TMI-11-062, Response to Request for Additional Information - Submittal of Relief Request RR-10-02 Concerning the Weld Overlay of the Pressurizer Spray Nozzle to Safe-End and Safe-End to Elbow Dissimilar Metal Welds

Response to Request for Additional Information - Submittal of Relief Request RR-10-02 Concerning the Weld Overlay of the Pressurizer Spray Nozzle to Safe-End and Safe-End to Elbow Dissimilar Metal Welds

ML110980257
Person / Time
Site:	Three Mile Island
Issue date:	04/06/2011
From:	David Helker Exelon Generation Co, Exelon Nuclear
To:	Document Control Desk, Office of Nuclear Reactor Regulation
Shared Package
ML110980255	List: TMI-11-062, Response to Request for Additional Information - Submittal of Relief Request RR-10-02 Concerning the Weld Overlay of the Pressurizer Spray Nozzle to Safe-End and Safe-End to Elbow Dissimilar Metal Welds
References
TAC ME4795, TMI-11-062 1000320.310, 1000320.314, 1000320.315, 1000320.316
Download: ML110980257 (152)

Text

Exelon Nuclear wwwexeloncorp,corn Exel n Exelon Way Nuclear Kennett PA 19348 PROPRIETARY INFORMATION - WITHHOLD UNDER 10 CFR 2.390 10 CFR 50.55a TMI-11-062 April 6, 2011 U.S. Nuclear Regulatory Commission Attn: Document Control Desk Washington, DC 20555-0001 Three Mile Island Nuclear Station, Unit 1 Renewed Facility Operating License No. DPR-50 NRC Docket No. 50-289

Subject:

Response to Request for Additional Information - Submittal of Relief Request RR-10-02 Concerning the Weld Overlay of the Pressurizer Spray Nozzle to Safe-End and Safe-End to Elbow Dissimilar Metal Welds

References:

1) Letter from P. B. Cowan (Exelon Generation Company, LLC) to U.S. Nuclear Regulatory Commission, "Submittal of Relief Request RR-10-02 Concerning the Weld Overlay of the Pressurizer Spray Nozzle to Safe-End and Safe-End to Elbow Dissimilar Metal Welds," dated September 30,2010

2) Letter from P. Bamford (U.S. Nuclear Regulatory Commission) to M. J.

Pacilio, "Three Mile Island Nuclear Station, Unit 1 - Request for Additional Information Regarding Relief Request RR-10-02, Weld Overlay of the Pressurizer Spray Nozzle to Safe-End and Safe-End to Elbow Dissimilar Metal Welds (TAC NO. ME4795)," dated February 28,2011

3) Letter from D. P. Helker (Exelon Generation Company, LLC) to U.S. Nuclear Regulatory Commission, "Response to Request for Additional Information -

Submittal of Relief Request RR-10-02 Concerning the Weld Overlay of the Pressurizer Spray Nozzle to Safe-End and Safe-End to Elbow Dissimilar Metal Welds," dated March 9, 2011 In the Reference 1 letter, Exelon Generation Company, LLC (Exelon) requested relief to perform a weld overlay of pressurizer spray nozzle to safe-end and safe-end to elbow dissimilar metal welds at Three Mile Island Nuclear Station (TMI), Unit 1. In the Reference 2 letter, the U.S.

Nuclear Regulatory Commission requested additional information. Reference 3 contained our response to questions 2, 3 and 4. As discussed in Reference 2, the response to question 1 would be provided by April 28, 2011.

Attachment 4 transmitted herewith contains Proprietary Information.

When separated from attachments, this document is decontrolled.

Response to Request for Additional Information Relief Request RR~ 10~02 Concerning the Weld Overlay of Dissimilar Metal Welds April 6, 2011 Page 2 is our response to question 1. Attachments 2 and 3 contain copies of Calculation No. 1000320.315, Revision 0 and Calculation No. 1000320.316, Revision 0, respectively. contains information proprietary to AREVA NP Inc. (AREVA) and Structural Integrity Associates (SI), Inc. AREVA and SI request that Calculation No. 1000320.310, Revision 0 and Calculation No. 1000320.314, Revision 0 be withheld from public disclosure in accordance with 10 CFR 2.390(b)(4). Attachment 5 contains non~proprietary versions of these two calculations (I.e., Calculations 1000320.310, Revision 0 and Calculation 1000320.314, Revision 0). Affidavits supporting AREVA and SI's request are contained in Attachment 6.

There are no regulatory commitments contained in this submittal.

If you have any questions concerning this letter, please contact Tom Loomis at (610) 765~5510.

Respectfully, David P. Helker Manager ~ Licensing & Regulatory Affairs Exelon Generation Company, LLC Attachments: 1) Response to Request for Additional Information ~ Submittal of Relief Request RR~10~02 Concerning the Weld Overlay of the Pressurizer Spray Nozzle to Safe~End and Safe~End to Elbow Dissimilar Metal Welds

2) Calculation No.1 000320.315

3) Calculation No. 1000320.316

4) Proprietary Version ~ Calculation No. 1000320.310 and Calculation No.

1000320.314

5) Non~Proprietary Version ~ Calculation No.1 000320.31 0 and Calculation No.

1000320.314

6) Affidavits cc: Regional Administrator, Region I, USNRC USNRC Senior Resident Inspector, TMI USNRC Project Manager, [TMI) USNRC

ATTACHMENT 1 Response to Request for Additional Information - Submittal of Relief Request RR-10-02 Concerning the Weld Overlay of the Pressurizer Spray Nozzle to Safe-End and Safe-End to Elbow Dissimilar Metal Welds

Attachment 1 Response to Request for Additional Information Relief Request RR-10-02 Concerning the Weld Overlay of Dissimilar Metal Welds Page 1 of 1 Question:

1. Section 5 of relief request (RR)-1 0-02 (page 4, third paragraph) states that the overlay design is currently in development. Please submit the weld overlay design information, including analyses, to demonstrate that the weld overlay design will mitigate the potential for primary water stress-corrosion cracking in the Alloy 82/182 dissimilar metal welds.

Response

The following calculations are attached:

1. Calculation No. 1000320.315, "ASME Code,Section III Qualification of Pressurizer Spray Nozzle with Weld Overlay Repair," Revision 0 (Attachment 2)

2. Calculation NO.1 000320.316, "Crack Growth Evaluation of Pressurizer Spray Nozzle with Weld Overlay," Revision 0 (Attachment 3)

3. Calculation No. 1000320.310, "Pressurizer Spray Nozzle Weld Overlay Sizing Calculation,"

Revision 0 (Attachment 4 - Proprietary Version, Attachment 5 - Non-Proprietary Version)

4. Calculation NO.1 000320.314, "Residual Stress Analysis of Pressurizer Spray Nozzle with Weld Overlay Repair," Revision 0 (Attachment 4 Proprietary Version, Attachment 5 -

Non-Proprietary Version)

ATTACHMENT 2 Calculation No. 1000320.315

Structurallntegrify Associates, Inc.

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Table of Contents 1.0 OBJECTIVE .

2.0 DESIGN CRITERIA 4 3.0 LOADS 4 4.0 LOAD COMBINATIONS 5 5.0 ASME CODE STRESS LIMITS EVALUATION 6 5.1 Service Level AlB Load Combination 7 5.1.1 Simplified Elastic-Plastic Analysis 9 5.2 Thenna1 Ratcheting 9 6.0 FATIGUE EVALUATION 10 6.1 VESLFAT Program 1I 6.1.1 Cyclic Data (*. CYC) 1I 6.1.2 Fatigue Data Input File (*.FDT) 11 6.1.3 Stress Data Input File (*.STR) 12 6.1. 4 Fatigue Uwge (*. FAT) 14

7.0 CONCLUSION

S 14

8.0 REFERENCES

15 APPENDIX A SUPPORT FILES A-I APPENDIX B EXAMPLE VESLFAT FILES B-l File No.: 1000320.315 Page 2 of27 Revision: 0 F0306-01RI

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List of Tables Table 1: Transient Load Pairs that Consider OBE 16 Table 2: Bounding Transients for Analysis 17 Table 3: Bounding Piping Interface Loads 18 Table 4: Load Combinations 18 Table 5: Allowahle Stress Intensities 19 Table 6: Service Level AlB Load Combination, Primary-Plus-Secondary Stress Intensity Evaluation 20 Table 7: Simplified Elastic-Plastic Evaluation Results 21 Table 8: Event Cycles 22 Table 9: Materials for Fatigue Evaluation 22 Table 10: Fatigue Usage Results 23 List of Figures Figure 1. Stress Path Definitions of Minimum Weld Overlay for ASME Code,Section III Evaluations 24 Figure 2. Stress Path Definitions of Maximum Weld Overlay for ASME Code,Section III Evaluations 25 Figure 3. Tapered Transition Section 26 Figure 4. Isometric Drawing of the Spray Piping 27 File No.: 1000320.315 Page 3 of27 Revision: 0 FOJ06*0IRI

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1.0 OBJECTIVE The objective of this calculation package is to detennine if the ASME Code,Section III design requirements are satisfied for a weld overlay repair of the pressurizer spray nozzle-to-safe end dissimilar metal weld (OMWl) and the safe end-to-elbow dissimilar metal weld (OMW2) with weld overlay repair at Three Mile Island Nuclear Generating Station, Unit 1 (TMI-I). A design drawing of the weld overlay repair is provided in Reference I.

