Text

Proposed Alternative to ASME Examination Requirements for Repairs Performed on Reactor Vessel Head Penetrations
	ML023600489
Person / Time
Site:	Arkansas Nuclear
Issue date:	12/16/2002
From:	; Entergy Operations
To:	; Document Control Desk, Office of Nuclear Reactor Regulation
References
	CNRO-2002-00058
	Download: ML023600489 (98)
	v • d • e

Entergy Operations, Inc.

1340 Echelon Parkway Jackson, MS 39213-8298 Tel 601 368 5758 Michael A. Krupa Director Nuclear Safety &Licensing CNRO-2002-00058 December 16, 2002 U. S. Nuclear Regulatory Commission Attn.: Document Control Desk Washington, DC 20555-0001

SUBJECT:

Entergy Operations, Inc.

Proposed Alternative to ASME Examination Requirements for Repairs Performed on Reactor Vessel Head Penetrations Arkansas Nuclear One, Unit 1 Docket No. 50-313 License No. DPR-51

REFERENCES:

1. Entergy Operations, Inc. Letter No. CNRO-2002-00054 to the NRC, "Proposed Alternative to ASME Examination Requirements for Repairs Performed on Reactor Vessel Head Penetrations,"

dated November 26, 2002

2. Entergy Operations, Inc. Letter No. CNRO-2002-00052 to the NRC, "Proposed Alternative to ASME Examination Requirements for Repairs Performed on Reactor Vessel Head Penetrations,"

dated October 28, 2002

Dear Sir or Madam:

In Reference 1, Entergy Operations, Inc. (Entergy) provided to the NRC staff four calculations performed by Framatome-ANP that supported Relief Request Nos. ANO1-R&R-003, Rev. 0 and ANOI-R&R-004, Rev. 0, which were submitted via Reference 2. As documented in Reference 1, Framatome-ANP considers information contained in Framatome Documents 32-5021538 and 32-5021539 (ANO Calculations 86-E-0074-156 and 86-E-0074-161) to be proprietary and confidential pursuant to 10 CFR 2.790(a)(4) and 10 CFR 9.17(a)(4).

In Reference 1, Entergy stated we would provide nonproprietary versions of the Framatome documents when they became available. This letter transmits those nonproprietary versions as Enclosures 1 and 2.

CNRO-2002-00058 Page 2 of 2 Should you have any questions regarding this letter, please contact Guy Davant at (601) 368-5756.

This letter contains no commitments.

Very truly yours, MAK/GHD/bal

Enclosures:

1. Framatome Document 32-5021538 (ANO Calculation 86-E-0074-156)

2. Framatome Document 32-5021539 (ANO Calculation 86-E-0074-161) cc: Mr. C. G. Anderson (ANO) (w/o)

Mr. W. R. Campbell (ECH) (w/o)

Mr. G. A. Williams (ECH) (w/o)

Mr. T. W. Alexion, NRR Project Manager (ANO-2)

Mr. R. L. Bywater, NRC Senior Resident Inspector (ANO) (w/o)

Mr. E. W. Merschoff, NRC Region IV Regional Administrator (w/o)

Mr. W. D. Reckley, NRR Project Manager (ANO-1)

ENCLOSURE I CNRO-2002-00058 FRAMATOME DOCUMENT 32-5021538 (ANO CALCULATION 86-E-0074-156)

20697-6 (2/2002)

A

? A 35 A KID CALCULATION

SUMMARY

SHEET (CSS)

I--KAukiVI/-,I %JIVII-V /A Document Identifier 32 -5021538 - 01 Title ANO-1 CRDM NOZZLE IDTB J-GROOVE WELD FLAW EVALUATION REVIEWED BY:

PREPARED BY:

METHOD. Z DETAILED CHECK nI INDEPENDENT CALCULATION NAME D.E. KILLIAN Zý

-(6V;

" NAME A.D. NANA SIGNATURE SIGNATURE

-I DATE 1 21/7-TITLE PRINCIPAL ENGR. DATE /2/c// 9 TITLE ADVISORY ENGR.

REF. TM STATEMENT:

COST CENTER 41026 PAGE(S) 37 REVIEWER INDEPENDENCE

- #1ý40 PURPOSE AND

SUMMARY

OF RESULTS:

Revision 1: This revision is a non-proprietary version of Revision 0.

suitability of leaving The purpose of the present analysis is to determine from a fracture mechanics viewpoint the following the repair of a CRDM degraded J-groove weld and butter material in the ANO Unit 1 reactor vessel head would combine with a nozzle by the IDtemper bead weld procedure. It is postulated that a small flaw in the head into the low alloy large stress corrosion crack in the weld and butter to form a radial corner flaw that would propagate and cooldown.

steel head by fatigue crack growth under cyclic loading conditions associated with heatup

[ ]" radial crack in Based on an evaluation of fatigue crack growth into the low alloy steel head, a postulated for 25 years of the Alloy 182 J-groove weld and butter would be acceptable from a fracture mechanics viewpoint operation.

THE DOCUMENT CONTAINS ASSUMPTIONS THAT THE FOLLOWING COMPUTER CODES HAVE BEEN USED IN THIS DOCUMENT: MUST BE VERIFIED PRIOR TO USE ON SAFETY RELATED WORK CODENERSION/REV CODENERSION/REV F[1 YES F7 NO Page 1 of 49

A 32-5021538-01 FRAMAI MAE ANP RECORD OF REVISIONS Affected Pages Description of Revision Date Revision 11/02 0 All Original release 12/02 1 All Revision 1 is a non-proprietary version of Revision 0.

2

A FRAMATOME ANP 32-5021538-01 CONTENTS Headinq Paqe Section Introduction .......................................................................................................... 4 1.0 Geom etry and Flaw Model .............................................................................. 6 2.0 3.0 Material Properties ............................................................................................... 9 Fracture Mechanics Methodology ................................................................ 11 4.0 5.0 Applied Stresses ........................................................................................... 12 Flaw Evaluations ........................................................................................... 24 6.0 Sum mary of Results ...................................................................................... 36 7.0 References ................................................................................................... 37 8.0 Appendix Heading Page A Nozzle 56 Considerations ............................................................................. 38 B Certification Document ................................................................................. 46 C Letter from Entergy for Reactor Trip Transients ............................................ 47 3

A 32-5021538-01 FRAMATOME ANP 1.0 Introduction Due to the susceptibility of Alloy 600 partial penetration nozzles to primary water stress corrosion cracking (PWSCC), an ID temper bead weld repair procedure has been developed for reactor vessel head control rod drive mechanism (CRDM) nozzles at ANO Unit 1 wherein the lower portion of a degraded nozzle is removed by a boring procedure and the remaining portion of the nozzle is welded to the low alloy steel reactor vessel head above the original Alloy 182 J-groove attachment

[1]

weld, as shown in Figure 1. This repair design is more fully described by the design drawing and the technical requirements document [2]. Except for a chamfer at the comer, the original J groove weld will not be removed. Since a potential flaw in the J-groove weld can not be sized by currently available non-destructive examination techniques, it is assumed that the "as-left" condition of the remaining J-groove weld includes degraded or cracked weld material extending through the to entire J-groove weld and Alloy 182 butter material. The purpose of the present analysis is of leaving degraded J-groove weld determine from a fracture mechanics viewpoint the suitability material in the vessel following the repair of a CRDM nozzle by the ID temper bead weld procedure.

[3]

Since it is known from analysis of the ANO-1 CRDM reactor vessel head nozzle penetrations that the hoop stress in the J-groove weld is greater than the axial stress at the same location, It is the preferential direction for cracking would be axial, or radial relative to the nozzle.

metal would propagate by PWSCC, through the postulated that a radial crack in the Alloy 182 weld weld and butter, to the interface with the low alloy steel head. It is fully expected that such a crack would then blunt and arrest at the butter-to-head interface [4]. On the uphill side of the nozzle, where the hoop stresses are highest [3] and the area of the J-groove weld is the largest [1], a radial crack depth extending from the comer of the weld to the low alloy steel head would be very deep, about 13/4" for the outermost row of nozzles (closest to the head-to-vessel flange). Since the penetration angle decreases toward the center of the head, crack depths would be progressively the less at each row of nozzles inward from the outermost row. For the present analysis of remaining J-groove weld, it is postulated that a small fatigue initiated flaw forms in the low alloy steel head and combines with the stress corrosion crack in the weld to form a large radial corner flaw that would propagate into the low alloy steel head by fatigue crack growth under cyclic loading conditions associated with the normal condition heatup and cooldown transients and the upset condition reactor trip transient The rod withdrawal accident transient, which is classified as an emergency condition transient for the repaired CRDM nozzle in Appendix B, is the most severe emergency or faulted condition transient for the J-groove weld region. The final faw size is then evaluated for this rod withdrawal accident transient.

The size of the postulated flaw is to the depth of the J-groove weld prep as measured along the inside surface of the head penetration. The repair design [1] specifies that the inside corner of the J-groove weld is to be progressively chamfered from the center to outermost penetrations to maintain a constant [ ]" effective J-groove weld prep at all locations. Thus for the present analysis, [ ]" is used for the initial size of the postulated nozzle corner flaw in the low alloy steel head.

Nozzle 56 was previously repaired in April 2001 by partially removing an axial flaw in the J groove weld and nozzle wall and welding with Alloy [ ] filler material. Appendix A presents an assessment of the effects of this repair on the flaw evaluation of the chamfered J-groove weld.

4

A

=InA, AATt A" A?,"f 32-5021538-01 Figure 1. IDTemper Bead Weld Repair 5

A 32-5021538-01 FRAMATOME ANP 2.0 Geometry and Flaw Model it is postulated that a radial flaw is present in the low alloy steel head, extending from the chamfered corner of the remaining J-groove weld to the interface between the butter and head.

Analytically, this flaw is crudely simulated using the comer flaw model shown below in Figure 2.

1 Stress Line Figure 2. Corner Flaw Model The flaw depth, "a", is the radius to the crack front. The stress line shown in the figure above depicts a typical direction for consideration of a one-dimensional variation of stress through the area represented by the corner flaw model.

The height of the original J-groove weld prep varies around the bore for all nozzle penetrations other than the one at the center of the head. In order to maintain a constant depth of remaining weld (and therefore postulated flaw depth) at all penetrations repaired by the ID temper bead repair procedure, the design drawing [1] specifies a chamfer at the inside corner of the J-groove weld to limit the height of the weld along the bored surface, from the inside corner to the low alloy steel head, to [ ]". Thus for the present flaw evaluations, the maximum initial flaw depth that need be considered is a=[ ]in.[1].

Fatigue stresses for the ID temper bead repair are obtained from a three-dimensional finite element structural analysis [6] that determined operating stresses in the remaining CRDM nozzle, the new weld, the remaining J-groove weld, and the reactor vessel head. The finite element model of the extreme hillside nozzle (38.50 penetration angle) includes a detailed geometrical representation of the remaining J-groove weld prep around the penetration.

Stresses are reported along a line originating at the inside comer (Point 0) and passing through the curved portion of the weld prep, as shown in Figure 3. The orientation of the stress line, relative to the vertical bored surface, is set at the maximum angle from the vertical (-15') where the distance along the line, from Point 0 to the interface between the butter and head, is still nearly equal to the height of the weld along the bore.

6

A 32-5021538-01 FRAMATOME ANP The design drawing [1] specifies a [ 3"tolerance on the size of the chamfer (for example, [ ]"

to [ I"for nozzles 62 through 69), which could create a smaller remaining weld height than the

[ 3"maximum. Although this would result in a smaller initial depth for the postulated flaw, the larger chamfer produces higher stresses in the weld. Accordingly, a maximum chamfer of [ ]"

was included in the finite element model to provide an upper bound on stress (weld height is approximately [ ]"). This results in a conservative analytical strategy, wherein bounding stresses (based on a maximum chamfer of [ ]") are coupled with a maximum flaw depth (based on a minimum chamfer of [ ]") in the fracture mechanics analysis.

7

A r- m PA KIM 32-5021538-01 Direction for flaw growth and variation of stress.

Point 0 Figure 3. Orientation of Stress Line 8

A FRAMATOME ANP 32-5021538-01 3.0 Material Properties The center portion of the ANO-1 reactor vessel head (closure head center disc) containing the CRDM nozzles is made from low alloy steel plate that is equivalent to [

I material [5, 71.

Yield Strength From the ASME Code,Section III, Appendix I [8], the minimum yield strength for the head material is 43.8 ksi at 600 OF. This is used as a conservative lower bound for yield strengths at operating temperatures less than 600 OF.

Reference Nil-Ductility Temperature Based on a highest measured RTNDT of [ ] OF for 13 heats of [ ] plate material

[9], a value of [ ] OF is conservatively used as the RTNDT of the [ ] low alloy steel head material.

Fracture Toughness The lower bound KI, curve of Section XI, Appendix A, Figure A-4200-1 [10], which can be expressed as K1, = 26.8 + 1.233 exp [ 0.0145 (T- RTNDT + 160)], [11]

represents the fracture toughness for crack arrest, where T is the crack tip temperature and RT*- is the reference nil-ductility temperature of the material. The corresponding fracture toughness curve for crack initiation is KI, = 33.2 + 2.806 exp [ 0.02 (T - RTNDT + 100)]. [111 K1, and K1, are in terms of ksi'lin and T and RTNDT are in OF. In the present flaw evaluations, both measures of fracture toughness are limited to a maximum value of 200 ksi4in for upper-shelf fracture toughness. For an RTNDT of [ ] OF, Kia and Kic equal 200 ksWin at crack tip temperatures of 242 OF and 164 OF, respectively. The lowest service temperatures considered in the present flaw evaluations are 250 OF for normal and upset conditions and 600 OF for the emergency rod withdrawal accident transient, confirming that operation is in the upper-shelf region of the fracture toughness curves.

9

A FRAMATOME ANP 32-5021538-01 Fatigue Crack Growth Flaw growth due to cyclic loading is calculated using the fatigue crack growth rate model from Article A-4300 of Section Xl [10],

d--a = Co(AK,I dN where AKI is the stress intensity factor range in ksi4in and da/dN is in inches/cycle. The crack growth rates for a surface flaw will be used for the evaluation of the corner crack since it is assumed that the degraded condition of the J-groove weld and butter exposes the low alloy steel bead material to the primary water environment.

Fatigue Crack Growth Rates for Low Alloy Ferritic Steels in a Primary Water Environment Source: ASME Code, Section Xl, 1992 Edition with No Addenda [10]

AKI = Klma, - KImn R = KImn / Klmax 0:_ R _0.25: AK, < 17.74, n = 5.95 Co = 1.02 x 10-12 x S S= 1.0 AKI > 17.74, n = 1.95 Co= 1.01 x 10-7x S S= 1.0 0.25*5 R < 0.65: AK4 < 17.74 [(3.75R + 0.06) / (26.9R - 5.725) ]o 25, n = 5.95 Co = 1.02 x 10-12 x S S = 26.9R - 5.725 AK1 _ 17.74 [(3.75R + 0.06) / (26.9R - 5.725) ]025, n = 1.95 C0= 1.01 x 10-7 x S S = 3.75R + 0.06 0.65:_ R < 1.0: AK4 < 12.04, n = 5.95 Co = 1.02 x 10-12 x S S = 11.76 K1 >_12.04, n = 1.95 Co= 1.01 x 10-7 x S S = 2.5 10

A FRAMATOME ANP 32-5021538-01 4.0 Fracture Mechanics Methodology The corner crack is analyzed using the following stress intensity factor solution:

K, =4F[O.-706(A0 +Ap)+0.537(2ajA, +O0.448 ajj)A2 + 0.393C4a 3A]

[ Ref. 11, Eqn. (G-2.2)]

where a is the depth of the crack and Ap is a term added to the Reference 11 solution to account for pressure on the crack face.

The stress distribution in the radial direction is described by the third-order polynomial, A o + A ix + A 2 x2 + A 3x 3 , [ Ref. 11, Eqn. (G-2.1)]

where x is measured from the inside corner of the chamfered weld, as shown in Figure 3.

From previous experience it is known that the comer crack model produces conservative results when compared to more rigorous stress intensity factor solutions calculated by the KCALC routine in the finite element code ANSYS.

Irwin Plasticity Correction The Irwin plasticity correction is used to account for a moderate amount of yielding at the crack tip. For plane strain conditions, this correction is defined by ry 6*K ,(a) 2 6nr ay, where, K1(a) = stress intensity factor based on the actual crack length, a, TY = material yield strength.

A stress intensity factor, K,(ae), is then calculated based on the effective crack length, ae =a+ry.

