ATTACHMENT 4

PNP 2018-054

Framatome Document No. 51-9292503,

Palisades CRDM & ICI Nozzle IDTB Repair - Life Assessment Summary

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Document No.: 51 - 9292503 - 000

Palisades CEOM Nozzle IOTB Repair - Life Assessment Summary

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Palisades CEOM Nozzle 10TB Repair - Life Assessment Summary

Safety Related? IZlYES DNO

Does this document establish design or technical requirements? D YES

Does this document contain assumptions requiring verification? 0 YES

IXjNO

IZINO

------bD'{lg~esli-'tHllYiIl!·s'_'dwgcumem..contain Customer Required...Eom!at .... ?-I::;D~YE ....... ,J------it:~'IV--------------------+

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Approved or Comments

Ryan Hosler Supervisory Engineer Materials & Fracture Mechanics Stephen Fyfitch . Technical Consultant Materials & Fracture Mechanics Tim Wiger Advisory Engineer Component Analysis Tomas Straka Advisory Engineer Component Analysis Brian Haibach RTM Structural Analysis & MODS

LP

LR

p

Note: PILP designates Preparer (P), Lead Preparer (LP) M designates Mentor (M) RlLR designates Reviewer (R), Lead Reviewer (LR)

All except Appendix A

All except Appendix A

u·~Lt .. j.~ Appendix A

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Document No.: 51-9292503-000

Palisades CEDM Nozzle IDTB Repair - Life Assessment Summary

Record of Revision Revision Pages/Sections/

No. Paragraphs Changed Brief Description / Change Authorization

00 All Original release (November 2018). The proprietary version of this document is 51-5047343-007.

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Palisades CEDM Nozzle IDTB Repair - Life Assessment Summary

Table of Contents Page

SIGNATURE BLOCK .. ..... ... ...... ... .... .......... ..... .. ..... ...... ... ........ .. ...... .. .... .. .. ... .... .. ... ...... ..... ... ...... ... .. ..... .... .. 2

RECORD OF REVISION ....... ... ..... ..... ....... ..................................... .... ........ .. .. ... ... .. ...... ....... .. .... ......... .. .... 3

LIST OF FIGURES ................. .. ........... .... ......................................................... ........................... .. ........... 5

1.0 INTRODUCTION ... ... .............. ... ..... ... .. .. ... ... ..... ....... ......... .. ..... ...... ....... .. ... .. .. .. ..... ... .... ... ... ... .... ..... 6

2.0 ASSUMPTIONS ..... .................................... .. ............................. ........... ..... .... ... ...... ... .. ... .. .. ........ ... 8 2.1 Assumptions Requiring Verification .. ..................... .. ........................... .. ....................... ... ... 8 2.2 Justified Assumptions .. .... ........... ......... .. ..... .. .......... ...... .............. .. .... ....... ........... ..... .. .. .. .... 8

3.0 RESULTS ..... ..... .... .... ... .. .... ..... .. ....... ..... .. ...... .. .. ..... ... ....... .. .. .... .. ... ... .... .. ........ ........ .. ..... .... ..... ... .... 9

4.0 CONCLUSION .. .. .... .. .. .. .... .... .. .. .... ...... ... .... ..... ..... .. ... ........ ... .... ....... .. .. ................ .. ...................... 11

5.0 REFERENCES ...... ... ..... ...... ...................................... ................. ............ ..... ....... .. ..... .. ........... .. .. . 11

APPENDIX A : ROLL EXPANSION STRESSES .. .. ......... .. ..... .... ..... ...... ........ .. .. .... .. ... .. ... .. ..... ........... 13

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Palisades CEDM Nozzle lOTS Repair - Life Assessment Summary

List of Figures Page

Figure 1-1: lOTS CEDM Nozzle Repair Configuration [2] ... ... .. .. ..... .. .. ... .... ... ..... .. .. .. ... ...... ....... ... ... ...... .... 7

Figure A-1: 10 Surface Stress (psi) vs. Distance (inches) Above the Roll Transition .... ... .. .................. . 15

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Palisades CEDM Nozzle lOTS Repair - Life Assessment Summary

1.0 INTRODUCTION Alloy 600 control rod drive mechanism/control element drive mechanism (CRDM / CEDM) nozzles and thermocouple (TC) nozzles at domestic pressurized water reactors (PWRs) have leaked via cracking attributed to primary water stress corrosion cracking (PWSCC) [1]. As a result of inspection findings in the Palisades Unit 1 reactor vessel (R V) head penetrations, Framatome has prepared a repair configuration for CEDM nozzles. The inner diameter temper bead (IOTB) CEDM nozzle repair method shown in Figure 1-1 will be used [2]. The proposed repair configuration will leave portions of the low alloy steel inside the R V head penetration exposed to the primary reactor coolant.

