Attachment 2 - Flaw Fracture Mechanics Evaluation to ...

Enclosure

Attachment 2 - Flaw Fracture Mechanics Evaluation to Support Restart

Attachment 2 -Flaw Fracture Mechanics Evaluation To Support Restart 0402-01-FOl (Rev. 017, 11/19/12)

A CALCULATION SUMMARY SHEET (CSS)AREVA

Document No. 32 - 9212942 - 001 Safety Related: H Yes 0 No

Title Palo Verde Unit 3 BMI Nozzle Repair - Section Xl Analysis for Restart

PURPOSE AND SUMMARY OF RESULTS:

Purpose

The purpose of the present fracture mechanics analysis is to determine the suitability of leaving degraded J-groove weld and butter material in the Palo Verde Nuclear Generation Station, Unit 3 (PVNGS3) reactor vesselbottom head (RVBH) following the repair of bottom mounted instrument (BMI) nozzle #3. It is postulated that asmall flaw in the head could combine with a large stress corrosion crack in the weld and butter to form a radialcorner flaw that could only propagate into the low alloy steel head by fatigue crack growth under cyclic loadingconditions.

The purpose of revision 001 is to incorporate customer comments and address the removal of "boat" sample fromremnant J-groove weld.

Summary of Results

Based on a combination of linear elastic and elastic-plastic fracture mechanics analysis of a postulated remainingflaw in the original Alloy 182 J-groove weld and butter material, bottom mounted instrument (BMI) nozzle #3 in thePalo Verde Nuclear Generation Station, Unit 3 reactor vessel bottom head is considered to be acceptable for onefuel cycle. Using EPFM analysis with safety factors of 3 on primary loads and 1.5 on secondary loads, it has beenshown that the applied tearing modulus (17.508) is less than the material tearing modulus at instability (26.580).Furthermore, with safety factors of 1.5 on primary loads and 1.0 on secondary loads the applied J-integral (0.953kips/in) is less than the J-integral of the low alloy steel head material (2.701 kips/in) at a crack extension of 0.1inch.

THE DOCUMENT CONTAINSASSUMPTIONS THAT SHALL BE

THE FOLLOWING COMPUTER CODES HAVE BEEN USED IN THIS DOCUMENT: VERIFIED PRIOR TO USE

CODE/VERSION/REV CODEN.ERSION/REV l YESANSYS 14.0 Z NO

Page 1 of 38

Attachment 2 -Flaw Fracture Mechanics Evaluation To Support Restart

0402-01-FO01 (Rev. 017, 11/19/12)P.. 1! : W.. .Document No. 32-9212942-001

Palo Verde Unit 3 BMI Nozzle Repair- Section Xl Analysis for Restart

Review Method: j[^ Design Review (Detailed Check)

D Alternate Calculation

Signature Block

P/RIAName and Title and Pages/Sections

(printed or typed) Signature LPILR Date Prepared/Reviewed/Approved

Samer MahmoudPrincipal Enginee, LP .1-2oI3 All Except Appendix A

Silvesler Noronha

Principal Engineer LR 11 -6-1OI3 All Except Appendix A

Jasmine Coo o 2-o 1;Principal Engineer Appendix A

Martin Kolar , G .•N O,, jEngineer IV R Appendix A

Tim Wiger A 6)• " All

Unit Manager A

Note: P/R/A 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

Maya Chaid rashekhar Project Manager (

Mentoring Information (not required per 0402-01)

Page 2


0402-01-FO1 (Rev. 017, 11/19/12)

Document No. 32-9212942-001A

AREVA

Palo Verde Unit 3 BMI Nozzle Repair - Section Xl Analysis for Restart

Record of Revision

Revision Pages/Sections/ParagraphsNo. Changed Brief Description / Change'Authorization

000 All Original Release001 Page 1 Added purpose for revision and updated results (minor

change to results)

Section 1.0 Page 8 Added discussion of "boat" sample removal

Section 3.2/Page 15 Corrected typo (butte to butter)

Section 5.2/Page 21 Updated computer files

Page 26 Corrected (2150 to 2235) and updated fonts for all equations

Page 27-28 Updated results (only minor change to the results)

Section 8 / Page 31 Updated Reference 1 and added reference to Boat sampledrawing

Appendix A Changed Page numbering to be sequential with the rest of thedocument

t t

Page 3


AARE VA Document No. 32-9212942-001

Palo Verde Unit 3 BMI Nozzle Repair - Section XI Analysis for Restart

Table of Contents

Page

SIG NATURE BLO CK ............................................................................................................................. 2

RECO RD O F REVISIO N ....................................................................................................................... 3

LIST O F TABLES ................................................................................................................................ 6

LIST O F FIG URES ................................................................................................................................ 7

1.0 INTRO DUCTIO N ......................................................................................................................... 8

2.0 ANALYTICAL M ETHO DO LOGY .............................................................................................. 9

2.1 Stress Intensity Factor Solution ................................................................................................. 10

2.2 Plastic Zone Correction .................................................................................................................. 11

2.3 Linear Elastic Fracture Mechanics ............................................................................................. 12

2.4 Elastic-Plastic Fracture Mechanics ........................................................................................... 12

2.4.1 Screening Criteria ....................................................................................................... 12

2.4.2 Flaw Stability and Crack Driving Force ........................................................................ 13

2.5 Sources of Stresses ....................................................................................................................... 14

3.0 ASSUM PTIO NS ....................................................................................................................... 15

3.1 Unverified Assumptions .................................................................................................................. 15

3.2 Justified Assumptions ..................... ........ ............ ......... 15

3.3 Modeling Simplifications ............................................................................................................ 15

4.0 DESIG N INPUTS ...................................................................................................................... 16

4 .1 M a te ria ls ......................................................................................................................................... 16

4.1.1 Yield Strength .................................................................................................................. 16

4.1.2 Reference Temperature .............................................................................................. 16

4.1.3 Fracture Toughness ..................................................................................................... 16

4.1.4 J-integral Resistance Curve ......................................................................................... 17

4.2 Basic Geometry .............................................................................................................................. 19

4.3 Operating Conditions ...................................................................................................................... 20

4.4 Applied Stresses ............................................................................................................................ 20

5.0 CO M PUTER USAG E ............................................................................................................... 21

Page 4

AAREVA


Document No. 32-9212942-001


Table of Contents(continued)

5.1

5.2

6.0 FLAW

6.1

6.2

6.3

Page

Hardware/Software ......................................................................................................................... 21

C o m p ute r F ile s ............................................................................................................................... 2 1

EVALUATIO N ................................................................................................................ 22

L E F M E va lu a tio n ............................................................................................................................ 2 2

EPFM Evaluation ............................................................................................................................ 26

Primary Stress Check ..................................................................................................................... 29

7.0 SUMMARY OF RESULTS AND CONCLUSIONS ................................................................ 29

7.1 Summary of Results ....................................................................................................................... 29

7 .2 C o n c lu s io n ...................................................................................................................................... 3 0

8.0 REFERENCES ......................................................................................................................... 31

APPENDIX A : COOLDOW N STRESS ANALYSIS ...................................................................................... 32