11

A FRAMATOME ANP 32-5021538-01 5.0 Applied Stresses Operational stresses are obtained from the results of a three-dimensional linear finite element analysis of the outermost CRDM nozzle head penetration that addresses the configuration after from repair by the ID temper bead weld procedure of Reference 1. Stresses are available around the nozzle bore for various Reference 6 at the 0', 450, 900, 1350, and the 1800 locations times during a combined normal and upset condition transient that includes heatup, cooldown, and a reactor trip. Additional stresses are available for the rod withdrawal accident transient, which is classified as an upset condition in the reactor coolant system functional specification

[12]. Since ANO-1 has not experienced any of these transients during its plant life [14], the rod withdrawal accident is analyzed as an emergency condition transient, with no contribution to fatigue crack growth. This exception to the functional specification is noted in the document certification statement contained in Appendix B.

The 00 and 1800 locations are at the downhill and uphill sides of the nozzle, respectively.

Stresses were reported in a cylindrical coordinate system relative to the nozzle so that the stress directions remain constant around the nozzle. The largest hoop stresses are found at the uphill side of the nozzle bore, or at the 1800 location. These stresses are perpendicular to the crack face and tend to open the corner crack. The operational stresses from Reference 6, calculated for the extreme outermost CRDM nozzle location, conservatively bound the stresses at all other nozzle locations.

The highest hoop stresses for the combined heatup-reactor trip-cooldown transient occur at 10.125 hours0.00145 days 0.0347 hours 2.066799e-4 weeks 4.75625e-5 months into the transient, during the reactor trip portion. Stresses are also significant at 4.8714 hours0.101 days 2.421 hours 0.0144 weeks 0.00332 months (heatup), 10.0 hours0 days 0 hours 0 weeks 0 months (steady state), and 12.939 hours0.0109 days 0.261 hours 0.00155 weeks 3.572895e-4 months (when the decay heat system is initiated during cooldown following a reactor trip). Due to the dominating influence of pressure on stress, stresses remain positive throughout the combined heatup-reactor trip cooldown transient. The hoop stresses for these four loading conditions are listed in Table 1 for the uphill (1800) location as a function of the radial position along the stress line shown in Figures'2 and 3. The maximum stresses that occur during the rod withdrawal accident transient are listed in Table 2. Although stresses are reported for 13 positions along the stress line, ranging from 0" to 8.82" from the inside corner of the nozzle bore, only the first 7 positions are used to form a third-order polynomial stress fit over a distance of approximately 1.92" from the surface, which includes the weld area and a small portion of the reactor vessel head.

The functional specification for the reactor coolant system [12] specifies one emergency condition transient, a stuck open turbine bypass valve, and two faulted condition transients, a steam line break and a loss of coolant accident. The pressures and temperatures associated with these transients are evaluated in Appendix F of the stress analysis [6] against the governing upset reactor trip transient at 10.125 hours0.00145 days 0.0347 hours 2.066799e-4 weeks 4.75625e-5 months . Reference 6 concludes that the stresses resulting from these three transients are bounded by those for the analyzed reactor trip transient-12

A

,'-DAb*AATCtNAIA I" l P 32-5021538-01 Table 1. Normal/Upset Condition Hoop Stresses [6]

Loading Condition Position Heatup (HU) Steady State (SS) Reactor Trip (RT) Cooldown (CD)

Along Time = 4.8714 hr. Time = 10.000 hr. Time = 10.125 hr. Time = 12.939 hr.

Stress Temp. = 559 OF Temp. = 604 OF Temp. = 531 °F Temp. = 250 OF Line Pres. = [ ] psig Pres. = [ ] psig Pres. = [ ] psig Pres. = [ ] psig x (in.) Stress (psi) Stress (psi) Stress (psi) Stress (psi) 0.0000 [ ] [ ] [ ] [ ]

0.2799 [ ] [ ] [ ] [ ]

0.5597 [ 1 [ 1 [ 1 [ 1 0.8396 [ 1 [ ] [ ] [ ]

1.1195 [ ] [ ] [ ] [ ]

1.5217 [ ] [ ] [ ] [ ]

1.9238 [ ] [ ] [ ] [

2.3260 [ ] [ ] [ ] [ ]

2.7282 [ [ ] [ ] ]

4.2523 [ ] [ ] ] [ ]

5.7764 [ ] [ ] [ ] [ ]

7.3004 [ ] [ ] [ ] [ ]

8.8245 [ ] [ ] [ ] [ ]

13

A

-- a .. arfltar! AKPfl 32-5021538-01 l--F-<.*

FR MM I t..*l",/ll:,,.

,11/* LJDIVEr /.-*t*

,INP f'"

Table 2. Emergency Condition Hoop Stresses [6]

Loading Condition Position Rod Withdrawal Accident (RWA)

Along Time = 0.0044 hours5.092593e-4 days 0.0122 hours 7.275132e-5 weeks 1.6742e-5 months Stress Temperature = 600 'F Line Pressure = [ ] psig x (in.) Stress (psi) 0.0000 [ ]

0.2799 [ ]

0.5597 [

0.8396 [ ]

1.1195 [ ]

1.5217 [

1.9238 [ ]

2.3260 [

2.7282 [ 1 4.2523 [

5.7764 [

7.3004 [ ]

8.8245 [ ]

14

A FRAMATOME ANP 32-5021538-01 Consideration of residual stresses Two three-dimensional elastic-plastic finite element analyses [3] were performed by Dominion Engineering, Inc. (DEI) to simulate the sequence of steps involved in arriving at the ID temper configuration of the CRDM nozzle and reactor vessel head after completion of the model depicted in Figure 4 to bead repair. The first set of analyses used the finite element as listed below.

determine stresses for five combinations of nozzle angle and yield strength, DEI Analysis Cases Nozzle Yield Angle* Strength Center nozzle 0.0 [ ] ksi 18.20 [ ] ksi 26.20 [ ] ksi Outermost nozzle 38.50 [ ] ksi Outermost nozzle 38.50 [ ] ksi

Relative to center of head.

This analysis simulated the laying of the original weld butter and the subsequent post-weld stress relief, the heatup of the original J-groove weld and adjacent material during the welding and process and the subsequent cooldown to ambient temperature, a pre-service hydro test, state loads were removed and the operation at steady state conditions. After the steady from the model. The structure was again at ambient conditions, the nozzle was removed stresses associated with this repair configuration represent the residual stresses in the vicinity of the original J-groove weld, butter, and adjacent head, prior to chamfering the J-groove weld and completing the ID temper bead repair in the region above the J-groove weld. The residual stresses for the outermost penetration (38.50) are listed in Table 3 and plotted in Figure 5.

The second analysis performed by Dominion Engineering addressed the final steps involved in completing the ID temper bead repair weld. For the sake of simplicity, the center nozzle was the modeled in an axisymmetric analysis (Figure 6) to simulate the removal of the nozzle below cut line [1], the deposition of the repair weld using four weld passes, and the chamfering of the J-groove weld (Figure 7). Although this analysis was performed for the center nozzle, the location of the repair weld above the J-groove and the size of the chamfer ([ ]") were specified to represent the worst case geometry at the outermost nozzle, where the distance between the two welds is a minimum and the size of the chamfer is a maximum. From the DEI analysis [3],

the minimum vertical distance between welds is [ ]" (difference between the Z-coordinates of nodes 2009 and 1409). The final residual stresses for the representative penetration are listed in Table 4 and plotted in Figure 8.

Although at shutdown, the residual hoop stress in the weld region is high, above [ ] psi (Figures 5 and 8), the stress decreases to zero just beyond the butter region and is compressive in the head. These residual stresses would be relieved as the crack propagates through the 15

A 32-5021538-01 FRAMATOME ANP weld and butter and a short distance into the head. Deeper cracks would then experience only compressive residual stress ahead of the crack tip. It can be seen from Figure 8 and Table 4 that the residual stresses are compressive at some distance less than [ ]" into the head.

The depth of the initial flaw size will therefore be increased by this amount so that residual stresses need not be considered in the present flaw evaluation.

16

A 32-502 1538-0 1 t- K IOL% IOA I IVI = 1 14 r-80006 80001 1 t phill Pljli.: \.'xke' a re tWV14'S.cric 623

%,i~dI\',ý..¶-Scie a NbwI I1)uiwrpo..d %tuKh4)D) in uedJ tlqfl Wit 7' t She1ID aLtxv*.%ldJrqx 23.at C4.d. o.f shell seMk'fl

\&.&\,,~vr~In~~k th L-n-h tr vi Itrbe, anl shell Itil tii~p S&\tuxnblim.¶ .~~

Nar I ahon6 th Itbandi Aelidjiti,~

Figure 4. DEl Model for CRDM Nozzle Stress Analysis 17

~-u O-U Framatome ANP Table 3.

Residual Hoop Stresses After Nozzle Removal Penetration angle = 38.5 degrees Nozzle yield strength = I ] ksi File: ANO1-38B.results.txt [3]

Time: 7006 Distance Global Coordinates Hoop into X Z ASO)} Location Stress Head Node (in.) (in.) (in.) (psi) (in.)

80606 2.0000 69.680 0.000 Inside Surface of Weld 80708 2.1128 69.976 0.317 Weld 80808 2.1059 70.183 0.514 Weld 80909 2.1979 70.440 0.785 Weld 81010 2.2760 70.670 1.028 Weld 81111 2.3403 70.874 1.242 Weld 81212 2.3881 71.050 1.424 Weld 81313 2.4246 71.202 1.580 Weld 81415 2.5142 71.313 1.712 Weld/Butter Interface 81516 2.5672 71.416 1.826 Butter 2.6230 71.511 1.934 Butter/Head Interface 0.000 81617 2.6869 71.726 2.158 Head 0.224 81717 2.7637 71.984 2.427 Head 0.493 81817 81917 2.8558 72.295 2.751 Head 0.817 2.9664 72.667 3.139 Head 1.205 82017 3.0992 73.114 3.606 Head 1.672 82117 82217 3.2586 73.650 4.165 Head 2.231 82317 3.4499 74.294 4.836 Head 2.902 82417 3.6795 75.066 5.642 Head 3.708 82517 3.9551 75.994 6.610 Head 4.676 82617 4.2860 77.108 7.772 Head 5.838 4.6832 78.444 9.166 Head 7.231 82717 5.1600 80.049 10.840 Head 8.906 82817

(') Distance from inside comer (node 80606) - see Figure 4.

18

Framatome ANP 32-5021538-01 Figure 5, Residual Hoop Stresses After Nozzle Removal for Outermost Penetration at 38.50 0o C/)

CL.

00 0r 0 2 4 6 8 10 12 Distance from Surface, in.

19

A 32-5021538-01 FRAMATOME ANP Node Numbers Increase by I00 up tilekicith of the tube and shell Node Numbers increase b% I aklomlý the tube and shell radius Nodes 609 ihroumh 1400 are coincident %%ith610 tlitough 1410 Figure 6. DEl Model for Center CRDM Nozzle with Weld Repair 20

A","

A trn ,A Ar-T(h IVI AA r ~l,, 32-5021538-01 F-m N\,.i,. Repair XeId R.!icsll S~450 Reg ioni Removed fbr Weld Chi mfer Figure 7. DEI Model for Center CRDM Nozzle After Weld Repair and Chamfer 21

Framatome ANP 32-5021538-01 Table 4.

Residual Hoop Stresses After Repair Weld and Chamfer Penetration angle = 0 degrees Nozzle yield strength = [ ] ksi File: ANO1-OB.results.txt [3]

Time: 16001 Distance Global Coordinates Hoop into Node X Z AS(') Location Stress Head (in.) (in.) (in.) (psi) (in.)

2.0000 86.686 0.000 Inside Surface of Weld 2.0852 86.924 0.253 Weld 2.0821 87.077 0.400 Weld 2.1579 87.278 0.613 Weld 2.2273 87.453 0.800 Weld 1114 2.2905 87.604 0.963 Weld 1215 2.3572 87.708 1.083 Weld 1316 2.4077 87.809 1.195 Weld 1418 2.5053 87.893 1.309 Weld/Butter Interface 1519 2.5578 88.002 1.429 Butter 1620 2.6141 88.100 1.542 Butter/Head Interface 0.000 1720 2.6933 88.286 1.744 Head 0.202 1820 2.7850 88.501 1.977 Head 0.436 1920 2.8911 88.749 2.247 Head 0.706 2020 3.0140 89.037 2.560 Head 1.019 2120 3.1562 89.370 2.922 Head 1.381 2220 3.3209 89.756 3.342 Head 1.801 2320 3.5114 90.203 3.828 Head 2.286 2420 3.7320 90.720 4.390 Head 2.849 2520 3.9874 91.318 5.040 Head 3.499 2620 4.2830 92.011 5.794 Head 4.252 2720 4.6251 92.813 6.666 Head 5.124 2820 5.0212 93.741 7.675 Head 6.133

(" Distance from original inside corner (node 609) - see Figure 6.

22

32-5021538-01 Framatome ANP Penetration Figure 8. Residual Hoop Stresses After Repair Weld and Chamfer for Representative I I II Butter/Head I II c II CL 0.

II Weld/Butter II II Interface II I p 7 8 6 7 8 9 0 1 2 3 4 5 Distance from Surface, in.

23

A FRAMATOME ANP 32-5021538-01 6.0 Flaw Evaluations The actual fracture mechanics calculations are presented in Tables 5 through 7. Evaluations are performed for a postulated radial corner crack on the uphill side of the outermost CRDM nozzle head penetration. The applicable hoop stresses (perpendicular to the plane of the postulated crack) are listed in Tables I and 2 for key time points from the finite element stress analysis [6].

As required by Article IWB-3612 [10], a safety factor of 410 is used for normal and upset conditions with the lower KIa fracture toughness for crack arrest. Article IWB-3612 also specifies that a safety factor of 42 must be used for emergency and faulted conditions, along with the higher K1, fracture toughness for crack initiation. As discussed in Section 5.0, the reactor trip stresses bound the emergency and faulted condition stresses. The flaw evaluation performed for the reactor trip transient therefore serves as a bounding analysis for the emergency and faulted condition analysis required by IWB-3612 (excluding the rod withdrawal accident that is classified as an emergency condition for the present analysis).

Based on the stresses listed in Tables 1 and 2, two groups of transient conditions are considered for fatigue crack growth. For the heatup/cooldown transient, the bounding load pair consists of heatup at 4.8714 hours0.101 days 2.421 hours 0.0144 weeks 0.00332 months and shutdown (zero stress state). Reactor trip stresses at 10.125 hours0.00145 days 0.0347 hours 2.066799e-4 weeks 4.75625e-5 months and steady state stresses at 10.0 hours0 days 0 hours 0 weeks 0 months are used for the reactor trip cyclic load conditions.

Fatigue crack growth is calculated on a yearly basis using the following pattern for accumulating cycles:

Fatigue Crack Growth Cycles Transient Heatup / Cooldown Reactor Trip Cycles / 40 Years 240 400 Cycles / Year 6 10 These cycles are distributed uniformly over the 25 year service life by linking the incremental crack growth between Table 5 (for heatup and cooldown) and Table 6 (for reactor trip). The final flaw size is evaluated in Table 7 for the rod withdrawal accident.

24

Framatome ANP 32-5021538-01 Table 5. Evaluation of CRDM Nozzle Corner Crack for Heatup/Cooldown INPUT DATA Initial Flaw Size: Distance to base metal: F Additional distance to compressive zone: L Total depth, a =E -in.

Material Data: Yield strength, Sy = 43.8 ksi Reference temp., RTndt= E ] F Upper shelf tough. = 200 ksi'/in Kla = 26.8 + 1.233 exp [ 0.0145 (T - RTndt + 160)]

Kla is limited to the upper shelf toughness.