The low alloy steel will be subject to general corrosion during operating and shutdown conditions. A materials evaluation has been performed to determine the maximum corrosion rate of the exposed RV head low alloy steel and to evaluate any other potential corrosion concerns involving the IOTB weld repair [3]. Detailed stress and fatigue analyses were performed to establish the minimum life expectancy [4]. The proposed repair involves leaving a portion of the original nozzle in place. A roll expansion is used to hold the nozzle in place during the welding process. A life assessment was also performed to evaluate the PWSCC susceptibility of the remaining Alloy 600 CEDM nozzle portion affected by the IOTB weld repair, which was performed in Section 3.0.

In addition, two fracture mechanics flaw evaluations were performed to evaluate the life expectancy of the repair with assumed flaw sizes and locations. The first analysis considered sub-critical growth of presumed pre-existing PWSCC cracks in the original Alloy 182 J-groove weld [5]. The second analysis evaluated a postulated weld anomaly in the CEDM nozzle temper bead weld. The postulated anomaly was assumed to be a O.l-inch semi-circular flaw that is 360 degrees around the circumference at the "triple point" location where there is a confluence of three different materials, the Alloy 600 CEDM nozzle, the Alloy 52/52M temper bead weld, and the low alloy steel head [6] .

The purpose of this life assessment summary document is to provide a summary of the calculations and evaluations performed to establish the life expectancy of the IOTB weld repair on Alloy 600 CEDM nozzles for Palisades.

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Palisades CEDM Nozzle IDTB Repair - Life Assessment Summary

Figure 1-1: 10TB CEOM Nozzle Repair Configuration [2]

Remaining Alloy 600 Nozzle

Alloy 52/52M IDTBWeld

Location A Exposed Low Alloy Steel

Counterbore

Replacement Lower Nozzle 1----- Alloy 690

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Palisades CEDM Nozzle lOTS Repair - Life Assessment Summary

2.0 ASSUMPTIONS

2.1 Assumptions Requiring Verification

There are no assumptions requiring verification.

2.2 Justified Assumptions

1. [ ] This assumption is justified based on the following:

a. [ ] factors that have historically caused PWSCC (elevated tensile stress (i)

and/or off-chemistry conditions (ii)) [ ] Furthermore, if shallow flaws were present, they would be observed (iii) and/or removed (iv).

i. For RVCH penetration nozzles, the primary cause of elevated tensile stress is residual stresses from the original l-groove welding process, which only affects the base metal directly adjacent to the weld (i.e., the heat-affected zone). [

] ii. Off-chemistry conditions have caused PWSCC flaws in CRDM nozzles in relatively low

stress locations at one international site where resin ingress caused prolonged periods of high sulfate levels [7]. In response, the Nuclear Regulatory Commission (NRC) requested in Generic Letter (GL) 97-01 that all u.s. PWRs report whether any resin intrusions exceeded the EPRI PWR water chemistry guidelines [8]. Palisades responded to GL 97-01 and the NRC was satisfied with the response [9]. Furthermore, since Palisades follows the EPRI PWR Primary Water Chemistry Guidelines (currently at Revision 7) [10], in the unlikely event of significant resin intrusion, it would be identified and addressed prior to becoming a concern.

iii. [ ] As part of the repair process, the whole length of the inner diameter of the remaining Alloy 600 nozzle affected by the repair is required to pass dye-penetrant testing (PT) after machining [2]. The lack of PT indications is commonly accepted as indicative of a flaw-free surface. While PT has missed reasonably deep cracks in CRDM nozzle welds, it is acknowledged that a higher quality inspection surface may help improve the reliability of the PT [11]. [

] iv. [

] After performing the IOTB weld, the inner diameter of the Alloy 600 nozzle will be bored, [

]

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3.0 RESULTS The materials evaluation addressed the potential corrosion concerns associated with the weld repairs planned for the CEDM nozzles. The PWSCC and general corrosion properties of Alloy 690 and its weld metals were addressed. It was concluded that Alloy 690 and its weld metals are the best commercially available materials for this application [3].