Two finite element models used to detennine structural and thennal operational stresses are developed in Reference 2. The two finite element models are constructed as 3-dimensional (3-D) "half-symmetry" (180 degrees) models. One of the 3-D models is constructed with the maximum weld overlay dimensions; whereas the other is constructed using minimum weld overlay dimensions. The model with the minimum overlay dimension is used for mechanical load evaluations. The model with the maximum overlay dimension is used for thennal stress evaluations. Further discussion of these finite element models can be found in Reference 2.

Several finite element stress and thennal analyses have been performed [3] to support the ASME Code evaluations. These analyses, together with the design requirements of the ASME Code [4], are used to detennine the adequacy of the repair.

2.0 DESIGN CRITERIA The weld overlay repair is designed to the requirements of the ASME Code,Section III, for Class I components. Thus, the rules of Article NB-3000 of Section III of the ASME Code, 2004 Edition [4] are used.

The weld overlay repair region affects the pressurizer spray nozzle, the safe end, and the attached spray piping. As a result, the nozzle portion of the repair at the pressurizer end will be evaluated using the rules of Subarticle NB-3200 of the ASME Code. For the attached safe end and piping, guidance from the rules of Subarticle NB-3600 of the ASME Code will be taken to satisfy NB-3200 acceptance criteria.

3.0 LOADS This evaluation only considers Service Level A and Service Level B operating conditions in regards to meeting ASME Code,Section III Service Level AlB allowables and fatigue. As such, thennal stresses resulting from Service Level C and Service Level 0 thennal transients are not considered [4, NB-3224.4 and Appendix F-13l0 (c)].

Primary stresses (such as mechanical loads due to deadweight, and seismic effects) resulting from Service Levels A, B, C and 0 operating conditions were previously evaluated in Reference 7, and are discussed in Section 4.0.

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Pressure Per Table 9 of Reference 6, the operating pressure loads range from 15 psia (0 psig) to 2,807 psia (2,792 psig) throughout the various thermal transients, whose temperatures range from 70°F to 682°F. The Hydro Test pressure ranges from 15 psia (0 psig) to 3,140 psia (3,125 psig) and the corresponding temperatures are 70°F. The pressures for various transients are summarized in Table 2. The gauge pressure (psig) is used in all calculations.

Thermal Transients Based on Table 9 of Reference 6, six bounding thermal transients (Plant Heatup (l A), Plant Cooldown (IB), Step Load Reduction! Reactor Trip Due to High Reactor Pressure/ Rod Withdrawal Accident (7 /8CIlI), Reactor Trip with Loss of Flow/Loss of Station Power (8A/15), Reactor Trip Due to High Reactor Temperature (8B), Rapid Depressurization (9)), Stratification Moment, and one test condition (Hydro Test) are considered. These thermal transients were evaluated in Reference 3. Details of the various transients are shown in Table 2.

The number of cycles shown in Table 2 are for the 60-year design operating period [6] for which the repair configuration will be evaluated.

Details of the thermal stress analyses are provided in Reference 3.

Mechanical Piping Loads The pressurizer spray nozzle is subjected to mechanical piping loads. These are defined in Table 3 of Reference 6. See Table 3 in this calculation for details. Note that Table 3 of Reference 6 also shows piping loads for the deadweight condition. However, deadweight loads are constant loads which occur for all load conditions. Therefore, deadweight loads do not contribute to the stress ranges for Service Levels A and B load combinations and fatigue evaluations (per NB-3222.2 and NB-3222.4 (for vessels) and NB-3653.1, NB-3653.2, and NB-3653.5 (for piping)) and are excluded for those evaluations.

4.0 LOAD COMBINATIONS The load combinations used in the repair design are:

1. Level A Load Combination

2. Level B Load Combination

3. Level C Load Combination

4. Level 0 Load Combination

5. Test Load Combination The weld overlay sizing evaluation [7] considered general primary membrane, Pm, and primary membrane-plus-bending, Pm+P b, stress intensities resulting from Service Levels A and B operating File No.: 1000320.315 Page 5 of27 Revision: 0 F0306-01RI

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conditions and Service Levels C and D conditions. The local primary membrane, PL , stress intensities are not specifically evaluated because the acceptability of Pm and P L+P b stress intensities indicate that PL stress intensities will be equally acceptable (local stress effects due to the WOL are expected to be minimal).

The sizing calculation does not specifically evaluate loads resulting from the Test Load Combination (Hydro Test). However, the Test Load Combination considers only primary stresses, which only result from pressure and mechanical loads. The added thickness of the weld overlay will only serve to reduce the general primary membrane, Pm, and primary membrane-plus-bending, PL +Pb, stress intensities (and as previously indicated, local primary membrane, PL , stress intensity) when compared to the original configuration. Therefore, the only stress category needed to be evaluated for stress acceptance is PL +Pb+Q criteria for all load combinations for Service Levels A and B. The specific load combinations are shown in Table 4. The allowable stress intensities for the primary + secondary stress category for these load combinations are shown in Table 5 [4]. Also, as indicated in Table 5, for Service Levels A and B, peak stresses and cyclic operation criterion must also be met.

Thus, this calculation, together with Reference 7, contains the ASME Code qualification for the weld overlay repair.

It should be noted that in using the ASME Code, Section 1II, Class 1 rules in NB-3200 (and NB-3600 rules for piping) [4], the bounding load combinations are used for evaluation of Service Levels AlB.

5.0 ASME CODE STRESS LIMITS EVALUATION Stress intensities are calculated for the various load combinations shown in Table 4 and compared to the allowable limits shown in Table 5. Linearized through-wall stresses are extracted through nine paths (Paths 1 through 9; see Figures 1 and 2) throughout the transient time histories and from the pressure and mechanical load analyses [3]. These calculated stress intensities are evaluated in accordance with ASME Code, Section 1II, Subarticle NB-3200 [4] for Paths 1,4, and 7, with guidance from Subarticle NB-3600 [4] for Paths 2,3,5,6,8 and 9.

Selection of the nine indicated paths was based on re-qualifYing the components impacted by the repair.

For this nozzle, the components in question are the spray nozzle, the safe end, and a portion of the spray piping. For the nozzle and piping, the critical locations are at the extreme ends of the weld overlay repair. Both locations, as a result of the repair, now have discontinuities, which impact the primary-plus-secondary-plus-peak stresses, and the corresponding fatigue usage. In addition, the pipe location will see the greatest impact in primary-plus-secondary stress ranges due to the thickness change resulting from the overlay and its impact on the thermal stresses per NB-3653.l.

The safe end was also selected as a separate component, though strictly speaking, the safe end is also a piping component per NB-1131(a). However, the safe end tends to have the greatest overlay thickness, which impacts the thermal secondary stresses, and is fabricated from a material weaker than the adjacent nozzle. In order to guarantee ASME Code compliance, the safe end was therefore included.

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5.1 Service Level AlB Load Combination Examination of the membrane-plus-bending stresses from Reference 3 does not provide an obvious pairing of stresses resulting from the various themlal transients for determination of the worst operating ranges. Thus, the VESLFAT program [8], developed by Structural Integrity, is used to calculate primary-plus-secondary membrane-plus-bending (P+Q) and total (P+Q+F) stress intensity ranges. The same program is used to perform the fatigue usage analysis described in Section 6.0.

The VESLFAT program computes stress intensity ranges based on component stress differences for all event pairs per NB-3216.2. It evaluates the stress ranges for primary-plus-secondary and primary-plus-secondary-plus-peak based on six component stresses (3 direct and 3 shear stresses). When more than one load set is defined for either of the event pair loading, the stress differences are determined for all of the potential loads, saving the maximum for the event pair. The principal stresses for the stress differences are determined by solving for the roots of the cubic stress equation per NB-3215.

The primary-plus-secondary membrane-plus-bending (P+Q) and total (P+Q+F) component stress values for thermal, pressure, and mechanical loads are combined prior to use in the VESLFAT program. The thermal component stresses resulting at each time increment from the various thermal transients are added to the component stresses resulting from corresponding pressure, and to the component stresses resulting from mechanical loading. The combination is performed in a series of Excel spreadsheets (file names are listed in Appendix A). Within the spreadsheet, the various component results are manipulated to produce the combined transient stress conditions, including:

• The primary-plus-secondary membrane-plus-bending (P+Q) and total (P+Q+F) stress components due to pressure are scaled from the 1,000 psi unit pressure evaluation performed in Reference 3. The actual pressure at specific time points for a given transient is defined in Reference 6 and shown in Table 2 of this calculation. The pressure between any two specified time points is assumed to vary linearly throughout each of the thermal transients.

• The primary-plus-secondary membrane-plus-bending (P+Q) and total (P+Q+F) stress components due to mechanical force and moment loads (resulting from thermal expansion) are calculated using the component stress results for a unit axial force (1,000 lb) and a unit moment (l,000 in-Ib) developed in Reference 3.

o Per Subparagraph NB-3653.l [4], piping force components (axial and shear) need not be included in the stress range determination and are, therefore, excluded for Paths 2, 3, 5, 6, 8, and 9 (safe end and piping).

o Equation 10 of Subparagraph NB-3653.1 provides a closed form solution to determine stress intensity range contributions for pressure and mechanical loads for Paths 2, 3, 5, 6, 8, and 9. However, this closed form solution is not used; rather, the actual stresses resulting from the finite element evaluations [3] are used.