Applied Loads:

Loading Conditions HU* SD**

Temperature (F) 559 70 Pressure, p (ksi)

Kia (ksin)

Position 200 41 x Hoop Stress (in.) (ksi) (ksi) 0.0000 0.2799 0.5597 0.8396 1.1195 1.5217 1.9238 2.3260 2.7282 42523 5.7764 7.3004 8.8245

Heatup at 4.8714 hours0.101 days 2.421 hours 0.0144 weeks 0.00332 months Shutdown 25

32-5021538-01 Framatome ANP Table 5. Evaluation of CRDM Nozzle Corner Crack for Heatup/Cooldown (Cont'd)

STRESS INTENSITY FACTOR 2 3 KI(a) = "(a)[0.706(Ao+Ap) + 0.537(2ahr)A 1 + 0.448(a /2)A 2 + 0.393(4a /31!)A 3 ]

where the through-wall stress distribution is described by the third order polynomial, S(x) = A, + A~x + A2x2 + A3x3, defined by:

Stress Loading Conditions Coeff. HU SD (ksi) (ksi)

A1 A2 A3 Effective crack size:

2 ae = a + 1/(61)*[KI(a)/Sy1 Effective stress intensity factor:

3 KI(ae) = (7aj)[ 0.706(Ao+Ap) + 0.537(2ajh)A 1 + 0.448(ae2/2)A2 + 0.393(4ae /3n)A 3 ]

26

32-5021538-01 Framatome ANP Table 5. Evaluation of CRDM Nozzle Corner Crack for Heatup/Cooldown (Cont'd)

FATIGUE CRACK GROWTH Transient

Description:

240 cycles over 40 years AN = 6 cyclesfyear HU SD HU SD HU SD Operating AKI Aa ae ae KI(ae) Kl(ae)

Tune Cycle a Kl(a) Kl(a)

(ksiin) (in.) (in.) (in.) (ksi~n) (ksiqin)

(yr.) (in.) (ksiqin) (ksiqin) 0.00107 46.80 0.00 0.0 0 46.29 0.00 46.29 46.31 0.00107 46.81 0.00 1.0 6 46.31 0.00 46.32 0.00107 46.83 0.00 2.0 12 46.32 0.00 46.34 0.00 46.34 0.00107 46.84 0.00 3.0 18 46.35 0 00 46.35 0.00107 46.86 0.00 4.0 24 46.36 0.00108 46.87 0.00 5.0 30 46.36 0.00 46.38 0.00 46.38 0.00108 46.88 0.00 6.0 36 46.39 0.00 46.39 0.00108 46.90 0.00 7.0 42 46.41 0.00 46.41 0.00108 46.91 0.00 8.0 48 46.42 0.00 46.42 0.00108 46.93 0.00 9.0 54 46.44 0.00 46.44 0.00108 46.94 0.00 10.0 60 46.45 0.00 46A5 0.00108 46.95 0.00 11.0 66 46.46 0.00 46.46 0.00108 46.97 0.00 12.0 72 46.48 0.00 46.48 0.00108 46.98 0.00 13.0 78 46.49 0.00 46.49 0.00108 47.00 0 00 14.0 84 46.51 0.00 46.51 0.00108 47.01 0.00 15.0 90 46.52 0.00 46.52 0.00108 47.02 0.00 16.0 96 46.53 0.00 46.53 0.00108 47.04 0.00 17.0 102 46.55 0.00 46.55 0.00108 47.05 0.00 1B.0 108 46.56 0.00 46.56 0.00108 47.07 0 00 19.0 ..114 46.58 0.00 46.58 0.00108 47.08 0.00 20.0 120 46.59 0.00 46.59 0.00109 47.09 0.00 21.0 126 132 46.61 0.00 46.61 0.00109 47.11 0.00 22.0 46.62 0.00 46.62 0.00109 47.12 0.00 23.0 138 46.63 0.00 46.63 0.00109 47.13 0.00 24.0 144 150 46.65 0.00 46.65 0.00109 47.15 0.00 25.0 27

32-5021538-01 Framatome ANP Table 5. Evaluation of CRDM Nozzle Comer Crack for Heatup/Cooldown (Cont'd)

FRACTURE TOUGHNESS MARGINS Period of Operation: Time = 25.0 years Final Flaw Size: a= [ ] in. (after reactor trip)

Margin = Kla / KI(ae)

Loading Conditions HU SD Fracture Toughness, Kla 200.0 41.3 ksi&in KI(a) 46.66 0.00 ksifin ae Kl(ae) 47.16 0.00 ksilin Actual Margin 4.24 #NIA Required Margin 3.16 #NIA 28

32-5021538-01 Framatome ANP Table 6. Evaluation of CRDM Nozzle Comer Crack for Reactor Trip INPUT DATA Beginning Flaw Size: Depth, Yield strength, Sy = 43.8 ksi Material Data:

Reference temp., RTndt = ] F Upper shelf tough. = 200 ksWin Kla = 26.8 + 1.233 exp [ 0.0145 (T - RTndt + 160)]

Kla is limited to the upper shelf toughness.

Applied Loads:

Loading Conditions RT* SS**

Temperature (F) 531 604 Pressure, p (ksi)

Kla (ksin)

Position 200 200 x Hoop Stress (in.) (ksi) (ksi) 0.0000 0.2799 0.5597 0.8396 1.1195 1.5217 1.9238 2.3260 2.7282 4.2523 5.7764 7.3004 8.8245 Reactor Trip at 10.125 hours0.00145 days 0.0347 hours 2.066799e-4 weeks 4.75625e-5 months Steady State at 10.000 hours0 days 0 hours 0 weeks 0 months 29

32-5021538-01 Framatome ANP Table 6. Evaluation of CRDM Nozzle Comer Crack for Reactor Trip (Cont'd)

STRESS INTENSITY FACTOR Kl(a) = %1(7Ta) [ 0.706(Ao+Ap) + 0.537(2a/h)A 1 + 0.448(a 2/2)A2 + 0.393(4a3 /3nt)A 3]

where the through-wall stress distribution is described by the third order polynomial, S(x) = Ak + Alx + A2x' + A3x3, defined by:

Stress Loading Conditions Coeff. RT SS (ksi) (ksi)

A0 A,

A2 A3 Effective crack size:

ae = a + 1/(6n)*[Kl(a)ISyI 2 Effective stress intensity factor:

2 Kl(ae) = 'I(ae) [ 0.706(Ao+Ap) + 0.537(2ahTc)AI + 0.448(ae /2)A 2 + 0.393(4ae3/337)A 3]

30

Framatome ANP 32-5021538-01 Table 6. Evaluation of CRDM Nozzle Comer Crack for Reactor Trip (Cont'd)

FATIGUE CRACK GROWTH Transient

Description:

400 cycles over 40 years AN = 10 cycles/year RT SS RT SS RT SS Operating ae KI(ae) KI(ae)

Kl(a) AKI Aa ae Time Cycle a KI(a)

(ksiqin) (in.) (in) (in.) (ksiqin) (ksi-in)

(yr.) (in.) (ksi'iin) (ksiin) 61.63 45.28 0.0 0 60.92 44.80 16.12 0.00057 61.64 45.29 1.0 10 60.93 44.81 16.12 0.00057 61.65 45.31 2.0 20 60.94 44.83 16.11 0.00057 61.66 45.32 3.0 30 60.95 44.84 16.11 0.00057 16.11 0.00057 61.67 45.33 4.0 40 60.96 44.86 61.69 45.35 5.0 50 60.98 44.87 16.11 0.00057 000057 61.70 45.36 6.0 60 60.99 44.88 16.10 0.00057 61.71 45.38 7.0 70 61.00 44.90 16.10 16.10 0.00057 61.72 45.39 8.0 80 61.01 44.91 0.00057 61.73 45.40 9.0 90 61.03 44.93 16.10 0.00057 61.74 45.42 10.0 100 61.04 44.94 16.10 16.09 0.00057 61.75 45.43 11.0 110 61.05 44.96 0.00057 61.76 45.45 12.0 120 61.06 44.97 16.09 0.00057 61.77 45.46 13.0 130 61.07 44.98 16.09 16.09 0.00057 61.78 45.48 14.0 140 61.09 45.00 16.08 000057 61.79 45.49 15.0 150 61.10 45.01 0.00057 61.80 45.50 16.0 160 61.11 45.03 16.08 16.08 0.00057 61.81 45.52 17.0 170 61.12 45.04 16.08 0.00057 61.82 45.53 18.0 180 61.13 45.06 16.08 0.00057 61.83 45.55 19.0 .190 61.15 45.07 0.00057 61.84 45.56 20.0 200 61.16 45.08 16.07 16.07 0.00057 61.85 45.57 21.0 210 61.17 45.10 16.07 000057 61.86 45.59 22.0 220 61.18 45.11 0.00057 61.87 45.60 23.0 230 61.19 45.13 16.07 16.06 0.00057 61.88 45.62 24.0 240 61.20 45.14 16.06 0.00057 61.89 45.63 25.0 250 61.22 45.16 31

32-5021538-01 Framatome ANP Table 6. Evaluation of CRDM Nozzle Corner Crack for Reactor Trip (Cont'd)

FRACTURE TOUGHNESS MARGINS Period of Operation: Time = 25.0 years Final Flaw Size: a = E- ] in.

Margin = Kla / KI(ae)

Loading Conditions RT SS 200.0 200.0 ksiqin Fracture Toughness, Kla 61.22 45.16 ksi4 in KI(a) ae KI(ae) 61.89 45.63 ksi4in Actual Margin 3.23 4.38 Required Margin 3.16 3.16 32

Framatome ANP 32-5021538-01 Table 7. Evaluation of CRDM Nozzle Comer Crack for Rod Withdrawal Accident INPUT DATA Final Flaw Size: Depth, a=[ ] in.(after reactor trip)

Material Data: Yield strength, Sy= 43.8 ksi Reference temp., RTndt =[ ] F Upper shelf tough. = 200 ksiqin KIc = 33.2 + 2.806 exp [ 0.02 (T - RTndt + 100)1 Kic is limited to the upper shelf toughness.

Applied Loads:

Loading Conditions RWA* #N/A Temperature (F) 600 #NIA Pressure, p (ksi)

N/A Kic (ksiqin)

Position 200 #N/A x Hoop Stress (in.) (ksi) (ksi) 0.0000 #N/A 0.2799 #N/A 0.5597 #N/A 0.8396 #NIA 1.1195 #NIA 1.5217 #NIA 1.9238 #N/A 2.3260 #N/A 2.7282 #N/A 4.2523 #NIA 5.7764 #NIA 7.3004 #NIA 8.8245 #N/A

Rod Withdrawal Accident at 0.0044 hours5.092593e-4 days 0.0122 hours 7.275132e-5 weeks 1.6742e-5 months 33

Framatome ANP 32-5021538-01 Table 7. Evaluation of CRDM Nozzle Comer Crack for Rod Withdrawal Accident (Cont'd)

STRESS INTENSITY FACTOR 2 3 KI(a) = 1-,(ia) [ 0.706(Ao+Ap) + 0.537(2a/n)A1 + 0.448(a /2)A 2 + 0.393(4a /3n)A 3 ]

where the through-wall stress distribution is described by the third order polynomial, S(x) = A0 + Ajx + A2x2 + A3 x3.

defined by:

Stress Loading Conditions Coeff. RWA #N/A (ksi) (ksi)

A0 #NIA A, #NIA A2 #NIA A3 #NIA Effective crack size:

2 ae = a + 1/(6t)*[Kl(a)/Sy]

Effective stress intensity factor-.

3 KI(ae) = q(7tae) [ 0.706(Ao+Ap) + 0.537(2aeht)A 1 + 0.448(ae 2/2)A2 + 0.393(4ae I/3i)A 3 ]

34

Framatome ANP 32-5021538-01 Table 7. Evaluation of CRDM Nozzle Comer Crack for Rod Withdrawal Accident (Cont'd)

FRACTURE TOUGHNESS MARGINS Period of Operation: Time = 25.0 years Final Flaw Size: a= - ] in.

Margin = KIc I KI(ae)

Loading Conditions RWA #NIA Fracture Toughness, KIc 200.0 #NIA ksi4in Kl(a) 66.99 #N/A ksi4in ae #N/A KI(ae) 68.34 #N/A ksiqin Actual Margin 2.93 #NIA Required Margin 1.41 #NIA 35

A 32-5021538-01 FRAMATOME ANP 7.0 Summary of Results A fracture mechanics analysis has been performed to evaluate a postulated large radial crack in the remnants of the original J-groove weld and butter after an IDtemper bead repair of a CRDM nozzle reactor vessel head penetration. Results of this analysis are summarized below for the controlling transients.

Reactor Trip at 10.125 hours0.00145 days 0.0347 hours 2.066799e-4 weeks 4.75625e-5 months (Upset Condition)

Temperature, T = 531 'F Initial flaw size, ai= [ ]in.

Final flaw size after 25 years, af = [ ]in.

Flaw growth, af- ai = 0.042 in.

Stress intensity factor at final flaw size, KI = 61.89 ksiqin Fracture toughness at 531 'F, Kia = 200.0 ksibin Safety margin, Kla / KI = 3.23 > ,/10 = 3.16 Rod Withdrawal Accident at 0.0044 hours5.092593e-4 days 0.0122 hours 7.275132e-5 weeks 1.6742e-5 months (Emeraencv Condition)

Temperature, T= 600TF Flaw size at 25 years, a= [ ]in.

Stress intensity factor, KI = 68.34 ksi*in Fracture toughness at 531 'F, Kla = 200.0 ksi4in Safety margin, Kla / Ki = 2.93 > 42 = 1.41 Conclusion Based on an evaluation of fatigue crack growth into the low alloy steel head, the above results demonstrate that a postulated radial crack in the Alloy 182 J-groove weld and butter would be acceptable for 25 years of operation.

36

A 32-5021538-01 FRAMATOME ANP 8.0 References Repair,

1. Framatome ANP Drawing 02-5021508E-2, "CRDM Nozzle ID Temper Bead Weld ANO-1."

2. Framatome ANP Document 51-5021517-01, "ANO-1 CRDM Nozzle ID Ambient Temperature Temper Bead Weld Repair Requirements," November 2002.

2001-0230

3. Framatome ANP Document 38-1290261-00, ANO Design Input Record No.

Calculations for IDTB 020-02, dated 11/6/2002 forwarding Dominion Engineering Stress Weld Repair.

of Low Alloy

4. Framatome ANP Document 51-5012047-00, "Stress Corrosion Cracking Steel," March 2001.

ANO-1.

5. Framatome ANP Drawing 02-135579E-9, "Closure Head Sub-Assembly,"

Weld Analysis,"

6. Framatome ANP Document 32-5012424-10, "CRDM Temperbead Bore November 2002.

ANO-1.

7. Framatome ANP Drawing 02-139669E-1 0, "Material List - Head and Vessel,"

of Nuclear

8. ASME Boiler and Pressure Vessel Code,Section III, Rules for Construction No Addenda.

Power Plant Components, Division 1 - Appendices, 1989 Edition with and

9. BAW-10046A, Rev. 2, "Methods of Compliance With Fracture Toughness Appendix G," B&W Owners Group Materials Operational Requirements of 10 CFR 50, Committee Topical Report, June 1986.

of

10. ASME Boiler and Pressure Vessel Code,Section XI, Rules for Inservice Inspection Nuclear Power Plant Components, 1992 Edition with No Addenda.

of ASME

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

ANO-I."

12. Framatome ANP Document 18-1173987-03, "RCS Functional Specification for 2001-0230

13. Framatome ANP Document 38-1290260-00, ANO Design Input Record No.

for Nozzle 56.

020-01, dated 1114/2002 forwarding Dominion Engineering Stress Results as

14. Letter from Entergy for Reactor Trip Transients, November 8, 2002 (attached Appendix C).

37

A FRAMATOME ANP 32-5021538-01 Appendix A Nozzle 56 Considerations Nozzle 56 was previously repaired in April 2001 by partially removing an axial flaw in the J groove weld and nozzle wall and welding with Alloy [ ] filler material. Dominion Engineering, Inc. (DEI) analyzed the final repair configuration and treated the remaining flaw indication as an embedded flaw. The flaw was modeled as a purely axial flaw in the nozzle (Figure A-I), starting at the nozzle outside surface at the top of the J-groove weld and extending 0.2" radially inward (toward the inside surface) and 1.3" axially upward (along the outside surface of the nozzle). In order to locate a plane of nodes at the center of the weld repair, the repair was modeled as extending over a total of 90 degrees, starting at the downhill position, with the flaw at the center of the 90 degree extent (or 45 degrees from the downhill position).

Dominion Engineering's embedded flaw model for Nozzle 56 is shown in Figure A-1 for a plane oriented 45' from the downhill position. Figure A-1 also indicates the node numbering scheme used for their finite element analysis. The nodes located in the same circumferential plane as the embedded flaw, or 45 degrees from the downhill position, are in the 20,000's range. In addition, for nodes on the open face of the crack (i.e., everywhere but the crack edges), the nodes are "doubled" to allow the crack face to open, and the doubled node numbers are increased by 300,000. On the crack face, the original 20,000's series nodes are kept on the elements closest to the downhill side of the model and the new 320,000's series nodes are transferred to the elements closer to the uphill side of the model.

The effect of the previous weld repair on Nozzle 56 will be assessed by comparing operating hoop stresses for both unrepaired and repaired nozzles. The angular penetration and yield strength of Nozzle 56 places this nozzle within the group modeled at 38.50 and 48.5 ksi yield.

The model for an original CRDM nozzle is shown in Figure 4. Stresses are obtained along a path from the inside comer of the J-groove weld, through the weld and butter, and out into the head, as indicated in Tables A-1 and A-2 for the unrepaired and repaired nozzles, respectively.

These stresses are plotted in Figures A-2 and A-3. In the DEI analyses, stresses at operation include residual stresses from welding and stresses due to pressure and temperature at steady state conditions.