The potential corrosion concerns of the RV closure head low alloy steel include general, galvanic, crevice, stress corrosion cracking (SCC), and hydrogen embrittlement. Galvanic corrosion, crevice corrosion, SCC, and hydrogen embrittlement of the RV head low alloy steel are not significant concerns based on previous operational experience with low alloy steel exposed to primary coolant. The general corrosion rate for the R V head low alloy steel, under the anticipated exposure conditions, is [ ] This corrosion rate is based on an 18-month operating cycle followed by a 2-month refueling cycle [3]. Because the corrosion rate is [

] Detailed stress and fatigue analyses of the IOTB CEDM nozzle weld repair were performed. The analyses demonstrated that the IOTB CEDM weld repair design meets the stress and fatigue requirements set by American Society of Mechanical Engineers Boiler and Pressure Vessel (ASME B&PV) Code, Section III, 1989 Edition without Addendum. The conservative fatigue analyses conclude that the fatigue usage factor for 27 calendar years of operation is [ ] for the CEDM weld repair [4].

The life expectancy of the rotary peened IOTB CEDM weld repair was evaluated with respect to the PWSCC concerns of the remaining Alloy 600 CEDM nozzle portion affected by the IOTB weld repair. [

] Rotary peening is a captive shot technology that was originally developed by 3M® Corporation for small and/or hard-to-reach surfaces. This process creates compressive residual stresses on the surface much like traditional shot peening, except that the rotary peening process is more controlled and thus results in a more consistent and uniform compressive stress layer. Surface remediation using rotary peening should inhibit PWSCC initiation based on the principle that PWSCC requires a sustained tensile stress and the peening process will create a uniform compressive stress layer.

The Electric Power Research Institute (EPRI) Materials Reliability Program (MRP) conducted a test program of various Alloy 600 R VCH penetration PWSCC mitigation techniques, which included flapper wheel and shot peening [12]. The flapper wheel technique discussed in Reference [12] is a grinding process and is not similar to rotary peening. Rotary peening is a captive shot technology, which is expected to perform very similarly to shot peening regarding inhibiting PWSCC initiation. The conclusions of the EPRI report indicate that shot peening was among the best techniques tested with no PWSCC observed in sample regions applicable to remediation performed at an operating plant. The sample region that was considered not applicable was at the cut edge ofthe sample where the compressive stress field was compromised.

[

] As discussed above, the compressive stress imparted by the rotary peening process removes the stress component ofPWSCC susceptibility. Rotary peening will be performed on all portions of the Alloy 600 remaining nozzle 10

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affected by the repair that are expected to have tensile stresses greater than 20 ksi prior to mitigation, which includes the IOTB weld HAZ and the roll transition region [2]. Note that the distance above the roll transition that requires rotary peening is addressed in Appendix A. Given the compressive stresses achieved by rotary peening, PWSCC initiation is not expected during the 27 year fatigue life of the repair, which is discussed below.

The remaining l-groove weld in the CEDM nozzles, after the IOTB repair, was analyzed by postulating a radial crack in the Alloy 182 l-groove weld and butter and evaluating fatigue crack growth into the low alloy steel head [5]. Since a potential flaw in the l-groove weld cannot be sized by currently available NDE techniques, it was assumed that the "as-left" condition of the remaining l-groove weld includes degraded or cracked weld material extending through the entire l-groove weld and Alloy 182 butter material. It was further postulated that a small fatigue initiated flaw forms in the low alloy steel head and combines with the primary water stress corrosion crack in the weld to form a large radial corner flaw that propagates into the head by fatigue crack growth under cyclic loading conditions. A flaw evaluation of any partial penetration nozzle l-groove weld is inherently difficult due to the presence of large residual stresses created during the welding process. An analysis of the Palisades head is particularly challenging because of a high reference nil ductility transition temperature (RT NDT) [ ] for the low alloy steel base metal.

A fatigue crack growth analysis was performed to determine final flaw sizes on the uphill and downhill sides of the l-groove weld after 27 calendar years of operation. Stress intensity factors were first determined using a three-dimensional finite element analysis for cracks extending to the butter/head interface and applying both residual and operating stresses for each of eight analyzed transients. For each increment of crack growth, stress intensity factors were increased by the [ ] This is a conservative approximation since both the residual stresses and the thermal gradient stresses decrease in the direction of crack propagation. Flaw growth into the head was calculated to be 0.610-inch on the uphill side and 0.324-inch on the downhill side [5].