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Per Reference 6, Section 4.3, the thennal expansion loads are applicable for a base fluid temperature of 555°F, and are scaled to other temperatures using a thennal load factor calculated as (T nu id-70)/(555-70), where Tnuid is the spray fluid temperature (Tspray ) in as listed in Table 2.

All forces related to piping mechanical loads are evaluated in the same manner for each event. The component forces are added together where appropriate and the resulting forces are then combined by SRSS to create a single bounding force load.

All moments related to piping mechanical loads are evaluated in the same manner for each event. The component moments are added together where appropriate and the resulting moments are then combined by SRSS to create a single bounding moment load.

Because it is not obvious which direction of applied moment loading produces the worst stress range, the evaluation considers both positive (Base case) and negative (Reverse case) loading directions for the thennal expansion loads and seismic loads in order to detennine the worst range.

o The thennal stratification loads are treated as separate event as shown in Table 8 with 7200 cycles. Depending on the resultant stresses, aBE is applied to this event as shown in Table 1.

o The aBE piping loads are defined in Section 3.0 of Reference 6 as having a total of 660 cycles. Therefore 660 cycles of aBE loading are applied to the thennal range loads with no self cycling. The aBE loads are added to the piping loads to one of the thennal transient pair events that produce the highest stress intensity range, as detennined on a per path basis. The transient load pairs on which aBE is being applied are shown in Table 1.

o Full range aBE is used for all paths to calculate the stress intensity range. Whereas half range aBE is considered while calculating fatigue.

Cyclic information, as well as material property data, are also needed to complete the VESLFAT input, though they do not playa direct role in the detennination of membrane-plus-bending stress intensity ranges. This data is needed to support the fatigue evaluations, and is discussed in detail in Section 6.0.

Table 6 presents the evaluation of the primary-plus-secondary stress intensity ranges for Service Level AlB operating conditions. The stress ranges extracted from VESLFAT files, with the extension *.FAT, are the stress intensity ranges that produce the greatest ratio of stress intensity range versus allowable stress.

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5.1.1 Simplified Elastic~Plastic Analysis Paths 6 (outside) and 9 (outside) have stress intensity which exceed the allowahle stress intensity range. A simplified elastic-plastic analysis is conducted per Subparagraph NB-3653.6 [4], since the 3S m limit is exceeded.

• NB-3653.6, Part (a), requires that "Equation (12) shall be met." Equation (12) is defined as:

C M* 3S m 21 I In Equation (12), M i ' is identified as the moment range that includes only the thermal expansion.

The largest thermal moment range is 173,263 in-Ibs at a fluid temperature range of 70°F to 636°F (extracted from Excel Spreadsheet " SectionlII_v120- TM1-SPRA Y-no-obe.xls" tab "Eq-12") for Path 6 Outside. The largest thermal moment range for Path 9 outside is 7604 in-Ibs at a fluid temperature range of 70°F to 636°F. These values are used to calculate the membrane-plus-bending stress intensity results from the 1,000 in-Ib unit moment results in Reference 3. For Path 6 outside, the unit out-of-plane moment results are used, since stratification is present.

According to the isometric drawing in Figure 4 for spray piping, it can be seen that the Stratification is being applied in the Out-of-plane direction on the nozzle. The resulting stress intensities are shown in Table 7 with comparison to the allowable values.

• NB-3653.6, Part (b), requires that "primary plus secondary membrane plus bending stress intensity, excluding thermal bending and thermal expansion stresses, shall be < 3S m ."

Equation (13) ofNB-3653.6, Part (b) is essentially identical to Equation (10) ofNB-3653.1. To bound the various load combinations, the membrane-plus-bending stress intensities for the worst thermal transient pressure (2792 psig for the 7/8CIlI transient) were conservatively added to the maximum membrane stress intensity range that occurs for any two thermal transients (for Equation (13), the thermal piping moment loads are excluded). The results are listed in Table 7.

5.2 Thermal Ratcheting The thermal stress ratchet is required to be evaluated to prevent cyclic growth in the component. The thermal stress ratchet effect is driven by internal pressure, as the component is subject to cyclic thermal stress. Paths 3, 6, 9 at the attached pipe location, are not protected by a thermal sleeve and have the thinnest cross-section, which produces the most conservative ~TI allowable. These paths are selected for thermal ratcheting evaluation per Subparagraph NB-3653.7 [4]. Therefore, the limiting range of through-wall temperature gradient is calculated as:

x ~(_l) 2792*4.5 ( I ) = 0.48 (bounding Path 3) 2* t Sy 2*0.438 29869 File No.: 1000320.315 Page 9 of27 Revision: 0 F0306-01RI

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where, P Maximum pressure for the set of conditions under consideration= 2,792 psig (see Table 2)

Do Outside diameter, inches for the pressurizer spray pipe [2].

t Wall thickness, 0.438 inches for the pressurizer spray pipe [2].

Sy Per Note 11 of Subparagraph NB-3222.5, I.5S m can be substituted for Sy, if greater. The value of 1.5S m is 29,869 psi [5] for SA-403, WP 316 (Path 3) at an average fluid temperature of 31 '" (Note: *' The pair with the maximum LlT I range is the Plant Heatup transient. The average fluid temperature is 313"F (555+70)/2). 313"F is used to calculate the value of l.5S m)

Therefore, the limiting range of ~TI can be calculated as:

AT Ll j

_ . (*C 4 ) -_ 2.08*29869 .3) 479 . 65°bF ( oun d'mg Path 3) 0.7* Ea 0.7*28.3*8.5 where, y' 2.08 (Per NB-3222.5 [4]; y' = 1/x = 2.08, for x 0.48).

C4 1.3 for stainless steel material E Modulus of elasticity, 28.3e6 psi, for SA-403, WP 316 at room temperature [5].

a Mean coefficient of thermal expansion, 8.5e-6 in/in/OF, for SA-403, WP 316 at room temperature [5].

The inside and outside surface temperatures are extracted from the thermal transients evaluated in Reference 3. The through-wall temperature difference (~T) is calculated for each time point of the transients. The maximum positive through-wall ~ T is subtracted from the minimum through-wall ~ T for the two limiting transients, and the resulting range conservatively considered the ~ T I range (see Excel Spreadsheet" SectionIII_v120-TMI-SPRAY-no-obe.xls If, Tab "ThermaIRatchet"). ~TI is defined in Subparagraph NB-3653.2 as the range of the temperature difference between the temperature of the outside surface and the temperature of the inside surface assuming moment generating equivalent linear temperature distribution. The maximum ~T for Paths 1 through 9 is 112°F, which is below the allowable temperature range of 480°F. Therefore, the thermal ratcheting criterion is met for all paths.

6.0 FATIGUE EVALUATION The fatigue evaluations are performed for Paths 1 through 9 for the weld overlay repair (see Figures 1 and 2). Both the inside and outside locations of the indicated paths are evaluated. The evaluations are performed in accordance with ASME Code,Section III, Subparagraph NB-3222.4(e) [4] (with guidance from NB-3600), using the Structural Integrity developed VESLFAT program [8].

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6.1 VESLFAT Program The VESLFAT program requires three input files. The first is the *.CYC file, which includes the number of cycles for each load combination. The data used in the *.CYC file is discussed in detail in Section 6.1.1. The second file is the *.FDT file, which includes the fatigue curve data, appropriate temperature dependent material properties, and simplified elastic-plastic limits and factors. These values are discussed in Section 6.1.2. The final input file is the *.STR file, which contains the component membrane-plus-bending and membrane-plus-bending-plus-peak (i.e., total) stresses for the various load conditions to be evaluated. Additional details are provided in Section 6.1.3. As several load conditions occur within each load case, these load conditions will be identified by a number, which matches the load condition to the load case. This number is defined in the *.CYC file. Each of these three files must be identically named, with the exception of the file extension.

A number of intermediate files are generated, which can be used to check the final results. The *.STI file is an echo output of the *.STR file but includes transformations to output the results in terms of psi.

The *.ALL file reflects all of the stress range pairs that are calculated. The *.PR file is a shortened version of the *.ALL file and lists only the significant (i.e., fatigue causing) pairs. The *.ORD file re-sequences the *.PR file such that the ordered pairs are arrayed in order of reducing alternating stress.

The actual final output file is labeled *.FAT. It echoes the input data, shows the significant cycle pairings, the cycle elimination, individual cycle pair fatigue contributions, and the final overall fatigue usage. See Section 6.1.4 for fatigue results.

6.1.1 Cyclic Data (*.CYC)

Reference 6 assigned a total number of cycles for each bounding event, which are tabulated in Table 8 of this calculation. See Appendix B for an example of a *.CYC file.

6.1.2 Fatigue Data Input File (*.FDT)

The materials at the surfaces of the stress paths indicated in Figures 1 and 2 are tabulated in Table 9.

The fatigue curve for the stainless steel and Alloy 52M components is per Reference 4,Section III Appendices, Table 1-9.2.

The fatigue curve for the SA-508 Class 1 material is also per Reference 4,Section III, Appendices, Table 1-9.1.