Since the weld sizes are different for the two models, the distance along the stress paths differ for a common point of reference, say the weld-to-butter interface. Table A-3 provides a basis for comparing the results by listing stresses by their relative position along each path line. These "normalized" stresses are plotted in Figure A-4 for the unrepaired and repaired nozzles. Figure A-4 reveals that the operating stresses are higher for the unrepaired nozzle in the butter and head regions of the structure, where the influence of residual stress diminishes. The higher stress in the weld region for the repaired nozzle is due to an increase in residual stress from the deposition of the Alloy [ ] weld filler. These stresses would be relieved as a postulated PWSCC flaw in the J-groove weld propagated through the weld and butter. Thus the repairs made to Nozzle 56 do not adversely affect the results of the J-groove weld flaw evaluation contained in this document.

38

A C: 00 hAA~r1&A9 A"93P 32-5021538-01 22208 22201 Embedded Flaw' 21401 21509,3215091 20601Z /

20607 Z 20611 Alloy 52 Repair Figure A-1 DEl Nozzle 56 Embedded Flaw Model at Center of Repair Zone (Oriented 450 from the Downhill Position) 39

32-5021538-01 Framatome ANP Table A-1 Operating Hoop Stresses at 450 Plane in Unrepaired Nozzle Penetration angle = 38.5 degrees Nozzle yield strength = [ ] ksi File: ANO1-38B.results.txt [3]

Time: 7007 Distance Global Coordinates Hoop into Node x Z ASO 1) Location Stress Head (in.) (in.) (in.) (psi) (in.)

20606 2.0000 66.306 0.000 Inside Surface of Weld 20708 2.1075 66.531 0.249 Weld 20808 2.1013 66.686 0.393 Weld 20909 2.1901 66.873 0.598 Weld 21010 2.2665 67.034 0.775 Weld 21111 2.3303 67.169 0.924 Weld 21212 2.3781 67.253 1.020 Weld 21313 2.4170 67.336 1.111 Weld 21415 2.5110 67.388 1.197 Weld/Butter Interface 21516 2.5675 67.488 1.311 Butter 21617 2.6216 67.588 1.425 Butter/Head Interface 0.000 21717 2.7462 67.743 1.619 Head 0.194 21817 2.8927 67.926 1.850 Head 0.425 21917 3.0650 68.140 2.121 Head 0.696 22017 3.2675 68.392 2.441 Head 1.016 22117 3.5057 68.688 2.818 Head 1.393 22217 3.7856 69.037 3.263 Head 1.838 22317 4.1148 69.446 3.786 Head 2.361 22417 4.5018 69.928 4.402 Head 2.977 22517 4.9568 70.494 5.127 Head 3.702 22617 5.4917 71.160 5.979 Head 4.555 22717 6.1206 71.942 6.982 Head 5.557 22817 6.8600 72.862 8.161 Head 6.736

) Distance from inside corner (node 20606*)

Node 20606 is located on the 450 plane at the same relative position as node 606 in Figure 4.

40

r%-It Framatome ANP 3,--:. , 0-U Table A-2 Operating Hoop Stresses at 450 Plane in Repaired Nozzle 56 Penetration angle = 38.5 degrees Nozzle yield strength = [ ] ksi File: Noz56.flawplane.results.out [13]

Time: 12001 Distance Global Coordinates Hoop into Node X Z AS() Location Stress Head (in.) (in.) (in.) (psi) (in.)

20611 2.0000 66.291 0.000 Inside Surface of Weld 20713 2.0903 66.543 0.268 Weld 20813 2.0843 66.730 0.447 Weld 20914 2.1564 66.951 0.678 Weld 21015 2.2164 67.151 0.887 Weld 21116 2.2642 67.330 1.072 Weld 21217 2.3067 67.462 1.210 Weld 21318 2.3303 67.600 1.350 Weld 21420 2.3828 67.726 1.485 Weld/Butter Interface 21521 2.4576 67.777 1.555 Butter 21622 2.5167 67.869 1.660 Butter/Head Interface 0.000 21722 2.5952 68.119 1.922 Head 0.262 21822 2.6805 68.392 2.208 Head 0.548 21922 2.7733 68.688 2.519 Head 0.858 22022 2.8741 69.010 2.856 Head 1.196 22122 2.9838 69.361 3.224 Head 1.563 22222 3.1030 69.742 3.623 Head 1.963 22322 3.2327 70.156 4.057 Head 2.396 22422 3.3736 70.606 4.528 Head 2.868 22522 3.5268 71.096 5.042 Head 3.381 22622 3.6934 71.628 5.599 Head 3.939 22722 3.8745 72.206 6.205 Head 4.544 22822 4.0715 72.835 6.864 Head 5.204 DOistance from inside corner (node 20611) - see Figure A-1.

41

a Framatome ANP 32-5021538-01 Figure A-2. Operating Hoop Stresses at 450 Plane in Unrepaired Nozzle I1 Butteir/Head

_nterfce II I I c:.

C')

(I) I I U) I I I I C')

I I 0~

0 I I 0

I I I I--I I I I I I I _____ __

I I I I 4 5 6 7 8 9 0 1 2 3 Distance from Surface, in.

42

32-5021538-01 Framatome ANP Figure A-3. Operating Hoop Stresses at 450 Plane in Repaired Nozzle 56' I I I I Butter/Head I I nterface

____', :- ...... .~.. - - .- - -.

I I F I I I I I I I 0*

V..

C.

0 0

""-r Weld/Butter i -I . . . . .

Interface

! tI 6 3 4 5 6 7 8 0 I 2 Distance from Surface, in.

43

32-5021538-01 Framatome ANP Table A-3 Operating Hoop Stresses at 450 Plane Penetration angle = 38.5 degrees Nozzle yield strength = [ ] ksi Relative Unrepaired Nozzle Repaired Nozzle 56 Node Hoop Node Hoop Nodal Number Stress Number Stress Position Location (psi) 20611 0 Inside 20708 20713 1 Weld Weld 20808 20813 2

Weld 20909 20914 3

21010 21015 4 Weld 21111 21116 5 Weld 21212 21217 6 Weld 21313 21318 7 Weld 21415 21420 8 Weld/Butter Interface 21516 21521 9 Butter.

21617 21622 10 Butter/Head Interface 21717 21722 11 Head 21817 21822 12 Head 21917 21922 13 Head 22017 22022 14 Head Head 22117 22122 15 Head 22217 22222 16 Head 22317 22322 17 Head 22417 22422 18 Head 22517 22522 19 Head 22617 22622 20 Head 22717 22722 21 22817 22822 22 Head 44

32-5021538-01 Framatome ANP Figure A-4. Operating Hoop Stresses at 450 Plane Cl) o*

Cl.

.4-,

0~

0 "r"

0 2 4 6 8 10 12 14 16 18 20 22 24 Relative Nodal Position Along Stress Line 45

A 32-5021538-01 FRAMATOME ANP Appendix B Certification Document ANO-1 CRDM NOZZLE IDTB J-GROOVE WELD FLAW EVALUATION SHEET 1 OF 1 1 certify that the flaw evaluations contained in this calculation package are evidence that the CRDM nozzle repair, as defined by Framatome ANP drawing 02-5021508E-2, meets the fracture mechanics requirements identified in Framatome ANP Document 51-5021517-01 and the ASME Boiler and Pressure Vessel Code,Section XI, 1992 Edition with no addenda, for the design transients provided in Framatome ANP Document 18-1173987-03.

Exception:

The rod withdrawal transient, identified as Transient No. 11 in the functional specification for the reactor coolant system at ANO-1 (Framatome ANP Document 18-1173987-03), is analyzed as an emergency condition transient since ANO-1 has not experienced any rod withdrawal transient transients during its plant life and the likelihood of even a single future occurrence is remote. It is even more unlikely that a CRDM nozzle would experience 25 cycles of this transient, which is a requirement for classification as an emergency event.

Attested to this date: November 8, 2002 By:_ _ __

Douglas E. Killian Framatome ANP, Inc.

Nuclear Engineering Business Unit Lynchburg, Virginia License No. 15308 Virginia Board for Architects, Professional Engineers, Land Surveyors, Certified Interior Designers, and Landscape Architects 46

A 32-5021538-01 FRAMATOME ANP Appendix C Letter from Entergy for Reactor Trip Transients 47

r-fI WU. Ui-* dqd '4 I IJV NOV-UH-2UO2 HKt UZ:IU Fr ri U $2-5'. " -::t ARKANSAS NUCLEAR ONE GEI3NERATION SUPPORT BUILDING 1448 S. R. 333 RUSSELLVILLE, AR 72301 t ~FAX COVERLETTER DATE ii-.o.1 . . -*

TO ie~

COmpANY NAME .

CITY, STATE FAX NUMBER VERIFY__________

NUMDER OF PAGES INCLUDING COVER FRoM .'%

COMPl'A.NY NAME - Z,14 fgr-'1 DEPARTMENTr FAXN1MTER(501) 858-4955 VERIFY (501) E5S-4302.

TO BE COMPLETED BY SENDER AF'ER TRANSMISSION IS COMPLETE, PLEASE RETURN RY MAIL___ DISCARD CALL FOR PICK-UP PHONE NUMBER

FAX NO. 501 858 4955 P. 02 NOV-08-2002 FRI 02:10 P1K 0 S 32- 5 E>38-o I Steve:

Based on a review performed by our System Engineering, as documented in our ER-ANO-2001 0230-024, ANO-I has not had a "high pressure. rod withdrawal accident" as described In transient No. 11 of RCS Functional Specification 18-1173987-03. See below for details.

++++++÷÷+++÷++++++++++++..........

From: MEANS, BRACY E Sont: Friday, November 08, 2002 11:20 AM To: TO, RAYMOND M; BAUMAN, DAVID N; CHISUM, MICHAEL R; DAIBER, BRYAN J; CHADBOURN, HARMON C: GRAY, BRIAN C; LEWIS, RAYMOND S

Subject:

Reactor Trip Transients Raymond.

The RCS Functional Specification Transient 1I (Rod Withdrawal Accident (upset condition) defines this condition as follows:

"The rod withdrawal accident occurs when one group of control rods is accidentally withdrawn from the core at the maximum rate when the reactor is operating at a low power, This transient is assumed to occur when the reactor is operating at 15% power. This causes a rapid rise In power znd a reactor trip."

In an effort to facilitate transient cycle counting, FTI performed a review of the functional specification transients and consolidated the above transient in what is currently tracked as Transient 119. This FTi report has been filed as Engineering Report 95R-1015-01 and Is FTI document number 51-1235146-01. In the FTI report, ANO historical translents though the year 1995 were reviewed and consolidated to reflect the current transient cycle recording document reflected in Procedure 1010.010. The transient cycle log procedure 1010.010 currently records reactor trip transients for the following conditions:

"£ Transient 17 - Reactor Trips with loss of RC flow. There have been 10 recorded transients in this category.

"* Transient 18 - Reactor Trips with Post Overcooling. There have been 0 transients recorded in this category. The 511811996 trip did not meet the overcooling requirements to be recorded in this category. It was captured in transient 19.

"* Transient 19 - All Other Reactor Trips. There have been 96 recorded transients In this ctegorythrough December 31,2001. There were two RPS high pressure trips recorded.

These high pressure trips did not cause a Pressurizer Code Safety valve relief with set point of 2500 psi.

The control room logs were reviewed in Autolog as well as PCRS with the plant computer data to assess related trips that may have met the criteria for Functional Specification Transient #11.

Based on this level of review, It can be stated that ANO-1 has not experienced any of the Functional Specification transient #11 events since commercial operation. Please be advised that the transients associated with plant performance for the 2002 operating year have not been finalized. However, it Is certain that none of the Functional Specification Transient #11 conditions have been experienced for ANO-1 for this year. Piease advise if there are any clarifications required for the data provided herewith or Ifany additional Input or research Is desired.

Bracy JaymotzdM. Tio"7 ANO-Design Engineering Structural Group L19

ENCLOSURE 2 CNRO-2002-00058 FRAMATOME DOCUMENT 32-5021539 (ANO CALCULATION 86-E-0074-161)

20697-6 (2/2002)

A FR:#A MAT)M E ANP CALCULATION

SUMMARY

SHEET (CSS)

I=AAOM N Document Identifier 32 - 5021539 - 01 Title ANO-1 CRDM NOZZLE IDTB WELD ANOMALY FLAW EVALUATIONS PREPARED BY: REVIEWED BY:

METHOD. 0 DETAILED CHECK E] INDEPENDENT CALCULATION NAME D.E. KILLIAN 4::ýk

./* . NAME A.D. NANA SIGNATURE /Mwu/ýý SIGNATURE DATE TITLE PRINCIPAL ENGR. DATE TITLE ADVISORY ENGR.

COST CENTER 41026 REF.

PAGE(S) 40,41 TM STATEMENT:

REVIEWER INDEPENDENCE )dDM PURPOSE AND

SUMMARY

OF RESULTS:

Revision 1: This revision is a non-proprietary version of Revision 0.

The purpose of this analysis is to perform a fracture mechanics evaluation of a postulated anomaly in the ANO Unit 1 CRDM nozzle ID temper bead weld. This anomaly is assumed to be a [ ] inch semi-circular flaw extending 360 degrees around the circumference at the "triple point" location where there is a confluence of three materials; the Alloy 600 nozzle, the Alloy [ ] weld, and low alloy steel head. Two potential flaw propagation paths are considered in the flaw evaluations. The analysis includes prediction of fatigue crack growth in an air environment since the anomaly is located on the outside surface of the new weld, just below the bottom of the severed nozzle. Flaw acceptance is based on the 1992 ASME Code Section XI criteria for applied stress intensity factor (IWB-3612) and limit load (IWB-3642).

The results of the analysis demonstrate that the [ ] inch weld anomaly is acceptable for a 25 year design life of the CRDM nozzle ID temper bead weld repair. Significant fracture toughness margins have been demonstrated for each of the two flaw propagation paths considered in the analysis. The minimum fracture toughness margin is 4.58, compared to the required margins of -410 for normal/upset conditions and "q2for emergency/faulted conditions per IWB-3612. Fatigue crack growth is minimal since the maximum final flaw size is less than 0.101 inch. The margin on limit load is 9.44 for normal/upset conditions and 6.67 for emergency/faulted conditions, compared to the required margins of 3.0 and 1.5, respectively, per IWB-3642.

THE FOLLOWING COMPUTER CODES HAVE BEEN USED IN THIS DOCUMENT: THE DOCUMENT CONTAINS ASSUMPTIONS THAT MUST BE VERIFIED PRIOR TO USE ON SAFETY RELATED WORK CODEVERSIONIREV CODENERSION/REV YES Ir NO Page I of 45

A "A A A' kA AKI- 32-5021539-01 rMFk1V1tA I %a M RECORD OF REVISIONS Affected Revision Paqes Description Date Original release 11/02 0 All Revision 1 is a non-proprietary version 12/02 1 All of Revision 0.

2

A 32-5021539-01 FRAMATOME ANP TABLE OF CONTENTS Section Title Page 1.0 INTRO DUCTIO N ....................................................................................................... 4 2.0 ASSUM PTIO NS ......................................................................................................... 5 3.0 W ELD ANO MALY ........................................................................................................ 6 4.0 MATERIAL PRO PERTIES .............................................................................................. 8 5.0 APPLIED STRESSES .................................................................................................. 11 6.0 FRACTURE MECHANICS METHODOLOGY .............................................................. 20 7.0 ACCEPTANCE CRITERIA ........................................................................................... 22 8.0 FLAW EVALUATIO NS ................................................................................................. 23 9.0 SUM MARY O F RESULTS ........................................................................................... 38 10.0 CO NCLUSIO N .................................................................................................................. 39

11.0 REFERENCES

................................................................................................................. 40

- Appendix Title Page A CERTIFICATIO N DO CUM ENT ................................................................................ 42 3

A 32-5021539-01 FRAMATOME ANP

1.0 INTRODUCTION

The CRDM nozzle ID temper bead weld repair is described by the design drawing (Reference 1).

"Thisweld repair establishes a new pressure boundary above the original J-groove weld. The five steps involved in the repair design are listed below.

1) Roll Expansion

2) Nozzle Removal and Weld Prep Machining

3) Welding

4) Grinding/Machining and NDE

5) Remediation and Original Weld Grinding During the welding process (step 3), a maximum [ ] inch weld anomaly may be formed due to lack of fusion at the "triple point", as shown in Figure 1. The anomaly is conservatively assumed to be a "crack-like" defect, 360 degrees around the circumference at the "triple point" location. The technical requirements document (Reference 2) provides additional details of the ID temper bead weld repair procedure. The purpose of the present fracture mechanics analysis is to provide justification, in accordance with Section XA of the ASME Code (Reference 3), for operating with the postulated weld anomaly at the triple point. Predictions of fatigue crack growth are based on a design life of 25 years.