Page 5


AAREVA Document No. 32-9212942-001


List of Tables

Page

Table 1-1: Safety Factors for Flaw Acceptance ............................................................................... 9Table 4-1: M aterial D esignation ...................................................................................................... 16T a b le 4 -2 : G e o m etry ........................................................................................................................... 19Table 4-3: A pplied T ransients ........................................................................................................ 21

Table 6-1: LEFM Evaluation of BMI Nozzle Corner Crack for Heatup/Cooldown ............................. 23Table 6-2: EPFM Evaluation of BMI Nozzle Corner Crack for Heatup/Cooldown ........................... 27

Table A -I: M aterial Properties ........................................................................................................ 33Table A-2: Reactor Coolant Temperature during Cooldown Transient ............................................ 33

Table A-3: Maximum Thermal Stresses in Lower Head during Cooldown Transient ....................... 35

T able A -4 : C om puter F iles .................................................................................................................. 37

Page 6

AAREVA


Document No. 32-9212942-001


Figure 2-1:

Figure 2-2:

Figure 4-1:

Figure 4-2:

Figure 4-3:

Figure 6-1:

Figure A-1:

Figure A-2:

Figure A-3:

Figure A-4:

Figure A-5:

List of Figures

Page

Postulated Flaw in the J-groove Weld ........................................................................... 10

Schematics of Nozzle Corner Flaw Used in SIF Solution ............................................. 11

Correlation of Coefficient, C, of Power Law with Charpy V-Notch Upper Shelf Energy ...... 18

Correlation of Exponent, m, of Power Law with Coefficient, C, and Flow Stress, Yo ..... 18

Sketch showing the geometric parameters .................................................................... 19

J-T D ia g ra m ...................................................................................................................... 2 8

Finite Element Model, Boundary Condition (Left) and Temperature field (Right) ........... 34

Temperature vs. Time (Left) and Temperature Difference vs. Time (Right) .................. 34

Thermal Stress in Radial (Left) and Hoop (Right) Directions ......................................... 35

Thermal Stress in Radial (Left) and Hoop (Right) Directions vs. Depth from ID to OD ...... 36

Temperature vs. Depth from ID to OD ........................................................................... 36

Page 7




1.0 INTRODUCTION

In 2013, inspection of the Alloy 600 Bottom Mounted Instrument (BMI) nozzles on the outside surface of thereactor vessel bottom head (RVBH) at Palo Verde Nuclear Generation Station, Unit 3 (PVNGS3), identified areasof potential reactor coolant (RC) leakage between the interface of the RVBH penetration(s) and the BMInozzle(s). Arizona Public Service (APS) Company (Owner) has contracted AREVA to develop and implement amodification for the BMI nozzle penetration(s) at PVNGS3.

The repair activity as described in Reference [1] will consist of removing the existing in-core instrument (ICI)guide tube to BMI weld, removing a portion of the existing BMI nozzle below the RVBH, the machineapplication of an Alloy 52M/52MSS temper bead weld pad on the outer surface of the RVBH around thepenetration, replacing the portion of BMI nozzle removed with an Alloy 690 nozzle, attaching it to the weld padwith a J-groove partial penetration weld, and attaching the ICI guide tube with a socket weld using Alloy52M/52MSS weld filler metal.

The present concem is that a flaw in the remnant J-groove weld could impact the structural integrity of the vessel.A flaw in the weld metal may propagate into the low alloy steel by fatigue. Since the hoop stress in the J-grooveweld is greater than the axial stress at the same location, the preferential direction for cracking is radial relative tothe nozzle. It is postulated that a radial crack in the Alloy 82/182 weld metal could propagate by PWSCC, throughthe weld and butter, to the interface with the low alloy steel head material, where it is fully expected that such acrack would then blunt, or arrest, as discussed in Reference [2]. Although primary water stress corrosion crackingwould not extend into the head, it is further postulated that a small fatigue initiated flaw forms in the low alloysteel head and combines with the stress corrosion crack in the weld to form a large radial comer flaw that couldpropagate into the head by fatigue crack growth under cyclic loading conditions. Linear-elastic (LEFM) andelastic-plastic (EPFM) fracture mechanics procedures are utilized to evaluate this worst case flaw in the original J-groove weld and butter. Since the current analysis considers one fuel cycle, fatigue crack growth will be minimal.

The boat sample removed from the BMI nozzle and J-groove weld, which is described in Reference [3], has noadverse impact on the flaw evaluation conducted in this document. Obtaining the boat sample involves removinga portion of the degraded J-groove weld and nozzle. From a fracture mechanics perspective, removing a portionof the degraded J-groove weld could have a twofold impact on the structural integrity of the RVBH. From a flawsize view point, the postulated flaw size in the vicinity of the boat sample would be smaller and thereforebeneficial. On the other hand, removing the boat sample would lead to a slight redistribution of residual andoperating stresses. The impact of any residual stress redistribution on the flaw evaluation is expected to be veryminimal from flaw growth consideration during one fuel cycle. Also, any impact on operational stresses will beinsignificant. In summary, removal of the boat sample from the BMI nozzle and J-groove weld will not alter theresults and conclusions of the current flaw evaluation.

The flaw evaluation of the as-left J-groove weld postulates a large planar flaw at the inside comer of the head, atthe location of the partial penetration weld between the nozzle and the head. The comer flaw model used togenerate the crack tip stress intensity factor is based on free surfaces along each "leg" of the comer, along theinside surface of the head as well as along the interface with the nozzle. These free surfaces maximize the crackopening displacement of the postulated flaw. Consideration of the nozzle would only serve to restrict the crackopening displacement of the comer flaw in the head, and thereby reduce the calculated stress intensity factors.Furthermore, even if there was any partial through-wall cracking of the nozzle, the uncracked volume of thenozzle would still provide some degree of restraint for the postulated flaw. It is therefore appropriate to use thepostulated comer flaw model to calculate stress intensity factors for the as-left J-groove weld flaw evaluation

The purpose of this calculation is to provide a Section XI [9] analysis for restart of PVNGS3.

Page 8

AAREVA


Document No. 32-9212942-001


Key features of the fracture mechanics analysis are:

" This analysis applies specifically to BMI nozzle penetration #3 in the PVNGS3 RVBH.

* Flaw acceptance is based on linear elastic fracture mechanics (LEFM) acceptance criteria of IWB-3612 ofSection XI of the ASME Code [9] considering the safety factors listed in Table 1-1.

* In the event that LEFM margins based on IWB-3612 of Section XI of the ASME Code [9] are not met,final flaw acceptance is based on elastic plastic fracture mechanics (EPFM) methodology considering theductile tearing resistance of the reactor vessel bottom head material and the safety factors listed in Table1-1. Table 1-1 lists the EPFM safety factors that were used in the current evaluation. In addition, Table1-1 lists the safety factors that are recommended in ASME Code Case N-749 [4]. Note that for the EPFMflaw evaluation, the current analysis used safety factors that are higher than the safety factors recommendby the ASME Code Case N-749 [4], which provides a significant degree of conservatism to the analysis.