The modulus of elasticity correction factor from the fatigue curves are based on Reference 5 temperature dependent modulus of elasticity values with a fatigue curve elastic modulus of 28.3e6 psi for the austenitic and Alloy 52M materials, and 30.0e6 psi for the SA-508 Class 1 material.

See Appendix B for an example *.FDT file.

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6.1.3 Stress Data Input File (*.STR)

Linearized membrane-plus-bending (P+Q) and membrane-plus-bending-plus-peak (P+Q+F) component stresses from the finite element stress analyses [3] were extracted for pressure, mechanical, and thermal transient loads. Stresses are scaled in cases (pressure and mechanical) where the applied load magnitude is not the same as that analyzed.

The resulting component stresses are then added to create the load combination for each thermal transient throughout the length of the event. Thus, thermal stresses are added to the scaled pressure stresses and to the scaled mechanical load stresses to create each membrane-plus-bending and membrane-plus-bending-plus-peak component stress entry.

Paths I, 4, 7, 3, 6, and 9 terminate on the outside at locations with geometric discontinuities. For these locations, the magnitude of the P+Q+F stress is determined by applying a fatigue strength reduction factor to the P+Q stress obtained from the finite element analysis (Reference 3). This is allowed per NB-3222.4(e)(2) [4] which allows for the use of theoretical techniques for the determination of a stress concentration factor. For only the inside locations (in contact with the fluid) with a fatigue strength reduction factor, the peak thermal stress components are added back into the total stress to capture the peak stress due to non-linear temperature gradients. The use of the fatigue strength reduction factor is shown as follows.

P+Q+F = (ANSYS membrane + bending)*FSRF + ANSYS thermal peak (Inside Locations)

P+Q+F (ANSYS membrane + bending)*FSRF (Outside Locations)

For those paths that do not occur at a geometric discontinuity, no fatigue strength reduction factor is used. Instead, the membrane-plus-bending-plus-peak (P+Q+F) component stresses from the finite element stress analyses [3] will be used directly.

Similarly, a fatigue strength reduction factor is applied to Paths 2, 5 and 8, inside which are located at areas of tapered transition. Note that Paths 2 and 5 are located at areas of tapered transition on the outside surface for the maximum weld overlay case only.

Fatigue strength reduction factors as shown in the following table are calculated and applied to the affected path locations as detailed in Section 6.2 of Reference 9. A pictorial representation of the tapered transition section is provided in Figure 3.

The fatigue stress reduction factor is calculated as follows:

. S Rd' 2tana Fattguetress e uctIOn Factor =( I. )

a+-sm2a 2

where a is the tapered transition angle in radian.

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Tapered transition Fatigue Strength Path No. Surface angle (a.) in degrees Reduction Factor Inside 0 1.00 1

Outside 30 1.21 Inside 18.4 1.07 2

Outside 18.5 1.07 Inside 0 1.00 3

Outside 24.7 1.13 Inside 0 1.00 4

Outside 30 1.21 Inside 18.4 1.07 5

Outside 14 1.04 Inside 0 1.00 6

Outside 29.8 1.20 Inside 0 1.00 7

Outside 30 1.21 Inside 18.44 1.07 8

Outside 0 1.00 Inside 0 1.00 9

Outside 49.1 1.41 The *.STR file includes the temperature of the location as it varies throughout the events and the pressure. The pressures vary as indicated in Table 2. The temperature at the stress path location is based on the actual metal temperature of the material rather than the fluid temperature. These metal temperatures were extracted from the prior stress evaluations in Reference 3 via the linearized stress results files which include the temperature data in the last field under "Total" stress.

Because it is not obvious as to which direction of applied force and moment loading produces the worst stress ranges, the fatigue evaluations will be considered in both positive (Base Case) and negative (Reverse Case) mechanical load directions, in order to capture the worst ranges and, therefore, the greatest fatigue usage.

The load combinations and the development of the *.STR file entries were performed III Excel spreadsheets named which are listed in Appendix A.

An example of the *.STR file is shown in Appendix B.

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6.1.4 (*,FAT)

The fatigue evaluation automatically the load pairs that create the greatest stress intensity range, performs a calculation, corrects for the modulus of elasticity, and performs the fatigue evaluation. It repeats this process selecting the next highest stress range until the available cycles are used up or the remaining stress fall below the endurance limit. An example *.FAT file is included in Appendix B. The intermediate solution files *.STl, *.ALL, *.PR, and *.URD are included with computer files.

Table 10 tabulates the total fatigue usage for each location. In addition, the table includes information on the load pairing which produces the alternating stress for each location, including the membrane-plus-bending stress intensity range, the calculated Ke elastic-plastic factor, and the alternating stress, Sa, for the specific load pair.

7.0 CONCLUSION

S An evaluation of the pressurizer spray nozzle weld overlay repairs for the Three Mile Island Nuclear Generating Station, Unit 1 (TMI-l) has been performed in accordance with the requirements the ASME Boiler and Pressure Vessel Code,Section III, for Class 1 components [4]. Stress intensities were conservatively determined for pressure, piping loads and bounding thermal transients provided in Reference 6, and compared to ASME Code allowable values for primary-plus-secondary stress effects.

In all cases, except for Path 6 outside and Path 9 outside, the reported values of stress intensity ranges are less than their corresponding allowable values (see Table 6). A simplified elastic-plastic analysis is conducted on Paths 6 outside and 9 outside, per Subparagraph NB-3653.6 [4], since the 3S m limit is exceeded. The resulting stress intensities are shown in Table 7 with comparison to the allowable values and it can be seen that the stress intensity values are less than their corresponding allowable values.

A detailed fatigue analysis was also performed. For the given number of expected cycles (see Table 8),

the total usage at all locations are less than the allowable of 1.0 (see Table 10).

In conclusion, the pressurizer spray nozzle weld overlay repair design shown on the design drawing [1]

satisfies the requirements of the ASME Code, and is qualified for the 60 years of cyclic operation.

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8.0 REFERENCES

l. SI Drawing No. 1000320.510 "Pressurizer Spray Nozzle Full Structural (FSWOL) Weld Overlay Design Drawing," (for revision number refer to SI Project Revision Log, latest revision).

2. SI Calculation No. 1000320.31 "Material Properties and Finite Element Models for Pressurizer Spray Nozzle with Weld Overlay Repair," (for revision number refer to SI Project Revision Log, latest revision).

3. SI Calculation No. 1000320.313, "Pressurizer Spray Nozzle Thermal and Mechanical Stress Analysis Calculation," (for revision number refer to SI Project Revision Log, latest revision).

4. ASME Boiler and Pressure Vessel Code, Section HI, Rules for Construction of Nuclear Facility Components, 2004 Edition.

5. ASME Boiler and Pressure Vessel Code,Section II, Part D, Material Properties, 2004 Edition.

6. SI Calculation No. 1000320.311, "Design Loads for Pressurizer Spray Nozzle with Weld Overlay Repair," (for revision number refer to SI Project Revision Log, latest revision).

7. SI Calculation No. 1000320.310, "Pressurizer Spray Nozzle Weld Overlay Sizing Calculation,"

(for revision number refer to SI Project Revision Log, latest revision).

8. VESLFAT, Version 1.42, Structural Integrity Associates, Inc., February 6, 2007.

9. J. F. Harvey, "Theory and Design of Modem Pressure Vessels," 2nd Edition, Van Nostrand Reinhold Company, 1974.

10. GPU Nuclear Drawing No. ID-212-23-028, Sheet 2 of 6, Rev. I, "LPSIIDECAY HEAT REMOVAL PIPING ANALYSIS," SI File Number 1000320.204.

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S}Structurat tntegrity Associates, Inc, Table 1: Transient Load Pairs tbat Consider OBE Primary-Plus-Secondary Fatigue Location Base Reverse Base Reverse Path I Inside 10 6 10 10 Path I Outside 6 ")

"- 6 6 Path 2 Inside 9 9 9 9 Path 2 Outside 3 6 3 6 Path 3 Inside 10 6 10 6 Path 3 Outside 6 3 6 3 Path 4 Inside 10 9 10 10 Path 4 Outside 9 2 9 6 Path 5 Inside 9 9 6 6 Path 5 Outside 9 9 9 9 Path 6 Inside 6 10 10 6 Path 6 Outside 9 9 9 9 Path 7 Inside 8 10 10 ]0 Path 7 Outside 2 6 6 6 Path 8 Inside 9 I 6 3 Path 8 Outside 3 6 3 6 Path 9 Inside 9 9 9 9 Path 9 Outside 3 6 3 6 Note:

I) The event ID # is identified in Table 8.