4

A 32-5021539-01 FRAMATOME ANP 2.0 ASSUMPTIONS evaluation.

Listed below are assumptions that are pertinent to the present fracture mechanics location

1) The anomaly is assumed to include a "crack-like" defect, located at the triple-point a continuous and extending all the way around the circumference. For analytical purposes, Another circumferential flaw is located in the horizontal plane at the top of the weld.

plane between the weld and reactor vessel (RV) continuous flaw is located in the cylindrical head.

defect

2) In the radial plane, the anomaly is assumed to include a quarter-circular "crack-like" represent the radial (see Figure 1). For analytical purposes, a semi-circular flaw is used to cross-section of the anomaly.

An RTNDT value of [ I 'F is conservatively assumed for the [

] low 3) on a highest measured value of 40 "F for alloy reactor vessel head material. This is based 13 heats of SA-533 Grade B plate material (Reference 4).

5

A 32-5021539-01 FRAMATOME ANP 3.0 WELD ANOMALY The anomaly is located in the triple point region as shown in Figure 1 below.

1 TRIPLE 2 MAX POINT I.

20" MIN

( MAX POSSIBLE LACK OF FUSION ANOMALY)

AS-WELDED SURFACE SHALL BE SUITABLE FOR PT Figure 1. Weld Anomaly in Temper Bead Weld Repair The region is called a "triple point" since three materials intersect at this location. The materials are:

a) the Alloy 600 CRDM nozzle material, b) the new [ ] filler weld material,* and c) the low alloy steel RV head material.

Per Reference 7, Specification 5.14, Par. A7.4.3, "Filler metal of this classification is used for welding nickel-chromium-iron alloy (ASTM B163, B166, B167, and B168 having UNS Number [ ])." This UNS number is associated with Alloy [ ] material.

6

A FRAMATOME ANP 32-5021539-01 3.1 Postulated Flaws The triple point weld anomaly is assumed to be semi-circular in shape with an initial radius of

[ ]", as indicated in Figure 1. It is further assumed that the anomaly extends 3600 around the nozzle. Three flaws are postulated to simulate various orientations and propagation directions for the anomaly. A circumferential flaw and an axial flaw on the outside surface of the nozzle would both propagate in a horizontal direction toward the inside surface. A cylindrically oriented flaw along the interface between the weld and head would propagate downward between the two components. The horizontal and vertical flaw propagation directions are represented in Figure 2 by separate paths for the downhill and uphill sides of the nozzle, as discussed below.

For both these directions, fatigue crack growth will be calculated considering the most susceptible material for flaw propagation.

Horizontal Direction (Path 1):

Flaw propagation is across the CRDM tube wall thickness from the OD of the tube to the ID of the tube. This is the shortest path through the component wall, passing through the new Alloy [ ] weld material. However, Alloy 600 tube material properties or equivalent are used to ensure that another potential path through the HAZ between the new repair weld and the Alloy 600 tube material is bounded.

For completeness, two types of flaws are postulated at the outside surface of the tube. A 3600 continuous circumferential flaw, lying in a horizontal plane, is considered to be a conservative representation of crack-like defects that may exist in the weld anomaly.

This flaw would be subjected to axial stresses in the tube. An axially oriented semi circular outside surface flaw is also considered since it would lie in a plane that is normal to the higher circumferential stresses. Both of these flaws would propagate toward the inside surface of the tube.

Vertical Direction (Path 2):

Flaw propagation is down the outside surface of the repair weld between the weld and RV head. A continuous surface flaw is postulated to lie along this cylindrical interface between the two materials. This flaw, driven by radial stresses, may propagate along either the new Alloy [ ] weld material or the low alloy steel head material.

7

A FRAMATOME ANP 32-5021539-01 4.0 MATERIAL PROPERTIES The region of interest for the present flaw evaluations is at the triple point, where three different materials intersect. These materials are the CRDM nozzle material, the new weld material and the reactor vessel head material.

The ANO Unit 1 CRDM nozzles are made from Alloy 600 material to ASME specification SB from 167 for tubular products (Reference 2). The new weld, as noted in Section 3.0, is made Alloy [ ] type material. The portion of the reactor vessel head that contains the CRDM nozzles is fabricated from [ ] (Reference 2).

4.1 Yield Strength Values of yield strength, Sy, are obtained from the 1989 Edition of the ASME Code (Reference 9), as listed below.

1Low Alloy Steel Plate Material (RV Head)

Room temperature 50.0 ksi Operating temperature of 600 OF 43.8 ksi SB-163 Material [ 1 (used for Alloy r 1 Weld Metal)

Room temperature 40.0 ksi Operating temperature of 600 OF 31.1 ksi SB-167 Material N06600 (Alloy 600 Material)

Room temperature 35.0 ksi Operating temperature of 600 OF 27.9 ksi 8

A FRAMATOME ANP 32-5021539-01 4.2 Fracture Toughness 4.2.1. Low Alloy Steel RV Head Material Fracture toughness curves for [ ] material are illustrated in Figure A 4200-1 of Reference 3. At an operating temperature of about 600 OF, the KIa and K1c fracture toughness values for this material (using an assumed RTNDT of [ ] OF) are above 200 ksi/in. An upper bound value of 200 ksi'in wilt be conservatively used for the present flaw evaluations.

4.2.2. Alloy 600 and Alloy [ I Materials In Table 7 of Reference 12, Mills provides fracture toughness data for unirradiated Alloy 600 material at 24 0C (75 OF) and 427 0C (800 OF) in the form of crack initiation values for the J integral, J, Using linear interpolation and the LEFM plane strain relationship between Jc and fracture toughness, K c, Kjc J=

_2

ý1-v 2 the fracture toughness at an operating temperature of 600 OF is derived as follows:

Note: v = 0.3 1 kN/m = I kNfm + 4.448 N/lb x 0.0254 m/in = 0.00571 kip/in Mills [12] Code [9]

Temp. Jc Jc E K.c (F) (kN/m) (kip/in) (ksi) (ksiqin) 75 382 2.18 31000 273 600 522 2.98 28700 307 800 575 3.28 27600 316 Since brittle fracture is not a credible failure mechanism for ductile materials like Alloy 600 or Alloy I ], these fracture toughness measures, provided for information only, are not considered in the present flaw evaluations. However it should be noted that the fracture toughness measures of these ductile materials is significantly greater than the fracture toughness measure of the low alloy RV head material reported in Section 4.2.1. The failure mechanism for the ductile Alloy 600 and [ ] materials is limit load.

9

A 32-5021539-01 r, K /A M A-k I %j IVI r- J-A 114 r-4.3 Fatigue Crack Growth Flaw growth due to fatigue is characterized by da = Co(AK1 )n dN where C, and n are constants that depend on the material and environmental conditions, AKI is the range of applied stress intensity factor in terms of ksWin, and da/dN is the incremental flaw growth in terms of inches/cycle. For the embedded weld anomaly considered in the present analysis, it is appropriate to use crack growth rates for an air environment. Fatigue crack growth is also dependent on the ratio of the minimum to the maximum stress intensity factor; i.e.,

R = (KI)min / (Ki)max

[ Low Alloy Steel Plate Material (RV Head)

From Article A-4300 of the 1992 Edition of Section Xl (Reference 3), the fatigue crack growth constants for subsurface flaws in an air environment are:

n = 3.07 CO = 1.99 x 10.10 S where S = 25.72 ( 2.88 - R )-307 for 0<R<1 Alloy 600 and Alloy [ TMaterials (used for Alloy [ 1Weld Metal)

Fatigue crack growth rates for austenitic stainless steels are used to conservatively predict flaw growth in the new Alloy [ ] repair weld. From Article C-3210 of the 1992 Edition of Section Xl (Reference 3), the fatigue crack growth constants for subsurface flaws in an air environment are:

n = 3.3 Co= CxS 2 3 where C = 10[ -10. 0 0 9 + 8.12E-4xT - 1.13E-6xT + 1.02E-9xT ]

S= 1.0 for R5;0

= 1.0 + 1.8R for 0< R<0.79

= -43.35 + 57.97R for 0.79 < R < 1.0 10

A FRAMATOME ANP 32-5021539-01 5.0 APPLIED STRESSES The applied stresses are the cyclic stresses that contribute to fatigue crack growth. Incremental crack growth is based on six design heatup/cooldown cycles per year of operation. Residual stresses are also developed in the repair weld from the ID temper bead welding process that forms the new pressure boundary.

5.1 Fatigue Stresses Fatigue stresses are obtained from the stress analysis contained in Reference 6. The maximum stresses, which occur during cooldown (at 10.004 hours4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months into the composite heatup/cooldown transient), are combined with a zero stress at shutdown to produce a maximum cyclic load since stresses remain positive during this transient due to the dominating effect of pressure. The reactor coolant pressure at the 10.004 hour4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months time point is [ ] psig (Reference 6). A slightly higher pressure ([ ] psia) occurs during a rod withdrawal accident, which is classified as an upset condition in the reactor coolant system functional specification (Reference 19). Since ANO-1 has not experienced any of these transients during its plant life (Reference 20), the rod withdrawal accident is analyzed as an emergency condition transient, with no contribution to fatigue crack growth. This exception to the functional specification is noted in the document certification statement contained in Appendix A. Stresses for the rod withdrawal transient will be obtained by multiplying the stresses at 10.004 hours4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months into the composite heatup/cooldown transient by the ratio of the pressures for the two transients.

Component stresses are obtained for the two crack propagation paths outlined on the finite element model in Figure 2. Stresses for Paths 1 and 2 are obtained from Appendix D of Reference 6. Stresses are reported in a cylindrical coordinate system relative to the CRDM nozzle and include the three component stresses (axial, hoop and radial) needed to calculate mode I stress intensity factors for the various postulated flaws. These stresses, provided at four uniform increments along each path, were derived for ligament thicknesses of 0.488" for Path 1 and 1.143 inches for Path 2.

The stresses in Reference 6 apply directly to a weld thickness of 0.488". When the inside surface of the weld is finished by grinding, the thickness of the weld relative to the outside surface of the nozzle is

([ J"- [ ]") 12 = 0.518" (Reference 1)

Since the actual weld thickness is greater than the analyzed thickness, no adjustment will be made to the Reference 6 stresses in the present flaw evaluations.

To ensure that the bounding stresses are captured for use in the present flaw evaluations, stresses are obtained at every 45 degrees from the downhill (00) to the uphill (1800) locations, as shown by the figure in Appendix D of Reference 6. It is concluded in that reference that the most limiting path is at the 1800 uphill location. The uphill stresses are presented in Tables 1 and 2 for Paths I and 2, respectively.

As noted in the conclusions of Appendix F of Reference 6, stresses due to emergency/faulted conditions are bounded by the controlling normallupset condition stresses. Therefore, the emergency/faulted condition stresses are bounded by the normal/upset condition stresses, considered above, for the fatigue crack growth analysis.

11

A 32-5021539-01 FRAMATOME ANP

-I Figure 2. Illustration of Crack Propagation Paths on the Finite Element Stress Model 12

32-5021539-01 Framatome ANP Table 1. Stresses for Flaw Evaluations Along Path 1 (from Reference 6)

Composite HoatuplCooldown Transient (Normal and Upset Conditions)

Triple Point Path: WA180 Length = 0.488 0.244 0.360 0.488 0.488 0,488 Location: 0,000 0,000 0.000 0,122 0,122 0,122 0,244 0.244 0.366 0.366

--SX.- --SX-- .. SY-- --Sz-- -.SY-. --SZ- .-SX.. --Sy- --SZ.- --SX-- .-SY-o .-SZA Pressure Time --Sy-- .-Sz.. LUsn Avl I P-1rini I-nnn Axial (psig) I (hr.) I Radial Hoop Axial Radial Hoop Mial i r-%au a, woo r Axial Radial Hoon 0.001 4.770 4.871 7.000 7.313 7.412 10.000 (1114 1.

1(I 10004j1 10.013 10.1 17 10.217 10.250 10.718 12.939 Ratioed Stresses for Rod Withdrawl Accident (Emergency Condition)

I psig / [ I psig) *HeatuplCooldown Stress Triple Point Note: Rod Withdrawal Accident Stress = ([

0 244 0 244 0.366 0.366 0 366 0 488 0.488 0488 Pressure JLocation: 0.000 0.000 0.000 0.122 0.122 0.122 0.244 --SY- --sz-

--SZ-- ..SX-. --SY-- --Sz-- --sx.. --SY-- --SZ- --sX-* --SZ-- --SX-- .-SY--

-SX- -SY-- A :{-, D, A ,.I krnn Axial (psig) Radial Hoop Axial Radial Hoop Axial dial.

aUl nuuRa . V r" ..

Legend: SX = radial stress SY = hoop stress SZ = axial stress 13 Path 1 ANO-1 Fatigue Stresses NP.xls

32-5021539-01 Framatome ANP Table 2. Stresses for Flaw Evaluations Along Path 2 (from Reference 6)

Composite HeatuplCooldown Transient (Normal and Upset Conditions)

Triple Point Path: WV18O Length= 1.143 02858 0.8573 0 8573 1.1430 1.1430 1.1430 0 0000 0.2858 0.2858 0.5715 0.5715 05715 0 8573 --SY-Location: 00000 00000 --SX . --SY-- --Sz-- --SX-- --SZ-

-SZ-- -SX-- --SY-- --Sz-- --SZ-- --SX-- Hoon Pressure Time -.SX- --SY-- '-'Ann Axial

.p2 _(ha.)

0.001 Radial Hoop Axial Radial Hoop Axial Rauial Houup Axdl . r-f ...... ....... ....

4.770 4.871 7.000 7.313 7.412 10.000j 10.004 1 10.013 10.117 10.217 10.250 10.718 12.939 Ratioed Stresses for Rod Withdrawl Accident (Emergency Condition)

Triple Point Note: Rod Withdrawal Accident Stress = ([ Ipsig I I psig)

Heatup/Cooldown Stress 0.244 0.366 0 366 0 366 0.488 0.488 0488 0.000 0.000 0122 0.122 0.122 0244 0 244 -- S)...S -

--SY-- --Sz-- --sY-- --SZ-- --SX-- -SY-- -xSZ-Pressre I-SX-- --SY-- -SZ- --SX-- --SX-- --SY- Axial PressureI Location: 0.000 tj-*

(psig) I Radial Hoop Axial Radial Hoop Axial Radial Hoop AM1d, .... .... "..

I I Legend: SX = radial stress SY = hoop stress SZ = axial stress Path 2 ANO-1 Fatigue Stresses NP.xls 14

A FRAMATOME ANP 32-5021539-01 5.2 -Residual Stresses A three-dimensional elastic-plastic finite element analysis (Reference 5) was performed by Dominion Engineering, Inc. (DEI) to simulate the sequence of steps involved in arriving at the configuration of the CRDM nozzle and reactor vessel head after completion of the ID temper bead repair. To simply the analysis of the complete repair process, only the center nozzle was modeled (Figure 3). Although this axisymmetric analysis was based on the geometry of the center nozzle penetration, adjustments were made to represent significant aspects of the controlling nozzle at the outermost hillside location (38) from the top of the vessel). In particular, the repair weld was positioned at the minimum distance above the J-groove and the J-groove weld was chamfered to simulate the largest chamfer (7/8"). The model also used the highest yield strength of any nozzle in the head ([ ] ksi). The 38' nozzle location was limiting for all three of these conditions.

The DEI analysis simulated the laying of the original weld butter and the subsequent post-weld stress relief, the heatup of the original J-groove weld and adjacent material during the welding process and the subsequent cooldown to ambient temperature, a pre-service hydro test, and operation at steady state conditions. After the steady state loads were removed and the structure was again at ambient conditions, the portion of the nozzle below the cut line (Reference 1) was deleted. Deposition of the repair weld was simulated using four weld passes, and the J-groove weld was chamfered as shown in Figure 4. The analysis of this final configuration provided residual stresses in the repair weld for use in the present flaw evaluations. These stresses are listed in Table 3.

The repair weld analysis of Reference 5 used a multi-linear isotropic hardening model to characterize the nozzle material and elastic-perfectly plastic material models for the welds, butter, cladding and head. The yield strengths for the non-strain hardening models were selected to represent the flow stress of the various materials. The following yield strength values were used in the DEi repair weld analysis:

Yield Strength Component Material at 600 OF Nozzle Alloy 600 [ ] ksi Repair weld Alloy [ ] [ ] ksi J-groove weld Alloy 182 [ ] ksi

.Butter Alloy 182 [ ] ksi Head Low alloy steel [ ] ksi Cladding Stainless steel [ ] ksi 15

A 32-5021539-01 FRAMATOME ANP

%7,171v0Cut I ine ,.