A combination of linear elastic fracture mechanics (LEFM) and elastic-plastic fracture mechanics (EPFM) was utilized to evaluate the final uphill and downhill flaw sizes after 27 calendar years of crack growth. At operating temperatures when EPFM is the appropriate analysis method, a l-integralltearing modulus (l-T) diagram was used to evaluate flaw stability with safety factors of 3 on primary (pressure) stresses and 1.5 secondary (residual plus thermal) stresses. The crack driving force was also checked against the l-integral resistance (J-R) curve at a crack extension of O.I-inch using safety factors of 1.5 and 1. 0 for primary and secondary stresses, respectively. Near room temperature, when the material is less ductile and LEFM is the more appropriate analysis method, stress intensity factors were compared to the crack initiation fracture toughness (K1c) using a safety factor of.y2 [5].

The highest crack tip stress intensity factors occur during [ ] when the pressure is [ ] and the temperature [ ] At these conditions, the applied l-integral at the uphill crack front is 1.470 kips/in, with safety factors on on pressure stresses and 1.5 on residual and thermal stresses, which is less than the l-integral for the material, 3.259 kips/in, at the point of instability. Flaw stability is also demonstrated by an applied tearing modulus of9.323 kips/in, which is well below a tearing modulus of62.04 kips/in for the material. For a flaw on the downhill side, the applied l-integral is 2.414 kips/in, compared to a l-integral at the point of instability of 3.270 kips/in. The applied tearing modulus is 15.18 kips/in and the corresponding tearing modulus for the material is 31.28 kips/in, again demonstrating flaw stability. As a final check on the EPFM analysis, the applied l-integrals for safety factors of 1.5 on pressure and 1.0 on residual and thermal loads are compared to the l-integral for the material at a crack extension of O.I-inch. It was determined that the applied l-integrals of 0.361 kips/in on the uphill side and 0.597 kips/in on the downhill side are both less than the required value of 1.711 kips/in for the material [5].

[ ] where LEFM is the appropriate method for flaw evaluation, it is widely recognized that l-groove weld residual stresses make it improbable that the ASME B&PV Code required fracture toughness margin can be satisfied, as was the case for the Palisades CEDM l-groove weld

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flaws. Although these residual stresses would tend to be relieved as a crack propagated farther into the head, additional analysis was performed for larger flaw sizes to demonstrate that the required fracture toughness margin could be met while still considering residual stress. New crack sizes were determined by reviewing stress contour plots and selecting crack extensions that would locate the uphill and downhill crack fronts in regions of compressive residual stress. On the uphill side, the crack was extended [ ] beyond the butter. A larger crack extension of [ ] was required on the downhill side to extend the crack into the compressive residual stress field. [

] The K1c fracture toughness at this temperature is [ ] At these larger crack sizes, the applied stress intensity factors were calculated to [

] on the uphill side and [ ] on the downhill side, both of which satisfy a K1c /.../2 acceptance criterion of [ ] In summary, a combination ofLEFM and EPFM was used to show that postulated flaws in the CEDM J-groove weld and butter are acceptable for 27 calendar years of operation [5].

The results of the triple point flaw analyses demonstrate that a O.IOO-inch weld anomaly is acceptable for 27 calendar years of operation following the CEDM nozzle ID temper bead weld repair, considering the transient frequencies of the applicable transients. Significant design margins have been demonstrated for all flaw propagation paths considered in the analysis. Flaw acceptance is based on the 2007 Edition through 2008 Addenda ASME B&PV Code Section XI criteria for applied stress intensity factor (lWB-3612) and limit load (lWB-3644). Fatigue crack growth is minimal along each flaw propagation path with the maximum final flaw size being only [ ] for the CEDM nozzle repair. The minimum fracture toughness margin is 3.58 for the CEDM nozzle, compared to the required margin of.../1O per IWB-3612. The margin on limit load is 7.96 for a CEDM nozzle, compared to the required margin of2.7 per Section XI, IWB-3644 [6].

4.0 CONCLUSION Based on the analyses and evaluations summarized above, the minimum life expectancy for the rotary peened repair is conservatively estimated at 27 calendar years for a CEDM nozzle.

5.0 REFERENCES

1. "PWR Materials Reliability Program Interim Alloy 600 Safety Assessments for US PWR Plants (MRP-44), Part 2 Reactor Vessel Top Head Penetrations," TP-IOOI491, Part 2, May 2001, Electric Power Research Institute, Palo Alto, CA.