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Table 2: Bounding Transients for Analysis 10800 370 370 173 1378 159 1.000 10800 370 70 22 21 216 1.000 12240 410 110 26 26 319 0.082 12240 410 110 141 427 319 0.082 27432 650 532 160 739 2242 0.953 27432 650 532 183 2149 2242 0.953 10800 430 300 177 1540 216 0.474 21600 400 280 177 1484 216 0.433 21600 400 210 152 578 159 0.289 25920 330 150 147 494 15 0.165 0

(bounds 0 650 566 182 2111 2242 1.023 2A, 3, 4, 4 682 586 182 2088 2792 1.064 6,7,8e, 12 682 590 182 2084 2792 1.072 10, II, 144 682 590 182 2084 2792 1.072 14, and 144 639 575 182 2101 2092 1.041 20B) 192 633 562 182 2116 1997 1.014 282 625 546 183 2134 1897 0.981 8 662 570 174 1261 2442 1.031 12 662 570 174 1261 2442 1.031 20 650 555 164 844 2242 1.000 30 632 550 154 612 1992 0.990 720 643 550 154 612 1992 0.990 720 643 636 173 1405 2142 Ll67 (also 1 651 544 183 2136 2260 0.977 bounds 4 654 562 182 2116 2312 1.014 17A) 6 657 564 182 2113 2348 1.019 10 661 560 182 2118 2419 1.010 15 666 576 182 2100 2507 1.043 22 653 568 182 2109 2297 1.027 22 653 568 41 44 2297 1.027 600 616 558 41 44 1757 1.006 (also 900 519 500 2093 842 0.887 Notes: (l) Table is from I 3.

(2) Temperatures are assumed to vary linearly with time between indicated time points.

(3) At the region covered by the thermal sleeve.

(4) At all regions, including pressurizer shell, except thermal sleeve.

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Table 3: Bounding Piping Interface Loads Forces,lb Moments,lb-in aBE 422 10932 20102 15723 Thermal 6 6468 704 348 Maximum Stratification Moment (2) Range = 173,263 Ib-in Note: (I) The above table is reproduced from Table 3 of Reference 6.

(I) The Stratification occurs at 555°F and at a pressure of 2242 psig [6].

Table 4: Load Combinations Load Combination LOADS Level A Level Pressure Note 1 Note 1 Temperature Note 2 Note 2 Mechanical Piping Loads X X Thermal Transients Plant Heatup X Plant Cooldown X I Step Load Reduction! Reactor Trip Due to High Reactor X Pressure/ Rod Withdrawal Accident Reactor Trip with Loss of Flow/Loss of Station Power X Reactor Trip Due to High Reactor Temperature X Rapid Depressurization X Stratification Moment X J§::dro test Note 3 Note 3 Notes:

1. Varies between 15 psia and 3140 psia depending on transient conditions [6], summarized in Table 2.

2. Varies between 70°F and 682°F depending on transient conditions [6], summarized in Table 2.

3. Hydro Test pressure varies between 15 psia and 3140 psia. Hydro test is covered under NB 3226, but included here conservatively.

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Table 5: Allowable Stress Intensities Load No (3) (3) (3) 3.08 m Notes:

1. The requirements of ASME Code, Section Ill, Subparagraph NB-3222.4(e) [4] (and NB-3653.5 for piping) for peak stresses and cyclic operation must be met.

2. Pc is not specifically listed and is included with Q. Pc is defined in Figure NB-3222-l [4] as stresses which result from the constraint of free end displacement. The piping loads provided in Reference 6 do not specifically detail the source of the loading (thermal expansion or thermal anchor movements for the thermal loads).

Therefore, the loads are all classified as secondary, Q.

3. Not evaluated as explained in Section 4.0.

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Table 6: Service Level AlB Load Combination, Primary-Plus-Secondary Stress Intensity Evaluation Maximum Maximum Patb Stress Intensity Stress Intensity Allowable Number(7) Surface 3S m (pSi)(6)(I1) Accept Range (So) Range (So)

(psi) Base(8)(IO) (psi) - Reverse(9)(lO)

Inside(l) 17502 17838 49066 Yes 1 Outside(L) 20145 19443 53872 Yes Inside(J) 19256 19028 69900 Yes 2 Outside(4 ) 20935 21278 69900 Yes Inside(5) 33832 34953 506]4 Yes 3 Outside(5) 369]4 35839 516]8 Yes Inside(l) 17213 ]7937 49066 Yes 4

OutsidelL ) 26722 ]96]3 53872 Yes nside lJ ) 23697 21822 69900 Yes 5 side(4 ) 21225 24480 69900 Yes Inside(5) 32]43 31601 50407 Yes 6 Outsidel.) 77277 67004 5]642 No Insidell ) ]6896 1675] 49066 Yes 7 Outside(L) ]9662 20334 53872 Yes Inside(J) 19835 20681 69900 Yes 8

Outside(4 ) 20577 20943 69900 Yes Inside(5) 35982 37707 52350 Yes 9 Outside(5) 48853 52459 51615 No Notes:

L Material at location is SA-240 TP 304 [3].

2. Material at loeation is SA-508 Class I [3].

3. Material at location is SB-166, treated as Alloy 600 [3J.

4. Material at location is Alloy 52M [3J.

5. Material at location is SA403 WP-316 [3J.

6. All material stress allowables [4J are based on the maximum Sn/3Sm ratio from VESLFAT output files ending in *.FAT (see Appendix B for example).

7. See Figures I and 2 for illustration of indicated locations.

8. Mechanical loads are applied in the positive direction (Base Case).

9. Mechanical loads are reverse from the Base Case (Reverse Case).

10. Sn are based on the maximum Sn/3Sm ratio from VESLFAT output files ending in *.FAT (see Appendix B for example).

I I. The calculation of 3Sm is performed by VESLFAT using the element temperatures along the selected path and interpolating between the ASME Code,Section II values.

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Table 7: Simplified Elastic-Plastic Evaluation Results Path(3) Surface Accept Notes:

I. Material at location is SA-403 WP316 [1].

2. Material stress allowabies evaluated are based on those generated in Table 6.

3. See Figures I and 2 for illustration of indicated locations.

4. Equation 12 evaluation can be found in Excel Spreadsheet" Sectionl11~vI20-TMI-SPRAY-no-obe.xls"in the tab "Eq-12." Stress intensity results were conservatively used.

5. The calculation on Sm is performed by VESLFAT using the element temperatures along the selected Path and interpolating between the ASME Code,Section II values [5J.

6. Equation 13 evaluation can be found in Excel Spreadsheet" Sectionlll~vI20-TMl-SPRAY-no-obe.xls"in the tab "Eq-13."

7. The calculation of 3S m is performed by VESLFAT using the element temperatures along the selected path and interpolating between the ASME Code,Section II values.

8. Moment stress from out of plane moment (Mx) is used to calculate Equation 12, since stratification load is responsible for the Service level AlB failure.

9. Moment range used to calculate this stress is taken from events 1 through 8, excluding stratification event, as this is not the event which contributes to maximum stress intensity range in Table 6.

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Table 8: Event Cycles Transient Event(l) rvrlp!il(.:=

Plant Heatup 1 240 Plant Heatup 2 240 Plant Cooldown 3 240 Step Load Reduction! Reactor Trip Due to High Reactor Pressure/ Rod Withdrawal 4 65938 Accident Reactor Trip with Loss of Flow/Loss of 5 80 Station Power Reactor Trip with Loss of Flow/Loss of 6 80 Station Power Reactor Trip Due to High Reactor 7 180 Temperature Rapid Depressurization 8 1480 Stratification Moment 9 7200 Hydro Test 10 20 OBE - 660 Notes:

1. Used by VESLFAT to identify event groupings.

2. The cycles provided in Table 9 of Reference 6 are for 60 years of operation.

Table 9: Materials for Fatigue Evaluation Path(l) Surface Material(Z)

Inside SA-240 TP304 1,4,7 Outside SA-508 Class 1 SB-166 (Treated as Inside 2,5,8 Alloy 600)

Outside Alloy 52M Inside SA-403 WP-316 3,6,9 Outside SA-403 WP-316 Notes:

I. See Figures I and 2 for illustration of indicated locations.

2. Identified in Reference 3.

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Table 10: Fatigue Usage Results Maximum Alternatin2 Stress Load Pair Path Total Surface Event So Alternating Fatigue Usage(3)

NO.(I)

Pair(4) (psi)

Ke Stress (psi)

Inside 6&10 16008 1 24472 0.0004334 (.2) 1 Outside 2&6 20233 I I 14077 0.0010798 Inside 1&9 19256 1 --- 0 2

Outside 3&6 21278 1 --- o (~)

Inside 6&10 34953 1 20109 0.000079 (.2) 3 Outside 3&6 37082 1 23263 0.0003368 Inside 6&10 15911 1 24398 0.0004795 (.2) 4 Outside 2&9 26722 1 18155 0.0041867 5 ~side 6&9 16862 1 15103 0.00003 utside 6&9 24480 I --- 0.0000000 (L)

Inside 1&6 30584 1 16873 0.0000766 (.2) 6 Outside 3&9 72648 2.293 110595 0.6640993 Inside 6&10 16704 1 24613 0.0004269 (2) 7 0.0011338 (L)

Outside 2&6 20864 1 14259 Inside 3&6 9050 I 14935 0.0000092 8 0(.2)

Outside 3&6 20943 1 ---

Inside 1&9 37707 1 20661 0.0005513 (2) 9 0.0080245 Outside 3&6 52459 1.055 43411 (2)

Notes:

I. See Figures I and 2 for illustration of indicated locations.

2. Reversed Piping Load Case.

3. Cumulative fatigue usage from all contributing load pairs.

4. Refer to Table 8 for event identification.

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Figure 1. Stress Path Definitions of Minimum Weld Overlay for ASME Code,Section III Evaluations (Figure isfrom Reference 3)

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F,r F:oflrr,:;

'-\i','!' In!