2N)0 Nozle Repair ..

Weld Repin Nozzle Element-,

Reimned During Repair Weld Prep Node Numbers Increase h% 100 up the lenr'th of the tube and shell Node Numbers incicase by I alonp Ith tube and shell radius I Nodes 609 through 1409 are coincident with 610 through 1410 Figure 3. DEl Model for Center CRDM Nozzle with Weld Repair 16

A 32-5021539-01 FRAMATOME ANP N.ozzk Repair Weld Rei~on Region Reimwed for \Vcld Chami fer Figure 4. DEI Model for Center CRDM Nozzle After Weld Repair and Chamfer 17

Framatome ANP 32-5021539-01 Table 3 Residual Stresses in Repair Weld after Chamfering J-Weld Penetration angle = 0 degrees Nozzle yield strength = [ ] ksi File: ANO1-OB.results.txt (Reference 5)

Time: 16001 Path Along Interface Between Repair Weld and Remaining Nozzle (Corresponds to Path 1)

Radial Hoop Axial Coordinates Location Node Stress Stress Stress X Z (psi) (psi) (osi) i (in.) (in.)

,a

-i r r

I1 Triple Point 2609 (psi) 2.0000 90.525 2608 1.9228 90.570 2607 1.8456 90.615 2606 1.7684 90.659 2605 1.6912 90.704 2604 1.6141 90.748 2603 1.5369 90.793 2602 1.4597 90.837 Inside Surface 2601 1.3825 90.882 18

Framatome ANP 32-5021539-01 Table 3 (Cont'd)

Residual Stresses in Repair Weld after Chamfering J-Weld Penetration angle = 0 degrees Nozzle yield strength = [ ] ksi File: ANOI-OB.results.txt (Reference 5)

Time: 16001 Path Along Interface Between Repair Weld and Reactor Vessel Head (Corresponds to Path 2)

Stresses in Weld Radial Hoop Axial Coordinates Relative Node Stress Stress Stress X Z Position Location (psi)

(psi) (psi) (in.) (in.) (in.)

ii i "11 I f I[ i * ] [

m Triple Point 2609 (psi) 2.000 90.525 0.000 2509 2.000 90.302 0.223 2409 2.000 90.079 0.446 2.000 89.855 0.670 2309 2209 2.000 89.632 0.893 2109 2.000 89.409 1.116 2009 2.000 89.185 1.340 Lower End Stresses in Head Radial Hoop Axial Coordinates Relative Node Stress Stress Stress X Z Position Location (psi) (psi) (psi) (in.) (in.) (in.)

Triple Point 2610 2.000 90.525 0.000 2510 2.000 90.302 0.223 2410 2.000 90.079 0.446 2310 2.000 89.855 0.670 2210 2.000 89.632 0.893 2110 2.000 89.409 1.116 2010 2.000 89.185 1.340 Lower End 19

A 32-5021539-01 FRAMATOME ANP 6.0 FRACTURE MECHANICS METHODOLOGY This section presents several aspects of linear elastic fracture mechanics (LEFM) and limit load analysis (to address the ductile Alloy 600 and Alloy [ ] materials) that form the basis of the present flaw evaluations. As discussed in Section 3.1, flaw evaluations are performed for flaw propagation Paths I and 2 in Figure 2.

Path 1 represents a section across the new Alloy [ ] weld metal which is equivalent to the thickness of the CRDM tube wall. Since the weld anomaly is located at the base of the OD of the CRDM tube and is assumed to be all the way around the circumference, a stress intensity factor (SIF) solution for a 360 degree circumferential crack on the OD of a circular tube is deemed appropriate. Therefore, the SIF solution of Buchalet and Bamford (Reference 13) is used in the analysis. However, this solution is applicable to a 360-degree part-through ID flaw.

To develop an SIF solution for a 360 degree part-through OD flaw, an F function is determined based on SIF solutions of Kumar (References 14 and 15). Appropriate F functions for internal and external circumferential flaws are determined for a cylinder subjected to remote tension.

The ratio of the F functions for the external and internal flaws is considered to be an appropriate multiplying factor for the Buchalet and Bamford SIF solution to extend its application to an external flaw. Similar ratios have been reported by Kumar (Reference 18). The materials to be considered for this path are the Alloy 600 tube material or the Alloy [ ] weld metal. Fatigue crack growth is calculated using crack growth rates for austenitic stainless steels from Appendix C of Section XI (Reference 3). A limit load analysis for an external circumferential flaw in a cylinder subjected to remote tension (Reference 15) is also performed for applied loads on the CRDM tube.

An axially oriented semi-circular OD surface flaw is also considered in the evaluation, as illustrated by the schematic below.

Flaw Propagation Path Conpone p t Wott

-t Semi-Elliptical.

where, a = initial flaw depth = [ J inch I =2c = flaw length =[ ]inch t = wall thickness = 0.518 inch An axial flaw is considered since the stresses in the CRDM penetration region are primarily due to pressure and therefore the hoop stresses are more significant. The SIF solution by Raju &

Newman (Reference 10) for an external surface crack in a cylindrical vessel is used in the 20

A FRAMATOME ANP 32-5021539-01 evaluation, considering growth in both the radial and axial directions. The fatigue flaw growth analysis for the axial crack is also performed using the austenitic stainless steel crack growth rates.

The Irwin plasticity correction is also considered in the SIF solutions discussed above. This plastic zone correction is discussed in detail in Section 2.8.1 of Reference 11. The effective crack length is defined as the sum of the actual crack size and the plastic zone correction:

ae =a+ry where ry for plane strain conditions (applicable for this analysis) is given by:

ry=-1 KI-6Tc a ys)

Path 2 represents the interface between the new repair weld and the RV head material. The potential for flaw propagation along this interface is likely if radial stresses are significant between the weld and head. This assessment utilizes an SIF solution for a continuous surface crack in a flat plate from Appendix A of the 1995 Edition of Section XI (Reference 16). Flat plate solutions are routinely used to evaluate flaws in cylindrical components such as the repair weld since the added constraint provided by the cylindrical structure reduces the crack opening displacements. The solution is therefore inherently conservative for this application. Crack growth analysis is performed considering propagation through the Alloy [ ] weld metal or the low alloy steel head material, whichever is limiting.

21

A FRAMATOME ANP 32-5021539-01 7.0 ACCEPTANCE CRITERIA The low alloy steel reactor vessel head material will be evaluated against the IWB-3612 acceptance criteria of Section Xl (Reference 3). For the highly ductile materials Alloy 600 and Alloy

] materials, the initial flaw depth to thickness ratio for the postulated weld anomaly is only about 20% and fatigue crack growth is minimal for these materials in an air environment. A convenient acceptance criterion on flaw size is the industry developed 75% through-wall limit on depth (Reference 8):

aa <!0.75 t

For the shallow cracks considered in the present analysis, this criterion is easily met. In addition, stress intensity factors will be calculated and evaluated against conservative fracture toughness requirements using a factor of safety of 410 for normal and upset conditions.

Another acceptance criterion for ductile materials is demonstration of sufficient limit load margin.

From IWB-3642 (Reference 3), the required safety margin, based on load, is a factor of 3 for normal and upset conditions and a factor of 1.5 for emergency and faulted conditions.

Since stresses for emergency/faulted conditions are bounded by the controlling normal/upset condition stresses (see Section 5.0) and the required fracture toughness margins are less stringent for emergency/faulted conditions, satisfying normal/upset conditions requirements implicitly satisfies those for emergency/faulted conditions as well.

22

A FRAMATOME ANP 32-5021539-01 8.0 FLAW EVALUATIONS The evaluation of the postulated external circumferential flaw for propagation along Path I is contained in Tables 4 and 5. The fatigue crack growth analysis is provided in Table 4 and a limit load analysis is presented in Table 5.

The evaluation of an external axial flaw for fatigue crack growth along Path 1 is contained in Table 6.

A continuous surface flaw along the cylindrical interface between the repair weld and the reactor vessel head is analyzed for fatigue crack growth along Path 2 in Table 7.

The flaw evaluations utilize the combined stresses resulting from the sum of the residual stresses from Reference 3 and the fatigue stresses from Reference 6. This is a conservative approximation of the actual state of stress since the elastic fatigue stresses are added directly to the elastic-plastic residual stresses, with no attenuation for additional plastic strain. It is therefore appropriate to use the yield strengths from the DEI stress analysis (Reference 3) when applying the Irwin plastic zone correction for crack length.

As required by Article IWB-3612 (Reference 3), a safety factor of '110 is used to evaluate applied stress intensity factors for normal and upset conditions, considering the lower Kl, fracture toughness for crack arrest. Article IWB-3612 also specifies that a safety factor of "42 must be used for emergency and faulted conditions, along with the higher K1, fracture toughness for crack initiation. Since the required safety margin for the emergency condition rod withdrawal accident is less than that for normal and upset conditions by a factor of 410 1 42 = 2.24 and the rod withdrawal accident stresses from Tables 1 and 2 are only [ ] psig / [ ] psig, or 1.05 times the maximum normal and upset condition stresses, the flaw evaluations performed for normal and upset conditions serve as a bounding analysis for the emergency condition rod withdrawal accident.

23

Framatome ANP 32-5021539-01 Table 4. Evaluation of Continuous External Circumferential Flaw for Fatigue Crack Growth Along Path I INPUT DATA Geometry: Outside diameter, Doo[

Di =

] in.in.

Inside diameter, Thickness, t= 0.518 in.

Ri/t = 2.861 Flaw Size: Flaw depth, a*[ ]in alt =

Temperature, T= 600F Environment ANO-1 Circ Flaw NP.xls 24 Circ. Input

32-5021539-01 Framatome ANP Table 4. Evaluation of Continuous External Circumferential Flaw for Fatigue Crack Growth Along Path I (Cont'd)

Variation of F Function between Continuous External and Continuous Internal Circumferential Flaws Using Solutions by V. Kumar et al.

Source: EPRI NP-1931 Topical Report, Section 4.3 for an internal circumferential crack under remote tension (Ref. 14).

The applied KI equation is given by the expression:

KI = o*q(n*a)*F(a/b, Ri/Ro) where a= P/(ir*(ROA2 - RiA2) and F is a function of a/b and b/Ri, where a/b =F b/Ri = L By extrapolation from Table 4-5 of EPRI-1 931, the internal F-factor is estimated to be:

Ftntemai = 1.14 Source: GE Report SRD-82-048, Prepared for EPRI Contract RP-1237-1, Fifth & Sixth Semi-Annual Report, Section 3.5 for an external circumferential under remote tension (Ref. 15).

For the external circumferential crack, the expressions for KI and a are as defined above for the internal circumferential crack, where a/b= F Ri/Ro =EL From Figure 3-11 of SRD-82-048, the external F-factor is estimated to be:

FextemaI " 1.25 Multiplying Factor:

To estimate the stress intensity factor for an external circumferential crack from the solution for an internal circumferential crack under remote tension, the appropriate multiplying factor is:

Fextyrt / Fintemat = 1.25 /1.14 = 1.10 An alternate source for this multiplying factor is EPRI NP-3607 (Ref. 18). From Figure 3-9. the multiplying factor for circumferential flaws with an a/t ratio of 0.2 can also be estimated to be:

Fextý.ma I Fntma - 1.10 ANO-1 Circ Flaw NP.xls 25 SIF Factor

Framatome ANP 32-5021539-01 Table 4. Evaluation of Continuous External Circumferential Flaw for Fatigue Crack Growth Along Path 1 (Cont'd)

STRESS INTENSITY FACTOR FOR CIRCUMFERENTIAL FLAW B3sis: Buchalet and Bamford solution for continuous circumferential flaws on the inside surface of cylinders (Ref. 13)

KI = ",(n*a) * [ A0 F, + (2afn) A1 F2 + (a2 /2) A2 F 3 + (4a 3 )/(37c) A3 F 4 ]

+ 2.2018(a/t)' - 0.2083(a/t)3 where, F1 = 1.1259 + 0.2344(a/t)

+ 0.6354(a/t)3 F2 = 1.0732 + 0.2677(a/t) + 0.6661(a/t)2 2

+ 0.6042(a/t)3 F3 = 1.0528 + 0.1065(a/t) + 0.4429(a/t)

+ 0.3750 (a/t)3 F4 = 1.0387 - O.0939(a/t) + 0.6018(a/t)2 and the through-wall stress distribution is described by the third order polynomial, S(x) =A 0 + Ax + A2z8 + A3x 3.

Applicability: Rit = 10 a/t < 0.8 Axial Stresses:

Wall Residual Normal/Upset Cond. Total Stresses Position Stress Stresses [61 at Operation x in Weld Cooldown Shutdown Cooldown Shutdown (in.) (ksi) (ksi) (ksi) (ksi) (ksi) 0.00000 0.12950 0.25900 0.38850 0.51800 Stress Coefficients:

Normal/Upset Stress Loading Conditions Coeff. NU1 NU2 (ksi) (ksi)

Ak A,

A2 A3 26 Circ. KI ANO-1 Circ Flaw NP.xls

32-5021539-01 Framatome ANP Table 4, Evaluation of Continuous External Circumferential Flaw for Fatigue Crack Growth Along Path 1 (Cont'd)

CRACK GROWTH FOR CIRCUMFERENTIAL FLAW (IN-AIR) - AUSTENITIC MATERIAL Basis: Aa = AN

Co(AKI)" AN = 6 fatigue cycles / year Sy = 27.9 ksl NU1 NU2 NU1 Operating ae Kl(aj)max Time Cycle a KI(a)max KI(a)min AKI R S Co Aa (in.) (ksiqin)

(yr.) (ksi/in) (ksi*/in) (ksiqin) 1.96E-10 1.52E-05 -21.50 0 0 -18.01 -35.63 17.62 0.00 1.00 0,0212 0.00 1.00 1.96E-10 1.52E-05 0.0212 -21.51 1 6 -18.01 -35.64 17.62 0.00 1.00 1.96E-10 1.52E-05 0.0212 -21.51 2 12 -18.01 -35.64 17.62 0.00 1.00 1.96E-10 1.52E-05 0.0212 -21.51 3 18 -18.02 -35.64 17.63 0.00 1.00 1.96E-10 1.52E-05 0.0212 -21.51 4 24 -18.02 -35.65 17.63 0.00 1.00 1.96E-10 1.52E-05 0.0212 -21.52 5 30 -18.02 -35.65 17.63 1.00 1.96E-10 1.52E-05 0.0212 -21.52 6 36 -18.03 -35.65 17.63 0.00 0.00 1.00 1.96E-10 1.52E-05 0.0212 -21.52 7 42 -18.03 -35.66 17.63 0.00 1.00 1.96E-10 1.52E-05 0.0212 -21.52 8 48 -18.03 -35.66 17.63 0.00 1.00 1.96E-10 1.52E-05 0.0212 -21.53 9 54 -18.03 -35.66 17.63 0.00 1.00 1.96E-10 1.52E-05 0.0212 -21.53 10 60 -18.04 -35.67 17.63 0.00 1.00 1.96E-10 1.52E-05 0.0212 -21.53 11 66 -18.04 -35.67 17.63 0.00 1.00 1.96E-10 1.52E-05 0.0212 -21.53 12 72' -18.04 -35.68 17.63 0.00 1.00 1.96E-10 1.52E-05 00212 -21.54 13 78 -18.04 -35.68 17.64 0.00 1.00 1.96E-10 1.52E-05 0.0212 -21.54 14 84 -18.05 -3568 17.64 1.00 1.96E-10 1,52E-05 0.0212 -21.54 15 90 -18.05 -35.69 17.64 0.00 0.00 1.00 1.96E-10 1.52E-05 0.0212 -21.55 16 96 -18.05 -35.69 17.64 1.00 1.96E-10 1.52E-05 0.0212 -21.55 102 -18.05 -35.69 17.64 0.00 17 1.53E-05 0.0212 -21.55

-18.06 -35.70 17.64 0.00 1.00 1.96E-10 18 108 0.0212 -21.55

-18.06 -35.70 17.64 0.00 1.00 1.96E-10 1.53E-05 19 114 0.0212 -21.56

-18.06 -35.70 17.64 0.00 1.00 1.96E-10 1.53E-05 20 120 0.0212 -21.56

-18.07 -35.71 17.64 0.00 1.00 1.96E-10 1.53E-05 21 126 0.0212 -21.56

-18.07 -35.71 17.64 0.00 1.00 1.96E-10 1.53E-05 22 132 0.0212 -21.56

-18.07 -35.72 17.64 0.00 1.00 1.96E-10 1.53E-05 23 138 0.0212 -21.57

-18.07 -35.72 17.65 0.00 1.00 1.96E-10 1.53E-05 24 144 0.0212 -21.57

-18.08 -35.72 17.65 0.00 1.00 1.96E-10 1.53E-05 25 150 Circ. Growth ANO-1 Circ Flaw NP.xls 27

Framatome ANP 32-5021539-01 Table 5. Limit Load Analysis for a Continuous External Circumferenital Flaw LIMIT LOAD Basis: GE Report SRDo-82-048, Combined Fifth and Sixth Semi-Annual Report by V. Kumar et al, Section 3.5 (Ref. 15).