Table 1-1: Safety Factors for Flaw Acceptance

Linear-Elastic Fracture Mechanics

Operating Condition Evaluation Method Fracture Toughness / K,

Normal/Upset Kia fracture toughness 410 = 3.16

Elastic-Plastic Fracture Mechanics

Operating Condition Evaluation Method Primary Secondary

Normal/Upset J/T based flaw stability 3.0 (2.14+) 1.5 (1.0')

Normal/Upset J0.1limited flaw extension 1.5 (1.5#) 1.0 (1.0+)

ý Safety factors based on ASME Code Case N-749 [4]. These safety factors are listed here forinformation only. The higher safety factors were used in the EPFM flaw evaluation analysis.

2.0 ANALYTICAL METHODOLOGY

A radial flaw at the inside comer of the nozzle penetration is evaluated based on linear elastic fracture mechanics(LEFM) and elastic-plastic fracture mechanics (EPFM), as outlined below.

1. Postulate a flaw in the J-groove weld, radial with respect to nozzle axis extending from the inside comerof the penetration to the interface between the J-groove weld and reactor vessel shell, as shown in Figure2-1 for flaw propagation in axial and meridional directions. The postulated flaw is shown Figure 2-1.

2. Obtain operating and residual stresses need for evaluating the stress intensity factor (SIF).

3. Calculate SIF based on the closed form solution provided in Reference [5] for a nozzle comer flaw.

4. Utilize the screening criteria of ASME Code Section XI, Appendix C to determine the failure mode andappropriate method of analysis (LEFM, EPFM, or limit load) for flaws in ferritic materials, considering

Page 9

AAREVA


Document No. 32-9212942-001


the applied stress, temperature, and material toughness. For LEFM flaw evaluations, compare the stressintensity factor at the final flaw size to the available fracture toughness, with appropriate safety factors.When the material is more ductile and EPFM is the appropriate analysis method, evaluate flaw stabilityand crack driving force.

Nozzle

U-

RVBH

IFigure 2-1: Postulated Flaw in the J-groove Weld

2.1 Stress Intensity Factor Solution

Section G of Reference [5], summarizes the stress intensity factor (SIF) solution that is applicable for a surfaceflaw in the blend radius of a nozzle, see Figure 2-2 for illustration of flaw model. Section G-2.2 of Reference [5]provides the stress intensity factor for a polynomial stress distribution for a circular crack. For a stress distributionof the form

a(x) = A0 + A1 x +A2 x2 + A3 X

where x is the distance from the inside comer (Figure 2-2), the stress intensity factor is given by Equation (G-2.2)of Reference [5] (corrected). The stress intensity factor solution for the nozzle comer flaw model is:

K, = x--na2x 0.706xA +0.537x2xaxA,+0.448xa xA2+0.393x4xa.2 3 7 xA 3K1 rxOOrA+ 57 2 37r

Page 10


Document No. 32-9212942-001A

AREVA


where a is the crack depth measured from the nozzle comer.

For any arbitrary stress distribution, it is necessary to .determine the constants A0 , A,, A2. A 3 by fitting a cubicpolynomial to the stress. Knowing the coefficients, the K value for a crack depth, a, can be determined from theabove equation. The above equation can be modified to included crack face pressure as:

I2xa a2KI = 1- xa x [0.706 x (AD + AP) + 0.537 x 2 x A, + 0.448 x - x A2 + 0.393 xL2

4 x a---3 x I33r3ir x A3

where Ap is the applied pressure.

Figure 2-2: Schematics of Nozzle Corner Flaw Used in SIF Solution

2.2 Plastic Zone Correction

The Irwin plasticity correction is used to account for a moderate amount of yielding at the crack tip. Theformulation of Irwin plasticity correction is given in Reference [6]. For plane strain conditions, this correction isgiven by

2ry= 6 KI(a) J

Page 11

AAREVA


Document No. 32-9212942-001


where KI(a)= stress intensity factor based on the actual crack size, a

Ty= material yield strength.

A stress intensity factor, Kx(aj), is then calculated for an effective crack size,

a,= a + ry,

2.3 Linear Elastic Fracture Mechanics

Section XI, Article IWB-3612 [9] requires that the applied stress intensity factor, K, at the final flaw size be lessthan the available fracture toughness at the crack tip temperature, with appropriate safety factors, as outlinedbelow.

Normal Conditions: K, < Kia / '/10

where KI, is the fracture toughness based on crack arrest.

Faulted Conditions: K, < Kic / 4I2

where K1 c is the fracture toughness based on crack initiation.

2.4 Elastic-Plastic Fracture Mechanics

Elastic-plastic fracture mechanics (EPFM) will be used as alternative acceptance criteria when the flaw relatedfailure mechanism is unstable ductile tearing. This type of failure falls between rapid, non-ductile crack extensionand plastic collapse. Linear elastic fracture mechanics (LEFM) would be used to assess the potential for non-ductile failure.

2.4.1 Screening Criteria

Screening criteria for determining failure modes in ferritic materials may be found in Appendix C of Section XI.Although Appendix C, Article C-4221 [9] contains specific rules for evaluating flaws in Class 1 ferritic piping, itsscreening criteria may be adapted to other ferritic components, such as the reactor vessel, as follows:

Let, Kr' =Klpp / Kic

Sr' = Umax / a'f

Then the appropriate method of analysis is determined by the following limits:

LEFM Regime:

EPFM Regime:

K,' / Sr' > 1.8

1.8 > Kr' / Sr' >0.2

Page 12


AARE VA Document No. 32-9212942-001


Limit Load Regime: 0.2 > Kr' / Sr'

2.4.2 Flaw Stability and Crack Driving Force

Elastic-plastic fracture mechanics analysis will be performed using a J-integral/tearing modulus (J-T) diagram toevaluate flaw stability under ductile tearing, where J is either the applied (Japp) or the material (Jmat) J-integral, andT is the tearing modulus, defined as (E/yf2) (dJ/da). The crack driving force, as measured by Jpp, is also checkedagainst the J-R curve at a crack extension of 0.1 inch (JO 1). Consistent with industry practice for the evaluation offlaws in partial penetration welds used to attach nozzles to vessels, different safety factors will be utilized forprimary and secondary loads. Flaw stability assessments for normal and upset conditions will consider a safetyfactor of 3 on the stress intensity factor due to primary (pressure) stresses and a safety factor of 1.5 for secondary(residual plus thermal) stresses. The crack driving force will be calculated using safety factors of 1.5 and I forprimary and secondary stresses, respectively.

The general methodology for performing an EPFM analyses is outlined below.

Let E' = E/(1-v 2)

Final flaw depth = a

Total applied K1 = KIapp

K, due to pressure (primary) = Kip

K, due to residual plus thermal (secondary) = Kis = Kjapp- Kip

Safety factor on primary loads = SFp

Safety factor on secondary loads = SF,

For small scale yielding at the crack tip, a plastic zone correction is used to calculate an effective flaw depth basedon

a,=a + [1/(67r)] [ (Kip+ K1,) / ay]2,

which is used to update the stress intensity factors based on

K'Ip = K1p(ae)

and K'1 s= Kis(a,).