Figure 2. Stress Path Definitions of Maximum Weld Overlay for ASME Code,Section III Evaluations (Figure isfrom Reference 3)

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Figure 3. Tapered Transition Section File No.: 1000320.315 Page 26 of27 Revision: 0 F0306-01Rl

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o c c

~

~_---,-_---,;_ "'*-.---i.=~~==J_---l;:;;;~~=';;'=....d::4-J Figure 4. Isometric Drawing of the Spray Piping (Note: Figure is obtainedfrom Reference 10)

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APPENDIX A SUPPORT FILES File No.: 1000320.315 Page A-I of A-2 Revision: 0 F0306-01 Rl

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File Name Description Excel e' used to Create *STR Files for use in VESLFAT Fatigue Program.

SectionllI v120-TMI-SPRAY-$$.xls

~

($$ no-obe, Pair-I, Pair-2, Fatigue-Pair-I, Fatigue-Pair-2)

Cycle Data Input File for Inside Surface Locations for VESLFAT Program where, P*-I.CYC

• ' Paths I through 9. 'r' in file name indicates Reverse Piping Load Case.

Cycle Data Input File for Outside Surface Locations for VESLFAT Program where, P*-O.CYC

• ' Paths I through 9. 'r' in file name indicates Reverse Piping Load Case.

Material Data Input File for Inside Surface Locations for VESLFAT Program where, P*'-I.FDT

• ' '" Paths I through 9. 'r' in file name indicates Reverse Piping Load Case.

Material Data Input File for Outside Surface Locations for VESLFAT Program P*'-O.FDT where,

• ' Paths I through 9. 'r' in file name indicates Reverse Piping Load Case.

Stress Data Input File for Inside Surface Locations for VESLFAT Program P*'-I.STR Created by Excel Spreadsheets Listed Above, where,

• Paths I through 9. 'r' in file name indicates Reverse Piping Load Case.

Stress Data Input File for Outside Surface Locations for VESLFAT Program P*'-O.STR Created by Excel Spreadsheets Listed Above, where,

• Paths I through 9. 'r' in file name indicates Reverse Piping Load Case.

P*-I.ALL Intermediate Result File for Inside Surface Locations Created by VESLFAT Program P*'-I.ORD where, P*-l.PR

• ' Paths I through 9. 'r' in file name indicates Reverse Piping Load Case.

P*'-I.STI P*-O.ALL Intermediate Result File for Outside Surface Locations Created by VESLFAT P*-O.ORD Program where, P*'-O.PR

• ' Paths I through 9. 'r' in file name indicates Reverse Piping Load Case.

P*-O.STI Fatigue Result File for Inside Surface Locations Created by VESLFAT Program P*'-LFAT where,

• ' '" Paths I through 9. 'r' in file name indicates Reverse Piping Load Case.

Fatigue Result File for Outside Surface Locations Created by VESLFAT Program P*'-O.FAT where,

• Paths I through 9. 'r' in file name indicates Reverse Piping Load Case.

File No.: 1000320.315 Page A-2 of A-2 Revision: 0 F0306-01Rl

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APPENDIXB EXAMPLE VESLFAT FILES File No.: 1000320.315 Page B-1 ofB-18 Revision: 0 F0306-01Rl

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VESLFAT INPUT FILE P3-0.CYC:

N Name N 1 TMI ~

1 40 2 n11 40 TMI ~

3- 660 4 TMI 4- 5938 5 80 6 TMI 6- 80 7 80 8 TMI 8-

~

1480 9 9- 7 00 0 TMI 10-

~

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VESLFAT INPUT FILE P3-0.FDT:

i 00 402.2

">94.2 429,8 575.9 70 "Ill 811 00 5.3 10.6 1U,9 e05.,,> 99 15.9 8116.? HO *. 6 lOll 21,2 toO,9 13)7.1 Ull 26,5 1210.0 1580,1 129 JLS 1385.8 lB36.9 139 37,1 ISH.l nO'.J 149 42 .*

1728,3 2335.S 159 47,1 File No.: 1000320.315 Page B-3 ofB-IS Revision: 0 F0306-01RI

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i" co 1:  ::

i,

i-.  ;",

U Hi n .~

n  ::

~ile . 1 11l::

Page B-4 ofB-i8 Revision: 0 F0306-01RI

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H  ::  ;  :

.~

i l

• • H

~

~

T

.~  !

File No: : 115 Page B-5 ofB-18 Revision: 0 F0306-01RI

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U i  :::  :~

U

~; ,

File N . 1 11';;

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,  :: i:

j

ile 11I11I ~ '.11 lll

i Page B-7 ofB-I8 Revision: 0 F0306-01RI

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VESLFAT RESULT FILE P3-0.FAT:

1. 2 / 9/2006 (Ves Ip42 )

19: 9 ion 1.4 - 01/03/ 007 (&VeslFatPairlp42)

Stres to convert to ps

["lax ( i) per unit Pressure (psi) = 0.01 Upper Limit on Sy for large number of cycles NB-5222. 5 (b) (ksi) 40 Temperature T, E, ks 3Sm, ks Sy, i 70 28 00 20.0 30.0 00 7500 20.0 25.9 300 27000 20.0 3.4 400 6400 19.3 21.4 500 25900 18.0 20.0 600 5 00 17.0 18.9 650 2 0 a 16.6 18.5 700 4800 16.3 18.2 7 0 4450 16.1 17.9 Stress ranges 13328 psi neglected Files:

Input Stress File p3-0.STR Converted Stress File p3-0.STI All Stress Ranges File p3-0.ALL S ficant Ranges File p3-0.PR Thermal Ratchet Case File p3-0.TRC File No.: 1000320.315 Page B-8 of B-18 Revision: 0 F0306*01Rl

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1 1

ion 1. 4 Ie File No.: 1000320.315 Page B-9 of B-18 Revision: 0 F0306-01RI

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1.4 12/ /200 1 9: 7 1.0E+01 708.00 1 1 .00

.OE+01 4 .00 1.0E+0 61.00

.OE+O 01.00

.OE+02 148.00 1.0E+0 119.00

.OE+03 97.00 5.0E+03 76.00 1.0E+04 64.00

5. 0

.OE+04 46. 0 1.0E+05 40.80

.OE+05 .90

.OE+O 1. 00 1.0E+06 28.20

.OE+06 22.80 5.0E+06 18.40 1.0E+07 16.40

.OE+07 1 .20

.OE+07 14. 0 1.0E+08 14.10 1.0t:+09 1 .90 1.0E+ 0 13.70

1. OE+11 13.60 Events: Index Name Num. Cycles 1 TMI 240 2 TMI 2- 240 TMI 3- 660 4 TMI 4- 65938 5 TMI 5- 80 6 TMI 80 7 TMI 7- 180 8 TM1 1480 9 TMI 9- 7200 10 TMI 10- 20 File No.: 1000320.315 Page B-lO ofB-18 Revision: 0 F0306-01Rl

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e Strocturat Integrity Associates, Inc, File No.: 1000320.315 Page B-I2 ofB-I8 Revision: 0 F0306-01Rl

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S3 Stlucturallnteglily Associates, Inc.

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lJ Structurat tntegrity Associates. Inc, Intens .ORD w/E-modulus

1. 1 # Sn Salt 1 TMI 6 TMI 6- 29795 1.0000 18692 6 and 10 TMI 10- 29795 1.0000 18692 L

" and 6 TMI 6- 29596 1.0000 18567 TMI - 7494 .0000 17249 4 and 10 TMI - . 10 26564 1.0000 16619 1 and 4 TMI 4- 26564 1. 0000 16619 4 TMI 4- 6 67 1.0000 16496 1 and TMI 2 47 1.0000 15929 VESLFAT Module ion 1. 42 - 12/29/2006 11-04-2010 12:19:47 s Input (ps - Sal w/E-modulus correction)  :

1. Eventl # 2 Sn Ke Salt 3 TMI and 10 TMI 25473 1.0000 15929 3 TMI and 4 TMI 4 25447 1.0000 15910 TMI - and TMI 3- 25270 1.0000 15802 TMI - l- and 7 TMI 7- 23691 1.0000 14803 7 TMI 7 and 10 TMI 10- 3691 1.0000 14803 2 TMI and 7 TMI 7- 23494 1.0000 14680 TMI 3- and 7 TMI 7- 22589 1.0000 14112 5 TMI - s- and 10 TMI 10- 21596 1.0000 13513 1 TMI - 1 and TMI 5- 21596 1.0000 13513 2 TMI 2- and 5 TMI 5- 21401 1.0000 13391 File No.: 1000320.315 Page B-17 ofB-18 Revision: 0 F0306-01Rl

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tVfodule ion 1.4 12/ I 00 010 1 :19:48 I Na lowed 40 97 1.000 18692 4.674 80 4- 5 8 6 64 1.000 1661 9. 10E+0 .000002 10 TM1 0 TMI 1 160 6 64 1.000 16619 9. 310E+06 .0000 7 4 nu . - 4 6 918 4 1't11 40 6367 1. 000 16496 9. 56E+06 .0000249 4 TM1 4- 7 8 TJ'vlI 660 25447 1.000 1 910 1.3185E+07 .0000 01 4 4 18 Total Usage 0.0001115 File No.: 1000320.315 Page B-I8 ofB-I8 Revision: 0 F0306-01RI

ATTACHMENT 3 Calculation No.1 000320.316

~Sirlfelurallntegrity Associates, Inc.~ File No.: 1000320.316 Project No.: 1000320 CALCULATION PACKAGE Quality Program: [8J Nuclear 0 Commercial PROJECT NAME:

TMI-l Pressurizer Spray Nozzle WOL CONTRACT NO.:

59091 CLIENT: PLANT:

Aquilex WSI, Inc. Three Mile Island Nuclear Generating Station, Unit 1 CALCULATION TITLE:

Crack Growth Evaluation of Pressurizer Spray Nozzle with Weld Overlay Project Manager Preparer(s) &

Document Affected Revision Description Approval Checker(s)

Revision Pages Shmature & Date Signatures & Date o I 34 Initial Issue A-I - A-3 I

Norman Eng NE 03/23/11 Apama Alleshwaram AA 03/23/11 (j~:n, .

