determined, da/dH C. (1i4,011-

'.(lI rmicro-in./oyale x 1O"6 in. /cyole 6a da/dN (240)

O.O001 in.

a1 0.5 + 0.O0,2ýl = 0.50N1 in.

Since xl g8 H.~ ( n a/a1)6- + (yb 14h 01aQO K,

r15.2 kaI in.

PREPARED B RL, Carey DATE 7/8/BB REVIEWED BY, 0't DATE PAGE NO. .... _ __

BWNP-M0687 (844)

-vomiah i~Y7) -a bacock& Wilcox-'~"

2 Even though the flaw size increased slightly, the stress intensity factor value.

did not change. The constant stress intensity factor range. assumption, was valid.

7.3 Acceptance Criteria Following the same procedure described in Section 3.3, T - RTNoT =280-60 r 2206F From Fig. A-4200-1 of Ref. 5, K,, r 200 ksi n, Kj, /Kt... =200/15.2 s 13.2 > 3,16 -1 Therefore, this flaw is acceptable by IWB-3612 standards..

8.0 Conolusion The five selected indications are found to be anceptable by IWB-3612 criteria.

PREPARED BY. £ce R. ,,Carey _ DATE_

REVIE BY IED .DATE PAGE NO. a

OWNP-2o0e? (644)

/.- ý,SbOock A&Wilcox.

i.M900rMolt bompany.

. . IJ~~OCIAD .3b 3 1

_m

~1 9.0 References I. O&W Volumetrio Examination Evaluation Report, TED 88-045, June.16, 1988.

2. O&W Volumetrio Examination Evaluation Report, TED 88-044, June 15, 1988,;

3. B&W Document, "Design Report,' Toledo Edison Company & C.E.I., B&W Contract No. 620-0014-55, NPGD Hicrofilm Roll Ho. 80-15 and 80-16,

4. O&W Document No. 18-1149327-00, "Functional Specification for Reactor Coolant System," Contract No. 582-7151, Toledo Edison Co., NSSE4.

5. ASME Boiler and Pressure Vessel Cdde, Section XI, 1977 Edition, PREPAtIM13Y ej R. L,# Cnrey DATE 7/8/88 REVIEWED BY_.-. DATE WO If PAGE NO. ________

_._j