The applied J-integral is then calculated using the relationship

Japp = (SFp*K'lp + SF,*K' S2/E'.

The final parameter needed to construct the J-T diagram is the tearing modulus. The applied tearing modulus,T~pp, is calculated by numerical differentiation for small increments of crack size (da) about the final crack size(a), according to

Page 13

AAREVA


Document No. 32-9212942-001


SJapp(a + da) - Japp(a - da))Tapp = 2KO 2 da )

Using the power law expression for the J-R curve,

JR = C(Aa)m,

the material tearing modulus, Tm,, can be expressed as

Tmat = (E/oy) Cm(Aa)m-i.

Constructing the J-T diagram,

J

UnstableRegionTapp a Tmat

Applied

Instability/- Point

MaterialStableRegionTapp -Tmat

T

flaw stability is demonstrated at an applied J-integral when the applied tearing modulus is less than the materialtearing modulus. Alternately, the applied J-integral is less than the J-integral at the point of instability.

To complete the EPFM analysis, it must be shown that the applied J-integral is less than J0.1, demonstrating thatthe crack driving force falls below the J-R curve at a crack extension of 0.1 inch.

2.5 Sources of Stresses

Pertinent stresses that contribute to the crack driving force are attributed to pressure, thermal, and residualstresses. Reference [7] performed a finite element analysis specific to PVNGS3 RVBH that simulated the J-groove weld fabrication and operating stress. The stresses in Reference [7] are attributed to welding residual stress

Page 14

Attachment 2 -Flaw Fracture Mechanics Evaluation To Support RestartA

ARE VA Document No. 32-9212942-001


plus stress due to operating pressure and temperature. The stresses reported in Reference [71 can be used todetermine the stress intensity factor at steady state conditions. From flaw stability perspective, it is also desirableto obtain the stresses during cooldown conditions since cooldown produces additional tensile stress. A simplifiedfinite element analysis is provided in Appendix A to estimate thermal stresses during cooldown conditions. Thesestresses are conservatively combined with the stresses from Reference [7] to provide a source of stress forevaluating cooldown.

3.0 ASSUMPTIONS

This section discusses assumptions and modeling simplifications applicable to the present evaluation of thePVNGS3 BMI nozzle remnant J-groove weld flaw.

3.1 Unverified Assumptions

There are no assumptions that must be verified before the present analysis can be used to support theinstrumentation nozzle repair at PVNGS3.

3.2 Justified Assumptions

" The size of the J-groove weld prep is based on the dimensions depicted in Reference [1 1]. Since withinthe RVBH shell, only the J-groove weld and butter are susceptible to SCC, it is assumed that thepostulated flaw extends through the entire J-groove weld and butter.

* A crack extension of 0.04" is conservatively assumed to account for crack extension during one fuelcycle. This assumption is based on the calculation provided in Reference [12] for South Texas ProjectUnit 1 (STP-1) RVBH following a half-nozzle repair of a bottom mounted instrumentation (BMI) nozzle.The crack growth for 50 years of operation calculated in Reference [12] is 0.305". Based on the resultsfrom Reference [12] crack growth for one fuel cycle is calculated as 0.009". The current analysisconservatively used 0.04" of crack extension for one fuel cycle.

3.3 Modeling Simplifications

The operating plus residual stresses extracted from Reference [7] do not include the repair modification. This isdeemed to be an appropriate modeling simplification since the repair weld (at the OD of the RVBH) issufficiently separated from the region of interest that is adjacent to the degraded J-groove weld, which is locatednear the ID of the RVBH.

Page 15

AAREVA


Document No. 32-9212942-001


4.0 DESIGN INPUTS

This section provides basic input data needed to perform flaw evaluation of the postulated flaw in the remnant J-groove weld.

4.1 MaterialsMaterial designations for the PVNGS3 RVBH from References [1, 11] are listed Table 4-1.

Table 4-1: Material Designation

Item MaterialRVBH material SA-533 Gr. B C1.1 [11]

Cladding mat. Stainless Steel [11 ]

Original nozzle SB-166, Alloy 600 [11]

Original J-groove weld and Alloy 182 [1]buttering I

4.1.1 Yield StrengthFrom the ASME Code [8] the specified minimum yield strength for the head material is 50.0 ksi at 100 'F and42.0 ksi at 600 'F. The yield strength value at the operating temperature (565 'F) is interpolated from Reference[8] to be 42.4 ksi.

4.1.2 Reference Temperature

The RTNDT value for the PVNGS3 RVBH is reported in Reference [1] to be -60 'F. The Charpy V-notch upper-shelf energy correlation for the J-integral resistance curve with a Charpy V-notch upper-shelf energy of 119 ft-lbs[1].

4.1.3 Fracture Toughness

From Article A-4200 of Section XI [9], the lower bound Kla fracture toughness for crack arrest can be expressedas

Kla = 26.8 + 12.445 exp [0.0 145 (T - RTNDT)],

where T is the crack tip temperature, RTNDT is the reference nil-ductility temperature of the material, KI, is in unitsof ksi'iin, and T and RTNDT are in units of 'F. In the present flaw evaluations, KIa is limited to a maximum valueof 200 ksix/in (upper-shelf fracture toughness). The crack arrest Kia upper shelf toughness of 200 ksi/in isachieved at T-RTNDT > 182 'F.

A higher measure of fracture toughness is provided by the Kjc fracture toughness for crack initiation,approximated in Article A-4200 of Section XI [9] by

KIc = 33.2 + 20.734 exp [0.02 (T - RTNDT)],

Page 16

AAREVA


Document No. 32-9212942-001


4.2 Basic Geometry

The dimensions used in this document are taken from References [11].pressure vessel and the instrumentation nozzle are described in Table 4-2.

Pertinent dimensions for the reactor

Table 4-2: Geometry

Item Uphill Side Downhill Side Reference. ..... .. ._(inches) (inches)

RVBH inside radius (to base metal) 93.35 93.35 [11]

RVBH thickness 6.5 (min.) 6.5 in. (min.) [11]

Cladding thickness 0.22 0.22 [11]

J-groove Depth 1.22 1.37 [11]

Nozzle Bore diameter 3.002 3.002 [11]

Horizontal distance of nozzle axis from sphere center 8.180 8.180 [11]

Figure 4-3: Sketch showing the geometric parameters

The initial flaw is postulated to extend from the inner surface of the nozzle to the interface between the butter andthe low alloy steel. The depth of the butter from the point where the toe of the weld fillet intersects the nozzle2.033" as modeled in Reference [7].

Page 19

AAREVA


Document No. 32-9212942-001


15

10

C

5

0.0 0.4 0.8 1.2 1.6

CVN/I002.0

Figure 4-1: Correlation of Coefficient, C, of PowerEnergy

Law with Charpy V-Notch Upper Shelf

0.7

0.6

0.5

0.4m

0.3

I I I

0

0

I I W

CT

0.2

TESTED ATIT-CT (OC)

120130

o 150170

o 200

II I

0.1

0 . .. . ..... , |

0 2 4 6 8 10

x= C+ 1.5( -)12 14 16

Figure 4-2: Correlation of Exponent, m, of Power Law with Coefficient, C, and FlowStress, a.