Ashwin Padmala AP 03/23/11 Page I of34 F0306-01Rl

Table of Contents 1.0 PURPOSE ..

2.0 METHODOLOGY 4 3.0 DESIGN JNPUTS 4 3.1 (Jeometry 4 3.2 Fracture Mechanics Models 5 3.3 Loads 6 3.4 Crack Growth Laws 9 3.4.1 Alloy 82/182 Fatigue Crack Growth Law 9 3.4.2 Alloy 82/182 PWSCC Growth Law 9 3.4.3 Treatment a/Negative Stress Intensity Factors 10 4.0 ASSUMPTIONS 10 5.0 CALCULATIONS 10 6.0 RESULTS AND CONCLUSIONS 12

7.0 REFERENCES

13 APPENDIX A COMPUTER FILE DESCRIPTIONS A-I List of Tables Table I: Stress Coefficients for Various Loadings at the DMW I, Safe End and DMW2 14 Table 2: Spray Nozzle Piping Interface Loads 17 Table 3: Bounding Thermal Transients for the Spray Nozzle 18 Table 4: Sequence of Events and Cycles for Fatigue Crack Growth 19 Table 5: Crack Growth Results 20 File No.: 1000320.316 Page 2 of34 Revision: 0 F0306-01Rl

List of Figures Figure 1: FEM Section Geometries Used For Crack Growth .

Figure 2: Critical Paths Used for Linearized and Through-wall Mapped Stresses 22 Fignre 3: Time Histories of Linearized M B Stress at the Inside Surfaee for Heatup Transient for DMWI 23 Figure Through-wall Residual Stress Distribution at 70°F and Curve Fits 24 Figure 5: K-vs-a Plots for Paths 1 and 2 at DMWI for Residual Stresses at 70°F 26 Figure 6: Circumferential Flaw Model, Under Arbitrary Through-Wall Stress Distribution 29 Figure 7: Circumferential Flaw Model, Moment Loading 30 Figure 8: Axial Flaw Model, Under Arbitrary Through-Wall Stress Distribution 31 Figure 9: K-vs-a at Normal Steady State Operating Conditions for DMWI 32 Figure 10: K-vs-a at Normal Steady State Operating Conditions for Safe End 33 Figure 11: K-vs-a at Normal Steady State Operating Conditions for DMW2 34 File No.: 1000320.316 Page 3 of34 Revision: 0 F0306-01RI

1.0 PURPOSE The purpose of this calculation package is to calculate crack growth in the susceptible material regions (SMRs), which include the nozzle-to-safe end weld (DMW 1), safe end (SE) and safe end-to-pipe weld (DMW2), for the spray nozzle with weld overlay (WOL) repair. Loads considered are external piping loads, internal pressure, thermal transients, local effect of stratification, and residual stress. Both fatigue crack growth (FCG) and Primary Water Stress Corrosion Cracking (PWSCC) are considered. PWSCC is due to sustained loading at steady state normal operating conditions.

2.0 METHODOLOGY Representative fracture mechanics models (Section 3.2) are used to determine stress intensity factors (K) within the susceptible material regions. The stress intensity factors for each type of load are computed as a function of assumed crack depth in the susceptible and superimposed for the various operating states. Stresses that contribute to fatigue crack growth and PWSCC are compiled from previous calculations (References 4 and 6). These stresses resulted from primary loads such as internal pressure and external piping loads, and secondary loads such thermal gradient stresses (due to thermal transient events), and residual stresses. The through-wall stresses are extracted and curve fitted to a third order polynomial. FCG (or combined FCG and PWSCC growth ifPWSCC is active) is computed using linear elastic fracture mechanics (LEFM) techniques. Potential for PWSCC is determined by computing the stress intensity factor versus flaw depth curve (K-vs-a) at steady state normal operating conditions. A crack growth law for Alloy 600 weld metals (Alloy 82/Alloy 182) with a multiplier to account for crack growth in a pressurized water reactor (PWR) environment is used at the susceptible material regions.

The time it takes for an initial flaw of 75% of the original base metal thickness to reach the overlay is reported.

3.0 DESIGN INPUTS 3.1 Geometry Details of the spray nozzle geometry are provided in the finite element model calculation package [1].

The sections evaluated for crack growth, and considered representative for the susceptible material regions are shown in Figure 1. The three sections considered here are named DMW1, safe end and DMW2 as shown in Figure 1. Note that in the figure, the minimum and maximum dimensions of the sections (due to the use of minimum and maximum overlay designs in the finite element models [1]) are shown. The thinner (minimum weld overlay) section is used for pressure, piping interface loads and residual stresses. The thicker (maximum weld overlay) section is used for all local thermal gradient stresses. See Section 3.3 for descriptions of the different loads applied to the nozzle.

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3.2 Fracture Mechanics Models Three fracture tuc<;llatltcs moclcls, EPRI LJU',U'" Fracture Handbook [2] and one model for moment loadlrlg Ric."",,,,,, 6 through 8. The three models are described below:

mnQ.!~!!u;!1Q!;!£!t. full 360 0 flaw a cylinder with the actual R/t ratio is used, where I :::; R/t

10. The R/t ratios are calculated as follows dimensions obtained from Figure 1):

Minimum WOL (DMW1): (1.8125)/(1.05) 1.73 Maximum WOL (DMW1): (1.8125)/(1.3) 1.40 Minimum WOL (Safe End): R/t= (1.7638)/(1.0986) 1.61 o Maximum WOL (Safe End): R/t= (1.7638)/(1.3486) 1.31 Minimum WOL (DMW2): (1.846)/(0.7715) 2.39 o Maximum WOL (DMW2): R/t= (1.846)/(1.1 1309) 1.66 It is determined in Reference 12 that the influence coefficients for calculating K must he computed using the plots provided in the reference as opposed to using the provided equation.

• Circumferential flaw, moment loading (Tada-Paris-Irwin model): full 360 0 flaw in a cylinder with the actual inside radius-to-outside radius ratio R/Ro is used. The R/Ro ratios are calculated as follows (geometric dimensions obtained from Figure 1):

o Minimum WOL (DMW1): R/Ro (1.8125)/(2.8625) 0.633 o Maximum WOL (DMWl): R/Ro (1.8125)/(3.1125) 0.58 o Minimum WOL (Safe End): R/Ro (1.7638)/(2.8624) = 0.616 o Maximum WOL (Safe End): R/Ro (1.7638)/(3.1124) 0.57 o Minimum WOL (DMW2): R/Ro (1.846)/(2.6175) 0.71 o Maximum WOL (DMW2): R/Ro (1.846)/(2.9601) 0.62

• Axial flaw under arbitrary through-wall stress distribution (EPRI Ductile Fracture Handbook model): semi-elliptical inside surface flaw with an aspect (depth-to-Iength) ratio of 0.2 in a cylinder is used for FCG and 0.5 for PWSCc. The reason a 0.5 aspect ratio was used for PWSCC is that, assuming stresses are compressive at normal operating pressure and temperature (for which PWSCC is evaluated, see Section 5.0 under "PWSCC"), a higher aspect ratio for semi-elliptical flaws would produce more conservative (less compressive) values ofK. Note that for tensile K distributions, it is the opposite: a lower aspect ratio produces more conservative (greater tensile) values for K. The same R/t ratios computed for the circumferential flaw non-moment loading are used, where I :::; R/t ::: 10. Self-similar growth is assumed. The K at the deepest point is used.

In addition, the same axial flaw model but with internal pressure loading is used for the pressure loading.

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Stress intensity factors using the above flaw models for all loads (Section 3.3) are all calculated via a Microsoft Excel Visual Basic for Applications (VBA) macro in spreadsheet "StresslntensityFactors l.xls." for DMWI, "StressIntensityFactorsSE.xls." for safe end, and "StresslntensityFactors2.xls." for DMW2. The macro is verified to ensure the equations from the EPRI Ductile Fracture Handbook [2] and the Paris-Tada-Irwin solution [3] are correctly used.

3.3 Loads Loads considered (described in detail in the following sub-sections) are internal pressure, interface piping loads, local thermal gradient stresses due to thermal transients, local etTect of stratification, and residual stresses. Through-wall stresses are curve fitted with a third order polynomial in the form shown below:

(1) where:

(J stress (axial or hoop) x distance from the inside surface Stresses were obtained for the above curve fits for through-wall paths in the susceptible material regions (Figure 2). These paths are defined in the thermal and mechanical stress analysis calculation package

[4]. It is important to note that the critical paths selected within the SMR have different path lengths in addition to the fact that mechanical stresses and thernlal stresses are extracted from models with the minimum and maximum overlays, respectively (Section 3.1). The intent for this approach is to characterize the stresses for each of the different loading conditions for three regions within the SMR.