For remote tension loading, Po = 2/3*ao*,E*(Rc2-Ri2) where Rc = Ro - a and co = 27900 psi (conservatively using the minimum yield strength)

Ro =in.

a = in.

Rc = in.

Ri = in.

Then Pc= E- Ibs From Reference 17, the applied loads on a typical B&W design CRDM tube are:

a) Normal/Upset conditions, P= ] lbs b) Emergency/Faulted conditions, P =L lbs The limit load margins are greater than those required by Article IWB-3642 of Section XI (Ref. 3), as shown below.

a) Normal/Upset conditions, Po/P = 9.44 > 3.0 b) Emergency/Faulted conditions, Po/P = 6.67 > 1.5 ANO-1 Circ Flaw NP.xls 28 Circ. Limit Load

Framatome ANP 32-5021539-01 Table 6. Evaluation of an External Axial Flaw for Fatigue Crack Growth Along Path I INPUT DATA Geometry: Outside diameter, Inside diameter, Doo=

Di =

in.

Thickness, t= 0.518 Ri/t = 2.861 I

Flaw Size: Flaw depth, in.

Flaw length, in.

Environment: Temperature, T= 600 F ANO-1 Axial Flaw NP.xls 29 Axial Flaw Input

Framatome ANP 32-5021539-01 Table 6. Evaluation of an External Axial Flaw for Fatigue Crack Growth Along Path I (Cont'd)

STRESS INTENSITY FACTOR FOR AXIAL FLAW Basis: Raju & Newman, "Stress Intensity Factors for Internal & External Surface Cracks in Cylindrical Vessels (Ref. 10) 5 5 25 + G 3 A 3 a3 ]

KI = [(-,T/Q) * [Go Ao ao +G1 A1 a1 +G 2 A 2 a where, from Table 4 of Reference 10, for an external surface crack with tIR = 0.25, alt = 0.2, a/c = 1.0, the influence coefficients are as follows:

Location: Deepest Point Surface (2ýIn = 1) (24ht= 0)

Go 1.030 1.163 G= 0.720 0.204 G2 = 0.591 0.077 G3 = 0.513 0.040 and Q= 2.464 = (1 + 1.464*(a/c)A1.65) and the through-wall stress distribution is described by the third order polynomial, S(x) = Ao + A1x + A2x2 + A3X3.

Hoop Stresses:

Wall Residual Normal/Upset Cond. Total Stresses Position Stress Stresses [61 at Operation x in Weld Cooldown Shutdown Cooldown Shutdown (in.) (ksi) (ksi) (ksi) (ksi) (ksi) 0.00000 0.12950 0.25900 0.38850 0.51800 Stress Coefficients:

Normal/Upset Stress Loading Conditions Coeff. NUI NU2 (ksi) (ksi)

Ak A1 A2 A3 ANO-1 Axial Flaw NP.xls 30 Axial Flaw Ki

32-5021539-01 Framatome ANP Table 6, Evaluation of an External Axial Flaw for Fatigue Crack Growth Along Patti 1 (Cont'd)

RADIAL CRACK GROWTH FOR AXIAL FLAW (IN-AIR) - AUSTENITIC MATERIAL Basis: Aa = AN

Co(AKI)" AN = 6 fatigue cycles / year Sy = 42.5 ksi NU1 NU2 NU1 Operating a, KI(aj)max a KI(a)max KI(a)min AKI R S C &Aa Time Cycle (ksl4ln)

(yr.) (kslqln) (ksilin) (in.)

(In.) (ksi4ln) 28.72 9.19 16.62 1.64 3.21E-10 2.06E-05 0.0196 0.0196 0 0 25.80 0.36 28.73 1 25.81 9.19 16.62 0.36 1.64 3.21E-10 2.06E-05 0.0196 6 28.73 25.81 9.19 16.62 0.36 1.64 3.21E-10 2.06E-05 0.0196 2 12 28.73 25.81 9.19 16.62 0.36 1.64 3.21E-10 2.06E-05 0.0196 3 18 28.74 25.82 9.19 16.62 0.36 1.64 3.21E-10 2.06E-05 0.0196 4 24 28.74 25.82 9.19 16.63 0.36 1.64 3.21E-10 2.06E-05 0.0196 5 30 28.74 25.82 9.20 16.63 0.36 1.64 3.21E-10 2.06E-05 0.0196 6 36 28.75 9.20 16.63 0.36 1.64 3.2IE-10 2.06E-05 0.0196 7 42 25.83 28.75 25.83 9.20 16.63 0.36 1.64 3.21E-10 2.06E-05 0.0196 8 48 28.76 25.83 9.20 16.63 0.36 1.64 3.21E-10 2.06E-05 0.0196 9 54 28.76 25.84 9.20 16.63 0.36 1.64 3.21E-10 2.06E-05 0.0196 10 60 28.76 25.84 9.20 16.63 0.36 1.64 3.21E-10 2.06E-05 0.0196 11 66 28.77 25.84 9.21 16.63 0.36 1.64 3.21E-10 2.06E-05 0.0196 12 72 28.77 25.84 9.21 16.64 0.36 1.64 3.21E-10 2.06E-05 0.0196 13 78 28.77 25.85 9.21 16.64 0.36 1.64 3.21E-10 2.06E-05 0.0196 14 84 28.78 25.85 9.21 16.64 0.36 1.64 3.21E-10 2.06E-05 0.0196 15 90 28.78 25.85 9.21 16.64 0.36 1.64 3.21E-10 2.07E-05 0.0196 16 96 28.79 9.21 16.64 0.36 1.64 3.21E-10 2.07E-05 0.0196 17 102 25.86 28.79 25.86 9.22 16.64 0.36 1.64 3.21E-10 2.07E-05 0.0196 18 108 28.79 25.86 9.22 16.64 0.36 1.64 3.21E-10 2.07E-05 0.0196 19 114 28.80 25.87 9.22 16.65 0.36 1.64 3.21E-10 2.07E-05 0.0197 20 120 2.07E-05 28.80 25.87 9.22 16.65 0.36 1.64 3.21E-10 0.0197 21 126 28.81 25.87 9.22 16.65 0.36 1.64 3.21E-10 2.07E-05 0.0197 22 132 28.81 25.87 9.22 16.65 0.36 1.64 3.21E-10 2.07E-05 0.0197 23 138 28.81 25.88 9.23 16.65 0.36 1.64 3.21E-10 2.07E-05 0.0197 24 144 28.82 9.23 16.65 0.36 1.64 3.21E-10 2.07E-05 0.0197 25 150 25.88 31 Radial Growth of Axial Flaw ANO-1 Axial Flaw NP.xls

Framatome ANP 32-5021539-01 Table 6. Evaluation of an External Axial Flaw for Fatigue Crack Growth Along Path I (Cont'd)

AXIAL CRACK GROWTH FOR AXIAL FLAW (IN-AIR) - AUSTENITIC MATERIAL Basis: Aa = AN

Co(AKI)" AN = 6 fatigue cycles / year Sy = 42.5 ksi Operating NU1 NU2 NU1 KI(a)max KI(a)min AKI S C. Aa a0 KI(a,)max Time Cycle a R S(yr.) (ksi-4in) (ksiqin) (in.) (kslqin)

(In.) (ksiqin) 9.57 20.03 0.32 1.58 3.10E-10 3.67E-05 0.0257 33.12 0 0 29.61 9.58 20.04 0.32 1.58 3.10E-10 3.67E-05 0.0258 33.12 1 6 29.61 29.62 9.58 20.04 0.32 1.58 3.10E-10 3.68E-05 0.0258 33.13 2 12 29.62 9.58 20.04 0.32 1.58 3.10E-10 3.68E-05 0.0258 33.13 3 18 29.63 9.58 20.05 0.32 1.58 3.10E-10 3.68E-05 0.0258 33.14 4 24 20.05 0.32 1.58 3.10E-10 3.68E-05 0.0258 33.15 5 30 29.63 9.58 20.05 0.32 1.58 3.10E-10 3.68E-05 0.0258 33.15 6 36 29.64 9.59 20.06 0.32 1.58 3.10E-10 3.69E-05 0.0258 33.16 7 42 29.65 9.59 20.06 0.32 1.58 3.1OE-10 3.69E-05 0.0258 33.16 8 48 29.65 9.59 20.07 0.32 1.58 3.10E-10 3.69E-05 0.0258 33.17 9 54 29.66 9.59 9.59 20.07 0.32 1.58 3.10E-10 3.69E-05 0.0258 33.18 10 60 29.66 20.07 0.32 1.58 3.10E-10 3.70E-05 0.0258 33.18 11 66 29.67 9.59 20.08 0.32 1.58 3.10E-10 3.70E-05 0.0259 33.19 12 72 29.67 9.60 20.08 0.32 1.58 3.10E-10 3.70E-05 0.0259 33.19 13 78 29.68 9.60 20.08 0.32 1.58 3.10E-10 3.70E-05 0.0259 33.20 14 84 29.68 9.60 20.09 0.32 1.58 3 10E-10 3.70E-05 0.0259 33.21 15 90 29.69 9.60 20.09 0.32 1.58 3.10E-10 3.71E-05 0.0259 33.21 16 96 29.69 9.60 20.09 0.32 1.58 3.10E-10 3.71E-05 0.0259 33.22 17 102 29.70 9.60 20.10 0.32 1.58 3.10E-10 3.71E-05 0.0259 33.22 18 108 29.70 9.61 20.10 0.32 1.58 3.10E-10 3.71E-05 0.0259 33.23 19 114 29.71 9.61 20.10 0.32 1.58 3.10E-10 3.72E-05 0.0259 33.24 20 120 29.71 9.61 20.11 0.32 1.58 3.10E-10 3.72E-05 0.0259 33.24 21 126 29.72 9.61 20.11 0.32 1.58 3.10E-10 3.72E-05 0.0260 33.25 22 132 29.73 9.61 20.11 0.32 1.58 3.10E-10 3.72E-05 0.0260 33.25 23 138 29.73 9.62 20.12 0.32 1.58 3.10E-10 3.72E-05 0.0260 33.26 24 144 29.74 9.62 20.12 0.32 1.58 3.10E-10 3.73E-05 0.0260 33.27 25 150 29.74 9.62 ANO-1 Axial Flaw NP.xls 32 Axial Growth of Axial flaw

Framatome ANP 32-5021539-01 Table 7. Evaluation of a Continuous Cylindrical Surface Crack for Fatigue Crack Growth Along Path 2 INPUT DATA Geometry. Plate thickness, t= 1.143 in.

Flaw Size: Flaw depth, a=

alt =

I in.

Temperature, T- 600 F Environment:

ANO-1 Cylind Flaw NP.xls 33 Cylind. Input

Framatome ANP 32-5021539-01 Table 7. Evaluation of a Continuous Cylindrical Surface Crack for Fatigue Crack Growth Along Path 2 (Cont'd)

STRESS INTENSITY FACTOR FOR CYLINDRICAL FLAW IN WELD Basis: Analysis of Flaws, 1995 ASME Code,Section XI, Appendix A (Ref. 16)

KI = [Ac Go + A1 G, + A2 G 2 + A3 G 3 1] I(7aIQ) where Q = 1 + 4.593*(afl)AI .65 - qy and qy = [ (A0 Go + A, G1 + A2 G2 + A3 G3) / ays ]2 / 6 For a/ = 0.0 (continuous flaw) a/t.<= 0.1 Go = 1.1945 G3 = 0.7732 G2 = 0.5996 G3 = 0.5012 Stresses are described by a third order polynomial fit over the flaw depth, S(x) = A0 + Al(X/a) + A2(x/a)2 + A3(x/a) 3 Radial Stresses in WelId:

Wall Residual Normal/Upset Cond. Total Stresses Position Stress Stresses [61 at Operation x in Weld Cooldown Shutdown Cooldown Shutdown (in.) (ksi) (ksi) (ksi) (ksi) (ksi) 0.000 0.223 0.446 0.670 0.893 Stress Coefficients: (a= 0.100 in.)

Normal/Upset Stress Loading Conditions Coeff. NU1 NU2 (ksi) (ksi)

Aa A,

Az A3 ANO-1 Cylid Flaw NP.xls 34 Weld KI

.1 32-5021539-01 Framatome ANP Table 7. Evaluation of a Continuous Cylindrical Surface Crack for Fatigue Crack Growth Along Path 2 (Cont'd)

CRACK GROWTH FOR CYLINDRICAL FLAW (IN-AIR) - AUSTENITIC MATERIAL Basis: Aa = AN

Co(AKI)n AN = 6 cycles/year Sy = 69.0 ksl Operating NU1 NU2 Cycle Q Kl(a)max KI(a)min AKI R S Co Aa qy 0(aj) Kl(ae)max Time a (in.) (ksi'in)

(yr.) (ksilin) (ksiqin) (ksl/in)

Sin., 2.14 4.18E-10 1.76E-05 0.175 0.825 43.62 0 1.000 39.62 25.00 14.63 0.63 0 4.18E-10 1.76E-05 0.175 0.825 43.63 1 1.000 39.63 25.00 14.63 0,63 2.14 6 1.76E-05 0.175 0.825 43.63 1.000 39.63 25.00 14.63 0.63 2.14 4.18E-10 2 12 1.76E-05 0.175 0.825 43.64 1.000 39.64 25.01 14.63 0.63 2.14 4.18E-10 3 18 0.175 0.825 43.64 1.000 39.64 25.01 14.63 0.63 2.14 4.18E-10 1.76E-05 4 24 0.175 0.825 43.64 1.000 39.64 25.01 14.63 0.63 2.14 4.18E-10 1.76E-05 5 30 1.76E-05 0.175 0.825 43.65 1.000 39.65 25.01 14.63 0.63 2.14 4.18E-10 6 36 0.175 0.825 43.65 1.000 39.65 25.01 14.63 0.63 2.14 4.18E-10 1.76E-05 7 42 1.76E-05 0.175 0.825 43.65 1.000 39.65 25.02 14.64 0.63 2.14 4.18E-10 8 48 0.175 0.825 43.66 1.000 39.66 25.02 14.64 0.63 2.14 4.18E-10 1.76E-05 9 54 1.76E-05 0.175 0.825 43.66 1.000 39.66 25.02 14.64 0.63 2.14 4.18E-10 10 60 0.175 0.825 43.67 1.000 39.66 25.02 14.64 0.63 2.14 4.18E-10 1.76E-05 11 66 0.175 0.825 43.67 1.000 39.67 25.03 14.64 0.63 2.14 4.18E-10 1.76E-05 12 72 0.175 0.825 43.67 39.67 25.03 14.64 0.63 2.14 4.18E-10 1.76E-05 13 78 1.000 0.825 43.68 39.67 25.03 14.64 0.63 2.14 4.18E-10 1.76E-05 0.175 14 84 1.000 0.825 43.68 1.000 39.68 25.03 14.65 0.63 2.14 4.18E-10 1.76E-05 0.175 15 90 0.175 0.825 43.69 39.68 25.03 14.65 0.63 2.14 4.18E-10 1.76E-05 16 96 1.000 0.825 43.69 39.68 25.04 14.65 0.63 2.14 4.18E-10 1.76E-05 0.175 17 102 1.000 0.825 43.69 1.000 39.69 25.04 14.65 0.63 2.14 4.18E-10 1.76E-05 0.175 18 108 0.175 0.825 43.70 1.000 39.69 25.04 14.65 0.63 2.14 4.18E-10 1.76E-05 19 114 0.175 0.825 43.70 1.000 39.69 25.04 14.65 0.63 2.14 4.18E-10 1.77E-05 20 120 1.77E-05 0.175 0.825 43.70 1.000 39.70 25.04 14.65 0.63 2.14 4.18E-10 21 126 0.175 0.825 43.71 1.000 39.70 25.05 14.65 0.63 2.14 4.18E-10 1.77E-05 22 132 1.77E-05 0.175 0.825 43.71 1.000 39.70 25.05 14.66 0.63 2.14 4.18E-10 23 138 1.77E-05 0.175 0.825 43.72 1.000 39.71 25.05 14.66 0.63 2.14 4.181-10 24 144 0.825 43.72 39.71 25.05 14.66 0.63 2.14 4.18E-10 1.77E-05 0.175 25 150 1.000 35 Weld FCG ANO-1 Cylind Flaw NP.xls