Page 18

AAREVA


Document No. 32-9212942-001


4.2 Basic Geometry

The dimensions used in this document are taken from References [11].pressure vessel and the instrumentation nozzle are described in Table 4-2.

Table 4-2: Geometry

Pertinent dimensions for the reactor

Item Uphill Side Downhill Side Reference_______________________________ (nches) (inches)______

RVBH inside radius (to base metal) 93.35 93.35 [11]

RVBH thickness 6.5 (min.) 6.5 in. (min.) [11]

Cladding thickness 0.22 0.22 [11]

J-groove Depth 1.22 1.37 [11]

Nozzle Bore diameter 3.002 3.002 [11]

Horizontal distance of nozzle axis from sphere center 8.180 8.180 [11]

1.37

Figure 4-3: Sketch showing the geometric parameters

The initial flaw is postulated to extend from the inner surface of the nozzle to the interface between the butter andthe low alloy steel. The depth of the butter from the point where the toe of the weld fillet intersects the nozzle2.033" as modeled in Reference [7].

Page 19


AREVA Document No. 32-9212942-001


Crack Propagation Consideration

The only credible mechanism for flaw growth in the low alloy steel in PWR environment is by fatigue. Becausethe intent of this document is to demonstrate acceptance of a postulated flaw in the existing J-groove weld for onefuel cycle, crack growth due to fatigue in one fuel cycle is marginal. To conservatively account for any potentialfatigue crack growth during one fuel cycle, an additional 0.04" of crack extension will be added to the depth ofthe postulated flaw, which is the depth of the existing J-groove weld and butter. Thus the total flaw depth afterone fuel cycle of operation is assumed to be 2.037" (2.033+0.04"). The 0.04" crack extension is conservativelyused based on the crack growth estimation provided in Reference [12] for South Texas Project Unit 1 (STP-1)RVBH following a half-nozzle repair of a bottom mounted instrumentation (BMI) nozzle (See Section 3.0 foradditional details).

4.3 Operating Conditions

Per Reference [1] the design pressure and temperature are 2500 psia and 650 'F, respectively. The operatingpressure and inlet temperature are 2250 psia and 564.5 'F [11] (565 'F was used).

4.4 Applied Stresses

Two sources of stress are considered for the present flaw evaluations, stresses that occur during normal operation,and residual stresses from welding.

Residual plus operating stresses are obtained from Reference [7], which performed a three-dimensional elastic-plastic finite element stress analysis that simulates the attachment of the original nozzle to RVBH. The analysisin Reference [7] includes the simulation of steady state operating conditions at operating temperature and pressureof 5657F and 2,235 psig.

As stated earlier, it is important to demonstrate flaw stability at cooldown conditions. A simplified finite elementanalysis (provided in Appendix A) was performed to estimate the thermal stresses during cooldown conditions.These stresses are conservatively combined with the stresses from Reference [7] to provide a source of stress to beused for evaluating the crack driving force during cooldown conditions.

The stresses used in the flaw evaluation are listed in Table 4-3 for several positions (x) along crack depth.

Page 20

AAREVA

Attachment 2 -

Flaw Fracture Mechanics Evaluation To Support Restart

Document No. 32-9212942-001


Table 4-3: Applied Transients

Position SS CD SS+CDx Hoop Stress

(in.) (ksi) (ksi) (ksi)0.0000 50.014 7.485 57.4990.2980 61.709 6.461 68.1700.5950 73.123 5.493 78.6160.8920 71.136 4.576 75.7121.1890 74.007 3.710 77.7171.4860 57.094 2.895 59.9891.7830 24.199 2.130 26.3292.0330 3.862 1.526 5.3882.2460 40.983 1.039 42.022

SS = Steady StateCD = Cooldown

5.0 COMPUTER USAGEThis section describes computer resources and stored computer files.

5.1 Hardware/SoftwareThe following computer resources were used in the present analysis.

1. Computer: Dell Precision Workstation - Tag#5VJV5SI2. Computer processor: Intel® CoreTM i7-2640M CPU @ 2.80 GHz3. Computer memory: 8.00 GB RAM4. Computer operating system: Microsoft Windows 7 Enterprise 2009 Service Pack 1

5.2 Computer FilesThe computer files listed below are stored in the AREVA ColdStor repository in the directory \cold\General-Access\32\32-9000000\32-9212942-00 \official\EXCELFiles".

File Name Time and Date File ColdStor Storage Date File Size ChecksumModified and Time

PV3 BMN HUCD EPFM.xlsx Oct 31 2013 15:00:28 Oct 31 2013 15:02:55 43660 46282PV3 BMN HUCD EPFM.xlsx Oct 30 2013 10:03:13 Oct 31 2013 15:02:56 248832 21664

Page 21




6.0 FLAW EVALUATION

6.1 LEFM Evaluation

Table 6-1 presents a fracture mechanics analysis wherein stress intensity factors are calculated for comparisonwith the fracture toughness requirements of Section XI. Article IWB-3612 [9] requires that a safety factor of 4I10be used when comparing the applied stress intensity factor to the fracture toughness for crack arrest. Calculationsare performed for a postulated radial comer crack in BMI nozzle head penetration.

Since the calculated fracture toughness margins are less than the Code required minimums, EPFM flawevaluations were performed in Section 6.2 to account for the ductile behavior of the low alloy steel under stablecrack propagation.

Page 22

AAREVA


Document No. 32-9212942-001


Table 6-1: LEFM Evaluation of BMI Nozzle Corner Crack for Heatup/Cooldown

INPUT DATA

Initial Flaw Size: Distance to base metal:Allowance for crack growth:

2.033 in.0.040 in.

Total depth, a = 2.073 in.

Material Data: Reference temp., RTndt = -60 OF

Upper shelf toughness =Transition temperature =

200 ksi'!in121.6 OF

Kia = 26.8 + 12.445 exp[ 0.0145 (T - RTndt )

Stresses:

Toe of the fillet

Butter/Head Interface

LoadingConditions

SS CD SS+CDTemperature (°F)

565 565 565Yield Strength

(ksi)42.4 42.4 42.4

Pressure (ksi)2.235 2.235 2.235

Kia (ksi'lin)Position 200 200.0 200.0

x Hoop Stress(in.) (ksi) (ksi) (ksi)

0.0000 50.014 7.485 57.4990.2980 61.709 6.461 68.1700.5950 73.123 5.493 78.6160.8920 71.136 4.576 75.7121.1890 74.007 3.710 77.7171.4860 57.094 2.895 59.9891.7830 24.199 2.130 26.3292.0330 3.862 1.526 5.3882.2460 40.983 1.039 42.022

SS = Steady StateCD = Cooldown

Page 23




Table 6-1: LEFM Evaluation of BMI Nozzle Corner Crack for Heatup/Cooldown(Continued)

STRESS INTENSITY FACTOR

Kl(a) = [0.706(Ao+A,) + 0.537(2a/7t)A 1 + 0.448(a 2/2)A2 + 0.393(4a 3/37r)A 3 ]

where the through-wall stress distribution is described by the third order polynomial,

S(x) = A0 + Ajx + A2x2 + A3x

3,

defined by:

LoadingStress ConditionsCoeff. SS CD SS+CD

(ksi) (ksi) (ksi)

A3 30.808 -0.002 30.806

A2 -127.324 0.296 -127.027

A1 122.528 -3.525 119.003

A0 43.826 7.485 51.311

Effective crack size:

ae = a + 1/(6rc)*[Kl(a)/Sv]2

Effective stress intensity factor:

KI(ae) = [ 0.706(Ao+Ap) + 0.537(2aeli)Aj + 0.448(ae 2/2)A 2 + 0.393(4ae3/37t)A 3 ]

Page 24

AAREVA


Document No. 32-9212942-001


Table 6-1: LEFM Evaluation of BMI Nozzle Corner Crack for Heatup/Cooldown

(Continued)

FRACTURE TOUGHNESS MARGINS

Final Flaw Size: a = 2.0730 in.