2. [ ]

3. [ ]

4. [ ]

5. [ ]

6. [ ]

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7. NRC Information Notice 96-11, "Ingress of Demineralizer Resins Increases Potential for Stress Corrosion Cracking of Control Rod Drive Mechanism Penetrations," February 14, 1996.

8. NRC Generic Letter 97-01, "Degradation of Control Rod Drive Mechanism Nozzle and Other Vessel Closure Head Penetrations," April 1, 1997.

9. Letter from Robert G. Schaaf, NRC, to Mr. Nathan L. Haskell, Palisades Plant, "Palisades Plant-Closeout of Generic Letter 97-01 'Degradation ofCRDM/CEDM Nozzle and Other Vessel Closure Head Penetrations' TAC No. M98582", November 29, 1999, NRC Accession Number ML993350545.

10. "Pressurized Water Reactor Primary Water Chemistry Guidelines: Volume 1, Revision 7" EPRl, Palo Alto, CA: 2014. 3002000505.

11. NRC Report NUREG/CR-6996 (PNNL-18372), "Nondestructive and Destructive Examination Studies on Removed-from-Service Control Rod Drive Mechanism Penetrations," July 2009, NRC Accession Number ML092170313 and ML092170314.

12. Materials Reliability Program: An Assessment of the Control Rod Drive Mechanism (CRDM) Alloy 600 Reactor Vessel Head Penetration PWSCC Remedial Techniques (MRP-61), EPRI, Palo Alto, CA: 2003. 1008901.

13. [ ]

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APPENDIX A: ROLL EXPANSION STRESSES

A.1 Background and Purpose:

Appendix B of Reference 1 reports residual plus operating stresses on the wetted surface of a CRDM nozzle as a function of the distance from the top of the effective length of the roll transition. While these results are based on a nozzle with a somewhat different geometry, they can be used to provide a conservative estimate of what the stresses in a Palisades nozzle would be if the same methodology were used to calculate the stresses at Palisades.

The nozzles are similar in the sense that:

1) Both nozzles are [ ] and roll expanded into a low alloy steel reactor vessel head. The low alloy steel head is not expected to plastically deform.

2) Both nozzles have nearly the same inside diameter [ ] 3) A similar level of roll expansion (i.e., permanent set) value is being targeted.

[ ]

A.2 Purpose:

The reason for calculating these stresses is that it is desired to ensure that any wetted surface that sees residual plus operating stresses at 100% steady state operation greater than 20 ksi are subjected to surface stress mitigation, which creates a compressive layer on the wetted surface of the component, which tends to reduce or eliminate the material's susceptibility ofPWSCC.

The existing analysis considers different cases regarding the relative location of the roll expansion with the top surface of the RVCH, [

] A.3 Assumptions:

There are no significant assumptions being made in this evaluation.

A.4 Discussion on Roll Expansion:

The roll expansion will be executed with the same tooling and same torque as that used for the nozzle in the existing analysis. Per References 2 and 3, both expansions [

] Since it is the deformation resulting in comparable strain in the rolled region that leads to the tensile stress above the rolled region, it should result in comparable tensile stresses. Therefore, it is reasonable to use the results of the existing calculation.

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A.5 Conversion:

To convert the results of the existing analysis into conservative results for Palisades, [

]

Based on this, for a given level of permanent set, the reduction in stress level at [ ] above the roll expansion zone in the existing analysis will be achieved in only [ ] for the Palisades design.

Even if we conservatively ignore the fact that the residual stresses will drop off faster for the Palisades geometry, it is shown below that the residual plus operating stress at Palisades will fall below 20 ksi at approximately [ ] above the roll transition zone. Figure A-I simply takes the existing component stresses and adds [ ] A.6 Conclusions: Based on the existing analysis performed for a similar CRDM geometry, and a conservative conversion of those stresses to the Palisades configuration, it has been shown that stresses attenuate to less than 20 ksi at [ ] inches above the effective length of the roll transition zone. Therefore, as long as the surface stress mitigating process extends above that location, there should not be any wetted surface stresses remaining above 20 ksi at 100% steady state operating conditions.

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Figure A-1: 10 Surface Stress (psi) vs. Distance (inches) Above the Roll Transition

A.7 References:

1. [ ]

2. [ ]

3. [ ]

4. [ ]

5. [ ]

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