Framatome ANP 32-5021539-01 Table 7. Evaluation of a Continuous Cylindrical Surface Crack for Fatigue Crack Growth Along Path 2 (Cont'd)

STRESS INTENSITY FACTOR FOR CYLINDRICAL FLAW IN HEAD Basis: Analysis of Flaws, 1995 ASME Code,Section XI, Appendix A (Ref. 16)

KI = [ A0 Go + A1 G1 + A2 G 2 + A3 G 3 ] 'I(7taIQ) where o = 1 + 4.593*(a/l)AI.65 - qy and qy = [ (Ao Go + A1 G1 + A2 G2 + A3 G3) / ys ]2 /6 For a/I = 0.0 (continuous flaw) a/t <= 0.1 Go = 1.1945 G, = 0.7732 G2 = 0.5996 G3 = 0.5012 Stresses are described by a third order polynomial fit over the flaw depth, 2 3 S(x) = A0 + Al(x/a) + A2(xla) + A3(x/a)

Radial Stresses in We Id:

Wall Residual Normal/Upset Cond. Total Stresses Position Stress Stresses [6) at Operation x in Weld Cooldown Shutdown Cooldown Shutdown (in.) (ksi) (ksi) (ksi) (ksi) (ksi) 0.000 0.223 0.446 0.670 0.893 Stress Coefficients: (a= 0.100 in. )

Normal/Upset Stress Loading Conditions Coeff. NU1 NU2 (ksi) (ksi)

A1 A2 A3 ANO-1 Cylind Flaw NP.xls 36 Head KI

Framatome ANP 32-5021539-01 Table 7, Evaluation of a Continuous Cylindrical Surface Crack for Fatigue Crack Growth Along Path 2 (Conrd)

CRACK GROWTH FOR CYLINDRICAL FLAW (IN-AIR) - FERRITIC MATERIAL Basis: Aa = AN

Co(AKI)" AN = 6 cycles/year Sy = 57.6 ksl Operating NUI NU2 Time Cycle a Q Kl(a)max KI(a)min AKI R S C0 Aa qy Q(ae) KI(ae)max (yr.) inI (ksiin) (ksiin) (ksiqin) (in.) (ksiqin) 38.08 23.46 14.63 0.62 2.09 4.16E-10 9.43E-06 0.232 0.768 43.45 0 0 1.000 1 38.08 23.46 14.63 0.62 2.09 4.16E-10 9.44E-06 0.232 0.768 43.45 6 1.000 38.09 23.46 14.63 0.62 2.09 4.16E-10 9.44E-06 0.232 0.768 43.46 2 12 1.000 38.09 23.46 14.63 0.62 2.09 4.16E-10 9.44E-06 0.232 0.768 43.46 3 18 1.000 1.000 38.09 23.46 14.63 0.62 2.09 4.16E-10 9.44E-06 0.232 0.768 43.46 4 24 38.09 23.46 14.63 0.62 2.09 4.16E-10 9.44E-06 0.232 0.768 43.46 5 30 1.000 38.09 23.46 14.63 0.62 2.09 4.16E-10 9.44E-06 0.232 0.768 43.46 6 36 1.000 38.09 23.46 14.63 0.62 2.09 4.16E-10 9.44E-06 0.232 0.768 43.47 7 42 1.000 1.000 38.10 23.46 14.63 0.62 2.09 4.16E-10 9.44E-06 0.232 0.768 43.47 8 48 38.10 23.47 14.63 0.62 2.09 4.16E-10 9.45E-06 0.232 0.768 43.47 9 54 1.000 38.10 23.47 14.63 0.62 2.09 4.16E-10 9.45E-06 0.232 0.768 43.47 10 60 1.000 38.10 23.47 14.63 0.62 2.09 4.16E-10 9 45E-06 0.232 0.768 43.47 11 66 1.000 1.000 38.10 23.47 14.63 0.62 2.09 4.16E-10 9.45E-06 0.232 0.768 43.48 12 72 38.11 23.47 14.63 0 62 2.09 4.16E-10 9.45E-06 0.232 0.768 43.48 13 78 1.000 1.000 38.11 23.47 14.64 0.62 2.09 4.16E-10 9.45E-06 0 232 0.768 43.48 14 84 1.000 38.11 23.47 14.64 0.62 2.09 4.16E-10 9.45E-06 0.232 0.768 43.48 15 90 38.11 23.47 14.64 0.62 2.09 4.16E-10 9.46E-06 0.232 0.768 43.48 16 96 1.000 1.000 38.11 23.47 14.64 0.62 2.09 4.16E-10 9.46E-06 0.232 0.768 43.49 17 102 1.000 38.11 23.48 14.64 0.62 2.09 4.16E-10 9.46E-06 0.232 0.768 43.49 18 108 38.12 23.48 14.64 0.62 2.09 4.16E-10 9.46E-06 0.232 0.768 43.49 19 114 1.000 1.000 38.12 23.48 14.64 0.62 2.09 4.16E-10 9.46E-06 0.232 0.768 43.49 20 120 1.000 38.12 23.48 14.64 0.62 2.09 4.16E-10 9.46E-06 0.232 0.768 43.49 21 126 1.000 38.12 23.48 14.64 0.62 2.09 4.16E-10 9.46E-06 0.232 0.768 43.50 22 132 23.48 14.64 0.62 2.09 4.16E-10 9.47E-06 0.232 0.768 43.50 23 138 1.000 38.12 1.000 38.12 23.48 14.64 0.62 2.09 4.16E-10 9.47E-06 0.232 0.768 43.50 24 144 38.13 23.48 14.64 0.62 2.09 4.16E-10 9.47E-06 0.232 0.768 43.50 25 150 1.000 ANO-1 Cylind Flaw NP.xls 37 Head FCG

A FRAMATOME ANP 32-5021539-01 9.0

SUMMARY

OF RESULTS The flaw evaluation results for 25 years of fatigue crack growth are as follows.

9.1 Propagation of a Continuous External Circumferential Flaw Along Path 1 a) Fatigue crack growth analysis:

Initial flaw size, a,= [ ] in.

Final flaw size, af< [ ]in.

Stress intensity factor at final flaw size, Ki (aef) < 0 ksi`/in Fracture toughness Kia = 200 ksiqin Fracture toughness margin, K,3 / Ki > 41i0 b) Limit load analysis:

Limit load, Po = [ ] lbs Applied loads: normal/upset, ] lbs emergency/faulted, lbs Limit load margins: normal/upset, Po / P= 9.44 > 3.0 emergency/faulted, Po / P= 6.67 > 1.5 9.2 Fatigue Crack Growth of a Semi-Circular External Axial Flaw Along Path 1 Initial flaw size, a,= [ ]in.

Radial Growth Final flaw size, af < [ ] in.

Stress intensity factor at final flaw size, K,(aef) = 28.8 ksi`1in Fracture toughness K1. = 200 ksi4in Fracture toughness margin, K13 / K, = 6.94 > 410 Axial Growth af < ]in.

Final flaw size, [

Stress intensity factor at final flaw size, Ki (aef) = 33.3 ksi4in Fracture toughness Kia = 200 ksi'lin Fracture toughness margin, Kia / K, = 6.01 > 10 9.3 Fatigue Crack Growth of a Continuous Cylindrical Flaw Along Path 2 Initial flaw size, a,= in.

Final flaw size, af<[ ]in.

Stress intensity factor at final flaw size, K,(aer) = 43.7 ksi`in Fracture toughness KI, = 200 ksi'lin Fracture toughness margin, Kia / KI = 4.58 > 410 38

A 32-5021539-01 FRAMATOME ANP

10.0 CONCLUSION

The results of the analysis demonstrate that the [ ] inch weld anomaly is acceptable for a 25 year design life of the CRDM ID temper bead weld repair. Significant fracture toughness margins have been demonstrated for both of the flaw propagation paths considered in the analysis. The minimum fracture toughness margins for flaw propagation Paths 1 and 2 have been shown to be 6.01 and 4.58, respectively, as compared to the required margins of 10 for normallupset conditions and 412 for emergency/faulted conditions per Section XI, IWB-3612 (Reference 3). Fatigue crack growth is minimal. The maximum final flaw size is less than 0.101 inch (considering both flaw propagation paths). A limriit load analysis was also performed considering the ductile Alloy 600 and Alloy [ ] materials along flaw propagation Path 1. The analysis showed limit load margins of 9.44 for normal/upset conditions and 6.67 for emergency/faulted conditions, as compared to the required margins of 3.0 and 1.5, respectively, per Section Xl, IWB-3642 (Reference 3).

39

A 32-5021539-01 FRAMATOME ANP

11.0 REFERENCES

1. Framatome ANP Drawing 02-5021508E-2, "CRDM Nozzle ID Temper Bead Weld Repair, ANO-1 ."

2- FramatomeANP Document 51-5021517-01, "ANO-1 CRDM Nozzle ID Ambient Temperature Temper Bead Weld Repair Requirements," November 2002.

3. ASME Boiler and Pressure Vessel Code, Section Xl, Rules for Inservice Inspection of Nuclear Power Plant Components, 1992 Edition with No Addenda.

4. BAW-10046A, 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.

5. Framatome ANP Document 38-1290261-00, ANO Design Input Record No. 2001-0230 020-02, dated 111612002 forwarding Dominion Engineering Stress Calculations for IDTB Weld Repair.

6. Framatome ANP Document 32-5012424-10, "CRDM Temperbead Bore Weld Analysis,"

November 2002.

7. ASME Section 11,Part C, "Specification for Welding Rods, Electrodes, and Filler Metals,"

1999 Addenda.

8. Framatome ANP Document 38-1288355-00, "Flaw Acceptance Criteria."

9. ASME Boiler and Pressure Vessel Code,Section III, Rules for Construction of Nuclear Power Plant Components, Division 1 - Appendices, 1989 Edition with No Addenda.

10. I.S. Raju and J.C. Newman Jr., "Stress Intensity Factors for Internal and External Surface Cracks in Cylindrical Vessels," Transactions of the ASME, Journal of Pressure Vessel Technology, pp. 293-298, Vol. 104, November 1982.

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

12. W.J. Mills, "Fracture Toughness of Two Ni-Fe-Cr Alloys," Hanford Engineering Development Laboratory Document HEDL-SA-3309, April 1985.

13. C.B. Buchafet and W.H. Bamford, "Stress Intensity Factor Solutions for Continuous Surface Flaws in Reactor Pressure Vessels," Mechanics of Crack Growth, ASTM STP 590, American Society for Testing and Materials, 1976, pp. 385-402.

14. EPRI Topical Report, EPRI NP-1931, "An Engineering Approach for Elastic-Plastic Fracture Analysis," Research Project 1237-1, prepared by V. Kumar et al of General Electric Company, July 1981..

40

A 32-5021539-01 FRAMATOME ANP

15. General Electric Report, SRD-82-048, "Estimation Technique for the Prediction of Elastic Plastic Fracture of Structural Components of Nuclear Systems," by V. Kumar et al, Contract RP1237-1, Combined Fifth and Sixth Semi-Annual Report, March 1982.

16. ASME Boiler and Pressure Vessel Code,Section XI, Rules for Inservice Inspection of Nuclear Power Plant Components, Appendix A, Analysis of Flaws, 1995 Edition with No Addenda (used for updated stress intensity factor solution).

17. Framatome ANP Document 32-5012403-00, "Oc-3 CRDM Nozzle Circumferential Flaw Evaluations," April 2001.

18. EPRI Topical Report, EPRI NP-3607, "Advances in Elastic-Plastic Fracture Analysis,"

Research Project 1237-1, prepared by V. Kumar et al of General Electric Company, August 1984.

19. Framatome ANP Document 18-1173987-03, "RCS Functional Specification for ANO-1."

20. Letter from Entergy for Reactor Trip Transients, November 8, 2002 (attached as Appendix B).

41

A 32-5021539-01 FRAMATOME ANP Appendix A Certification Document ANO-1 CRDM NOZZLE IDTB WELD ANOMALY FLAW EVALUATIONS SHEET 1 OF 1 I certify that the flaw evaluations contained in this calculation package are evidence that the CRDM nozzle repair, as defined by Framatome ANP drawing 02-5021508E-2, meets the fracture mechanics requirements identified in Framatome ANP Document 51-5021517-01 and the ASME Boiler and Pressure Vessel Code,Section XI, 1992 Edition with no addenda, for the design transients provided in Framatome ANP Document 18-1173987-03.

Exception:

The rod withdrawal transient, identified as Transient No. 11 in the functional specification for the reactor coolant system at ANO-1 (Framatome ANP Document 18-1173987-03), is analyzed as an emergency condition transient since ANO-1 has not experienced any rod withdrawal transient transients during its plant life and the likelihood of even a single future occurrence is remote. It is even more unlikely that a CRDM nozzle woi~ld experience 25 cycles of this transient, which is a requirement for classification as an emergency event.

Attested to this date: November 8, 2002 By: WIA Douglas E. Killian Framatome ANP, Inc.

Nuclear Engineering Business Unit Lynchburg, Virginia License No. 15308 Virginia Board for Architects, Professional Engineers, Land Surveyors, Certified Interior Designers, and Landscape Architects 42

A FRAMATOME ANP 32-5021539-01 Appendix B Letter from Entergy for Reactor Trip Transients 43

"3-Z-SozIS3 -o ARKANSAS NUCLEAR ONE GE3NERATION SUPPORT BUILDING 1448 S.R. 333 RUSSELLVILLE, AR 72801 i ~FAX COVERLETTER DAT, . /-*,-.o. .

CO.,ANY NAME CITY, STATE FAX NUMBER VERIFY NUIMBE* OF PAGES INCLUDrNG COVER COMPAINYNAMEN .- i Iz '.

DEP'ARTMENT FAX NUMBlER (501) 858.4955 VERIFY (501) 858-4302 TO BE COMPLETED BY SENDER AFTER TRANSMISSION IS COMPLETE, PLEASE RETURN BY MAIL___ DISCARD CALL FOR PICK-UP PHONE NUMBER

SAM NOV-8-;2-002 FRI 02:10 PH 110 S FAX NO. 501 858 4955 P. 02 S'3- 60oZ ~

Steve:

Based on a review performed by our System Engineering, as documented In our ER.ANO-2001 0230-024, ANO-1 has not had a "high pressure. rod withdrawal accident" as described In transient No. 11 of RCS Functional Specification 18-1173987-03. See below for details.

From: MEANS, BRACY E Sonrt: Friday, November 08, 2002 11:20 AM To: TO, RAYMOND M; BAUMAN, DAVID N; CHISUM, MICHAEL R; DAIBER, BRYAN J; CHADBOURN, HARMON C: GRAY, BRIAN C; LEWIS, RAYMOND S

Subject:

Reactor Trip Transients Raymond, The RCS Functional Specification Transient 11 (Rod Withdrawal Accident (upset condition) defines this condition as follows:

"The rod withdrawal accident occurs when one group of control rods is accidentally withdrawn from the core at the maximum rate when the reactor Is operating at a low power. This transient is assumed to occur when the reactor is operating at 15% power. This causes a rapid rise in power and a reactor trip."

In an effort to facilitate transient cycle counting, FTI performed a review of the functional specification transients and consolidated the above transient in what is currently tracked as Transicnt #19. This FTI report has been filed as Engineering Report 95R-1015-01 and Is FTI document number 51-1235146-01. In the FTI report, ANO historical transients though the year 1995 were reviewed and consolidated to reflect the current transient cycle recording document reflected in Procedure 1010.010. The transient cycle log procedure 1010.010 currently records reactor trip transients for the following conditions:

"* Transient 17 - Reactor Trips with loss of RC flow. There have been 10 recorded transients in this category.

"* Transient 18 - Reactor Trips with Post Overcooling. There have been 0 transients recorded in this category. The 511811996 trip did not meet the overcooling requirements to be recorded Inthis category. It was captured in transient 19.

"* Transient 19 - All Other Reactor Trips. There have been 96 recorded transients in this category through December 31,2001. There were two RPS high pressure trips recorded.

These high pressure trips did not cause a Pressurizer Code Safety valve relief with set point of 2500 psi.

The control room logs were reviewed in Autolog as well as PCRS with the plant computer data to assess related trips that may have met the criteria for Functional Specification Transient #11.

Based on this level of review, It can be stated that ANO-1 has not experienced any of the Functional Specification transient #11 events since commercial operation. Please be advised that the transients associated with plant performance for the 2002 operating year have not been finalized. However, it Is certain that none of the Functional Specification Transient #11 conditions have been experienced for ANO-I for this year. Please advise if there are any clarifications required for the data provided herewith or Ifany additional Input or research Is desired.

Bracy

/27 Encit egy ANO-Design Engineering Structural Group

CNRO-2002-00058, Proposed Alternative to ASME Examination Requirements for Repairs Performed on Reactor Vessel Head Penetrations

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