Margin = Kia / Ki(ae)

LoadingConditions

SS CD SS+CDFracture Toughness, 200 200.0 200.0

KiaKI(a) 108.63 11.86 116.46

ae 2.4212 2.0771 2.4733KI(ae) 109.44 11.86 117.37

Actual Margin 1.83 16.86 1.70

Required Margin 3.16 3.16 3.16

ksiin

ksi'in

in

ksi'in

Page 25

Attachment 2 -

Flaw Fracture Mechanics Evaluation To Support Restart



6.2 EPFM Evaluation

Because the EPFM methodology applies different safety factors to the primary (pressure) and secondarycontributions of the total stress intensity factor, is necessary to separate the total stress intensity factor to itsprimary and secondary contributions. The pressure stress intensity factor (KIp) is calculated using the closed formsolution used in this document as

_= J x [o.7o6x(A 0 + A)]

Where a is the crack depth, A0 is the hoop stress with stress concentration due to the hole, Ap is the term used toaccount for crack face pressure (pressure value is used). A0 is evaluated using the equation below.

= RmAD = SCFxP x-

2xt

Where SCF is stress concentration factor (a value of 2 is used for a hole in a flat plate under biaxial stress filed), Pis the applied pressure, Rm is the mean radius, and t is the thickness. Substituting into the above equation

96.6A0 = 2 x 2235 9 = 33215-5 psi

2 x 6.5

Substituting into Kip equation

Kip = ýrx 2073 x [0.706x (33215.5 + 2235)] = 63870-8 psi Fm

Once Kip is known then the secondary stress intensity factor (Kis) can be evaluated as

KIS = KI p

where K, is the total stress intensity factor.

Table 6-2 contains the EPFM evaluation. The J-T diagram is shown in Figure 6-1.

Page 26

AAREVA


Document No. 32-9212942-001


Table 6-2: EPFM Evaluation of BMI Nozzle Corner Crack for Heatup/Cooldown

EPFM Equations: Jrmt = CA~

Tt= (E/a 2)-CM(Aa)r"l

C = 7.68m = 0.45

Japp = [Kl'(ae)] 2/E'

Tapp = (E/cyf 2)*(dJappda)

Ductile Crack Growth Stability Criterion: Tap < Tret

At instability: Tapp --- Tmat

Safety Factors Kl*p KI s Kl*(a) ae Kl'(ae) Japp Tapp Stable?

Primary Secondary (ksiqin) (ksiin) (ksi'in) (in.) (ksi'in) (kips/in)1.00 1.00 63.870 52.592 116.462 2.4733 127.209 0.533 1.897 Yes2.00 1.00 127.740 52.592 180.332 3.0326 218.114 1.568 5.576 Yes3.00 1.50 191.610 78.887 270.497 4.2322 386.498 4.924 17.508 Yes5.00 1.00 319.350 52.592 371.942 6.1554 640.920 13.539 48.146 No7.00 1.00 447.090 52.592 499.682 9.4411 1066.361 37.479 133.278 No

Iterate on safety factor until Tapp = Tt to determine Jinstbiy:

2.6589 2.6589 169.824 139.836 309.660Jmnstability Tapp Trt

4.9027 476.213 7.475 26.580 26.580

at Jnt = 4.924 kips/in, Tra,= 43.915 ( Tapp - Tat = 0.000

Applied J-Integral Criterion: Japp J0.1

J0.1 = Jrt at a = 0.1 in.where,

Safety Factors Kl*p KIls Kl*(a)Primary Secondary (ksi'/in) (ksi'/in) (ksi'in)

a, Kl'(a.) Japp JO.1(in.) (ksi',in) (kips/in) (kips/in)

OK?

1.50 1.00 95.805 52.592 148.397 2.7229 170.073 0.953 2.701

* Note that the value of the Young's Modulus (E) at 565 TF from Reference [8] is 27610 ksi.

Yes

Page 27

AAREVA


Document No. 32-9212942-001


Figure 6-1: J-T Diagram

10

9

8

7

6

0~

0)

5

0 5 10 15 20 25 30 35 40 45 50

Tearing Modulus

Page 28


Document No. 32-9212942-001A

AREVA

Palo Verde Unit 3 BMI -Nozzle Repair - Section Xl Analysis for Restart

6.3 Primary Stress Check

The flaw acceptable criterion of IWB-3610(d)(2) requires that a the primary stress limits of NB-3000 be satisfiedassuming a local area reduction of the pressure retaining membrane that is equal to the area of the flawed material.This primary stress check can be met by satisfying the reinforcement requirements of NB-3332 for openings inshells and formed heads since these requirements provide for adequate compensation for material removed fromthe pressure boundary, in a similar fashion to the area of degraded material associated with a postulated ordetected flaw.

The Section III based justification for continued operation [13] provides a reinforcement calculation for thePVNGS3 lower head considering the enlarged opening at penetration #3. Per NB-3335 both the original J-grooveand repair weld may be counted for the area of reinforcement. NB-3336 requires the area of the reinforcement bemultiplied by the ratio of the design stress intensity (Sm) of the reinforcement material to the design stressintensity of the removed material. The repair weld material Sm is equal to the Sm for the original J-groove weldand the area of the repair J-groove weld and weld pad is obviously greater than the area of the original J-grooveweld and butter. Since the repair weld and weld pad were not considered in satisfying the reinforcementrequirements of NB-3332, the additional reinforcement provided by the repair weld and weld pad may be used toimplicitly satisfy the primary stress check of JWB-3610(d)(2).

7.0 SUMMARY OF RESULTS AND CONCLUSIONS

Linear-elastic and elastic-plastic fracture mechanics has been used to evaluate a postulated radial flaw in theremnant J-groove weld and butter of a BMI nozzle reactor vessel bottom head penetration after repair. It isdetermined that the flaw size would remain acceptable after one fuel cycle, as summarized below.

7.1 Summary of Results

Flaw SizesInitial flaw size,Assumed Flaw Extension,Final flaw size,

ai = 2.033 inAa = 0.040 inaf = 2.073 in

Operating Conditions

Temperature,Material tearing modulus,Material J-integral at 0. 1" crack extension,Safety factors (primary/secondary),Applied tearing modulus (.<.Treat)Safety factors (primary/secondary),Applied J-integral (< J0 .1)

T =565 'FTmat = 26.580Jo1 = 2.701 kips/inSF = 3 / 1.5Tapp = 17.508SF= 1.5/1

Japp = 0.953 kips/in

Page 29




7.2 Conclusion

Based on a combination of linear elastic and elastic-plastic fracture mechanics analysis of a postulated flaw in theoriginal Alloy 182 J-groove weld and butter material, the Palo Verde Nuclear Generation Station, Unit 3(PVNGS3) reactor vessel bottom head (RVBH) is considered to be acceptable for one fuel cycle following theproposed repair.

Page 30




8.0 REFERENCES

References identified with an (*) are maintained within [PVNGS3] Records System and are not retrievable fromAREVA Records Management. These are acceptable references per AREVA Administrative Procedure 0402-01,Attachment 8. See page [2] for Project Manager Approval of customer references.

1. AREVA Document 08-9212780-001, "Palo Verde Unit 3 Reactor Vessel Bottom Mounted Instrument

Nozzle Modification."

2. AREVA Document 51-5012047-00, "Stress Corrosion Cracking of Low Alloy Steel."

3. *Document N001-0301-00633, Revision 0, "Boat Sample Extraction General Layout Drawing.". OEMDoc No. BBE-2205, Rev. 0.

4. Cases of ASME Boiler and Pressure Vessel Code, Code Case N-749, "Alternative Acceptance Criteria forFlaw in Ferritic Steel Components Operating in the Upper Shelf Temperature Range," Section XI,Division 1, Approval Date: March 16, 2012.

5. Marston, T.U., "Flaw Evaluation Procedures- Background and Application of ASME Section XI,Appendix A," EPRI Report NP-719-SR, August 1978.

6. T.L. Anderson, Fracture Mechanics: Fundamentals and Applications, CRC Press, 1991.

7. *Dominion Engineering, Inc., Calculation No. C-7789-00-2, Revision No. I "Palo Verde Bottom HeadInstrumentation Nozzle Stress Analysis."

8. ASME Boiler and Pressure Vessel Code, Section III, Subsection NB, 1971 Edition, through Summer1973 Addenda.

9. ASME B&PV Code Section XI, "Rules for Inservice Inspection of Nuclear Power Plant Components",2001 Edition, including Addenda through 2003."

10. NUREG-0744, Vol. 2, Rev. 1, "Resolution of the Task A-Il Reactor Vessel Materials Toughness SafetyIssue," Appendix D, Materials Toughness Properties, Division of Safety Technology, Office of NuclearReactor Regulation, U.S. Nuclear Regulatory Commission, Washington, D.C. 20555, October 1982.

11. *Report N001-0301-00214, Revision 007, "Reactor Vessel, Unit 3, Analytical Report, V-CE-30869,

30AU84."

12 AREVA Document 32-5027942-002, "STP-1 BMI Nozzle Original J-Groove Weld Flaw Evaluation."

13. AREVA Document 32-9212915-001, "Palo Verde Unit 3 - Instrument Nozzle Repair Section III OneCycle Justification."

Page 31

AAREVA


Document No. 32-9212942-001


APPENDIX A: COOLDOWN STRESS ANALYSIS

A.1 Purpose

The purpose of the analysis in this appendix is to determine the maximum hoop thermal stress in the Palo Verdereactor vessel lower head developed during cooldown transient.

A.2 Methodology

1. Generate a 2D axisymmetric finite element model to simulate a simplified reactor vessel lower head withan inner radius of 93.3 inches and a thickness of 6.5 inches (Reference [A. ]);

2. Perform thermal transient analysis for cooldown condition to determine the temperature field of thereactor vessel lower head,

3. Get temperature field and thermal gradients for each time point;

4. Identify maximum thermal gradient across thickness and the time point of its occurrence;

5. Perform structural analysis, using temperature field identified in Step 4, to determine the thermal stressdistribution through the thickness of the head.

A.3 Assumptions

1. The finite element model represents a perfect hemisphere. Any feature other than the sphere portion ofthe base metal of the lower head, such as cladding, weld, and penetration elements are not included;

2. The fluid temperature data during cooldown transient are taken from Figure 2 of Reference [A.2]. It hasan approximately 100 °F/hr temperature drop rate;

3. The initial condition of the lower head is assumed to be a uniformly distributed temperature of 565 'F.

A.4 Material Properties

Per Reference [A.l], the material of the reactor vessel lower head is SA-533 Gr. B Class 1 (C-Mn-Mo-0.4-0.7Ni).The material properties are taken from Reference [A.3] except the material densities are taken from Reference[A.5].

Page 32

AAREVA


Document No. 32-9212942-001


Table A-i: Material Properties

Modulus of Thermal Thermal Specific Heat DensityTemp. Elasticity Expansion Conductivity (k) (C) (p)

Coefficient (u)(C()

OF × 106, psi x10-6, 1/-F Btu/hr-in-0 F Btu/lb-0 F lb/in3

100 29.80 6.13 2.5833 0.1147 0.2839

200 29.50 6.38 2.5000 0.1169 0.2831

300 29.00 6.60 2.4250 0.1210 0.2823

400 28.60 6.82 2.3417 0.1251 0.2817

500 28.00 7.02 2.2667 0.1292 0.2809

600 27.40 7.23 2.1833 0.1333 0.2802

700 26.60 7.44 2.1083 0.1393 0.2794

Reference [A.3] [A.3] [A.3] Calculated" [A.5]Note: *C = K/(p" Td), where Tais thermal diffusivity from the same source as thermal conductivity (k in the table).

A.5 Finite Element Model and Boundary Conditions and ResultsDefinition of the reactor coolant temperature history for cooldown transient is listed in Table A-2 (Reference[A.2]). The temperature data is input as bulk temperatures of the inner surface of the head in the thermal transientanalysis.

Table A-2: Reactor Coolant Temperature during Cooldown Transient

Time (hr) 0 4.8 8.0Reactor Coolant Temperature (F) 565 70 70

A convection coefficient of 1000 Btu/hr-ft2 -°F is applied on the inner surface of the base metal of the head. Thisvalue is based on experiences from similar projects performed in the past. The convection coefficient on the outersurface of the lower head is assumed to be 0.150 Btu/hr-ft2-'F and the ambient air temperature is assumed to be70'F during cooldown. The lower head is assumed to be initially under uniformly distributed temperature of5650F.

Figure A-I shows Finite element model boundary conditions and the temperature field. Figure A-2 shows thehistory of temperature vs. time and the history of temperature gradient between inside and outside surface of thehead vs. time. Note that curves identified with TEMPI, TEMP_2 and TEMP_3 in the left graph of this figure aretemperature histories for node located on inner surface, at depth of 1.5 inches from inner surface, and on outersurface. Figure A-3 shows radial and hoop thermal stresses in the lower head at the maximum temperaturedifference time point during cooldown transient (at time of 1.31343 hours). Table A-3 lists radial and hoopthermal stresses in a path across the thickness of the lower head (path is shown in Figure A-1). Figure A-4provides graphs for the thermal stresses vs. depth from ID to OD of the lower head. Figure A-5 shows thetemperature vs. depth from ID to OD.

It is seen that the maximum hoop thermal stress on the inner surface of the lower head during Cooldown transientis about 7.5 ksi.

Page 33

AAREVA


Document No. 32-9212942-001


Figure A-i: Finite Element Model, Boundary Condition (Left) and Temperature field (Right)

AN AN

%%WJ VAm

Tim TIM

Figure A-2: Temperature vs. Time (Left) and Temperature Difference vs. Time (Right)

Note: TEMPI, TEMP_2, and TEMP 3 represents locations at inner surface, 1.5 inches from the inner surface,and the outer surface of the lower head, respectively.

Page 34

AAREVA


Document No. 32-9212942-001


XB0 14.0CCT 28 201311:48: 136129 MD. 10UPAL 06WrEclO

RSS-1

OW -.:28237SN6 -. 004827

OW -190.635SXB-090 635

*21185942 36763.5481847292-105 91127.091148 272169 454190 635

NV0 14.0=rY 28 201311:48.21PLOW ND. 1N29L SOUMMl92 6-1T0B-1

ICK -. 2602375W -3521.6SN--3521.65N4 -7485.09566-7485 08

-3521 6* 2298663

-1075.67147.2971370 262593 223816.

19

50m• 156M62 127485.08

I ture pn onl Ogy L Mtue odyI

Figure A-3: Thermal Stress in Radial (Left) and Hoop (Right) Directions

Note: Thermal stresses are calculated based on temperature field at the time point, during cooldown transient,with maximum temperature difference between ID and OD of the lower head.

Table A-3: Maximum Thermal Stresses in Lower Head during Cooldown Transient

Palo Verde (Ri-93.35", Thk•-6.5") .....

Depth from ID to OD Temperatur (F) SX'* (psi) SY*(psi), SZ* (psi)0.0 433.26 2 7485

0.5 438.58 72 5797

1.0 443.46 124 4255

1.5 447.90 160 28582.0 451.90 182 1604

2.5 455.46 191 491

3.0 458.59 189 -482

3.5 461.29 177 -13174.0 463.57 158 -2015

4.5 465.42 133 -2579

5.0 466.85 103 -3009

5.5 467.86 70 -3309

6.0 468.46 35 -3479

6.5 468.66 0 -3522Note: * The stresses are under spherical coordinate system. SX represents the stress in radial direction, and SYand SZ represent the stresses in the hoop directions.

Page 35

AAREVA


Document No. 32-9212942-001


Radial and Hoop Stresses vs. Depth from IDto OD

-40.m- Palo Verde (Ri=9335", Thk=6.5") SK250 SX (Radial Stress) --- Palo Verde (Ri=93.35', Thk=6.S5) SY

201)0

r,-4160D0j

ri 1so 00

o -6oo

42000

_ _ _ _ _6000

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0 5.5 6.0 6.5

Depthfrom IDt OD (In)

Figure A-4: Thermal Stress in Radial (Left) and Hoop (Right) Directions vs. Depth from ID toOD

Note: Thermal stresses are calculated based on temperature field at the time point, during cooldown transient,with maximum temperature difference between ID and OD of the lower head.

Temperature vs. Depth from ID to OD

475.00

470.00 _

465.0_

460.00 ___ ___ ___

* 455.00

450100 /

-w- Palo Verde (Ri=93.35", Thk=6.5S) Temp. (F)445.00 ___le

440.00 /

435.00f _

430.000.0 0.S 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0 5.5 6.0 6.5

Ditance fhmn ID to OD (In)

Figure A-5: Temperature vs. Depth from ID to OD

Page 36

AAREVA


Document No. 32-9212942-001


A.6 Hardware, Software and Computer Files

A.6.1 Hardware and softwareThe EASI listed computer program ANSYS Release 14.0 (Reference [A.4]) is used in this calculation.Verification tests of similar applications are listed as follows:

" Error notices for ANSYS Release 14.0 are reviewed and none apply for this analysis." Computer hardware used:

o Dell Precision (Computer Name: MOCAO2, Service Tag #: 5VKT5S1) with Intel® CoreTM i7-2640M CPU @ 2.80GHz, 2.80 GHz, 8.00 GB of RAM and Operating System is MicrosoftWindows 7 Enterprise Version 2009 Service Pack 1.

o Name of person running tests: Jasmine Cao* Date of tests:

o October 27, 2013 on computer "MOCAO2" (Service Tag #: 5VKT5S1)* Acceptability: Results shown in files vm5.out and vm28.out show that the test runs are acceptable.

A.6.2 Computer FilesThe computer files for this evaluation are stored in the ColdStor under /cold/General-Access/32/32-9000000/32-9212942-000/official directory. The computer files are listed below:

Table A-4: Computer Files

File Name Time and Date File ColdStor Storage Date File Size ChecksumModified and Time

CD tr.inp Oct 21 2013.17:05:53 Oct 30 2013 15:18:01 1299 25823SA533_TypeB_Class2_ASME_1971.mp Oct 28 2013 11:20:45 Oct 30 2013 15:18:04 1773 50847postjpv.out Oct 28 2013 11:45:14 Oct 30 2013 15:18:01 13978 34722poststress.inp Oct 27 2013 12:47:46 Oct 30 2013 15:18:02 800 11532pv st studydiff Ri.inp Oct 28 2013 09:39:24 Oct 30 2013 15:18:02 2864 05548rpv_pv.out Oct 28 2013 11:45:08 Oct 30 2013 15:18:03 1076520 31094vm28.out Oct 27 2013 12:17:36 Oct 30 2013 15:18:04 18528 26608vm5.out Oct 27 2013 12:17:39 Oct 30 2013 15:18:05 40368 37793

Page 37




A.6.3 References for Appendix A

References identified with an (*) are maintained within [PVNGS3] Records System and are not retrievable from AREVARecords Management. These are acceptable references per AREVA Administrative Procedure 0402-01, Attachment 8. Seepage [2] for Project Manager Approval of customer references.

[A.1 ]. *Report NOO1-0301-00214, Revision 007, "Reactor Vessel, Unit 3, Analytical Report, V-CE-30869,30AU84."

[A.2]. *Customer Document, N001-0301-00006, OEM Document No. 00000-PE-1 10, Rev. 05, B3, OEM Title"General Specification for Reactor Vessel Assembly."

[A.3]. ASME Boiler and Pressure Vessel Code, Section III, Subsection NB, 1971 Edition, through Summer1973 Addenda.

[A.4]. ANSYS Finite Element Computer Code, Version 14.0, ANSYS Inc., Canonsburg, PA.

[A.5]. AREVA Document NPGD-TM-500 Rev. D, "NPGMAT, NPGD Material Properties Program, User'sManual (03/1985)"

Page 38

Date post:	14-Jan-2022
Category:	Documents
Upload:	others
View:	2 times
Download:	0 times

Attachment 2 - Flaw Fracture Mechanics Evaluation to ...

Documents