As long as the correct stresses are extracted from the correct model (minimum or maximum overlay model depending on the load type) and correct path, the resulting stress intensity factors are correct because crack growth is computed for only one representative seetion (Figure I) using the most limiting path per each section (DMW1, safe end and DMW2 sections).

Curve fits for residual stresses are calculated in spreadsheets" I000320-3l4.xls" obtained from the residual stress analysis [6]. Linearized stresses and curve fits for local thermal gradient stresses (and stresses due to mechanical loads) are in spreadsheet "CG_SPRAY_DMWl.xls" for DMW1, "CG_ SPRAY_DMWSE.xls" for safe end, and "CG_ SPRAY_DMW2.xls" for DMW2. The curve fit coefficients are used to calculate the K-vs-a values using a K-solution from Paris-Tada-Irwin [3] for circumferential Haw moment loading (Figure 7) or input into the EPRI Ductile Fracture Handbook models [2] for the circumferential (non-moment loading) and axial flaws (Figures 6 and 8). These spreadsheets are included in the computer files. Table I shows the polynomial coefficients for all loads at the DMWI, safe end and DMW2 susceptible material regions. Details of how the Table I coefficients are obtained are explained in the following sub-sections.

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Internal Pressure In the thermal and mechanical stress analysis calculation package [4], a unit internal pressure load (1000 psig) was applied to the spray nozzle finite element model. The resulting through-wall stress distributions within the SMR were extracted and given a third order polynomial curve fit. Files "TMI~PR~P* -,MAP.OUT" (* = I through 18), listed in Appendix A in the table labeled "Stress Output,"

contain the through-wall stress maps. These output files are obtained from [4]. The CSV file versions of these files contain the resulting stress coefficients for the unit pressure loading condition. The resulting stress coefficients are shown in Table I for DMWl, safe end and DMW2 regions. Stress intensity factors resulting from the stresses in Table I are scaled by actnal pressure values during the transient.

Note that a constant through-wall stress of 1000 psig is added to the Co coefficient ofthe "Unit Pressure" stresses to account for the internal pressure acting on the crack face.

Piping Interface Loads The piping interface loads were determined in the design loads calculation package [5] and shown in Table 2. Finite element analyses for an applied unit axial force (1000 lbs) and unit moment (1000 in-Ib) at the free end of the piping interface were performed in [4]. Through-wall stresses within the SMR were extracted from the analysis in [4] and included in this calculation package. See files "TMI_AXIAL_P* MAP.OUT (*= I through 18), and "TMI_MOM$~P* _MAP.OUT" ($ X and Z;

• I through 18) in Appendix A in the table labeled "Stress Output." These files are included in the computer files. For the axial stresses due to the unit moment load, only the maximum stress in the section is used because a constant stress is required in the fractnre mechanics model that is used for moment loading (see Figure 7). Stress intensity factors using the fractnre mechanics models described in Section 3.2 were determined from these stresses. Thermal piping interface loads are scaled to the actnal temperatnres during the transient. The resulting stress intensity factors are scaled in two ways:

(1) by the axial force values (Fx) in Table 2 for the force Ks, (2) by the square root sum of the squares (SRSS) of the transverse moment values (Mx and Mz) in Table 2 for the moment Ks.

The resulting stress coefficients for the unit axial and unit moment cases are shown in Table 1.

Local Thermal Gradient Stresses Bounding thermal transients for crack growth analysis are shown in Table 3, which were developed in the design loads calculation package [5]. Determining the thermal gradient stresses due to these transients is a two-step process. First, a bounding path in the SMR is selected by comparing the thermal gradient stresses at the six representative paths in the SMR (Paths 1, 2, 7, 8, 13 and 14 for DMW 1; Paths 3, 4,9,10,15 and 16 for safe end; paths 5, 6, 11,12,17, and 18 for DMW2, in Figure 2) for the transients. Second, the times at which the maximum and minimum inside surface membrane-plus-bending axial and hoop stresses during each transient listed in Table 3 are determined so that the worst through-wall thermal gradient stresses may be extracted for the bounding path.

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Thennal gradient stresses for Hydro Test will be similar to the thennal gradient stresses for Transient IA at 0 seconds. Thennal gradient stresses for Stratification will be similar to the thennal gradient stresses for Transient 1Bat 0 seconds, from Reference 5, it can be seen that stratification occurs at 555°F and which are the same conditions as Transient 1B at 0 seconds.

To select the bounding path in the SMR, the time histories of the inside surface linearized membrane-plus-bending stresses for each transient are plotted for axial and hoop stresses. A typical plot is shown in Figure 3 for axial and hoop stresses for the Heatup transient The linearized membrane-plus-bending stresses and corresponding mapped stresses fitted with a third order polynomial are contained in files "TMCSTR~$$~P* ~MAP.csv" <<$$ event IDs I through 6 in Table 4 and

• 1 through 18). Figure 3 shows, for example, that Path I is used for the maximum axial stress and Path 2 for the minimum axial stress for the Heatup transient Table 1 shows the bounding paths used at the SMR.

Once the bounding path for the SMR is detennined, the times at which the maximum and minimum membrane-plus-bending axial and hoop stresses during each transient may now be detennined. Note that although the inside surface linearized membrane-plus-bending axial and hoop stresses were used in selecting the critical times, the stress coefficients resulting from the through-wall mapped stresses at the critical times are total stresses. The resulting bounding thennal gradient stress coefficients due to the thennal transients at the times of maximum and minimum inside surface linearized membrane-plus-bending stresses are shown in Table 1.

Residual Stresses Residual axial and hoop stresses were extracted from the residual stress analysis [6] for six paths within the SMR. These paths from [6] are Paths 1 through 6, and correspond (approximately) to Paths 1 through 18 in Figure 2. Stresses were obtained from file 1000320-314.xls" of Reference 6. The If resulting stresses were curve fit with a third order polynomial in this spreadsheet in worksheet "Path Data." The bounding (least compressive or most tensile stress) residual stress curve fits were detennined for two paths of the SMR for DMW1, safe end and DMW2. These bounding residual stresses are used for the crack growth analysis. Residual stresses at 70°F are shown in Figure 4.

The bounding path for the axial and hoop stress for the six paths within the SMR can be detennined either by calculating the K-vs-a curve via the "StresslntensityFactors1.xls",

"StressIntensityFactorsSE.xls", and "StresslntensityFactors2.xls" spreadsheets for DMW1, safe end and DMW2 SMR's respectively. The residual stress coefficients (at 70°F) for all six paths are input into the spreadsheet and the K-vs-a plotted. This comparison is shown in Figure 5, which shows that at 75% of the original wall thickness, Path 2 gives the worst K distribution for the axial flaw and circumferential flaw for DMW1 region; Path 3 gives the worst K distribution for the axial flaw and circumferential flaw for safe end region; Path 5 gives the worst K distribution for the axial flaw and circumferential flaw for DMW2 region. The Ks for these selected paths that give the worst Ks are considered representative for the entire PWSCC susceptible material.

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The resulting residual stress coefficients at 70°F and at steady state pressure (2155 psig) and temperature (650°F) are shown in Table 1.

3.4 Crack Growth Laws 3.4.1 Alloy 82/182 Fatigue Crack Growth Law NUREG/CR-6907 [7] indicates that "The CGRs of Alloy 82/182 in the PWR environment are a factor ~

5 higher than those of Alloy 600 in air under the same loading conditions." Thus the fatigue crack growth rate (FCGR) for the SMR used in this analysis is that for Alloy 600 in air (Equation (2) below) multiplied by 5. The FCGR for air obtained from NUREG/CR-672I [8] is given by:

(daldN)air = C A6 00 (l-0.82Rr22 (,1,K/ 1 , units ofm/cycle (2) where:

C A600 4.835xlO- 14 + 1.622xlO- 16T 1.49 xlO- 18T 2 + 4.355 x10- 21 T 3 T temperature inside pipe, °C (taken as the maximum during the transient)

R R-ratio = (Kmin/Kmax)

,1,K K max - Kmin range of stress intensity factor, Mpa_m05 Note that Equation (2) in accordance with NUREG/CR-6907 is independent of rise time.

3.4.2 Alloy 82/182 PWSCC Growth Law When appropriate PWSCC growth is computed using Equation (3) below. This is for Alloy 182 weld metal and obtained from Eq. 4.5 ofMRP-115 [II, pgs 4-4 to 4-5 for US customary units].

. [Qg [

a=exp-- - J] (

I - -I . aK .)/1 R T Tref (3) where:

a= crack growth rate in/hr Qg 31 kcallmole R I.IOE-03 kcal/mole-oR T 650 F Tref= 1076.67 R (617°F) a 2.47E-07 crack growth amplitude K= Stress intensity factor ksi-i nJ\0.5

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3.4.3 Treatment a/Negative Stress Intensity Factors Because of the beneficial compressive residual stresses produced by the weld overlay, the stress intensity factors for many load cases are negative. This condition is handled as follows in the FCG analyses:

L For Alloy 600 and its weld metals (Alloy 82/Alloy 182), the crack growth law in Equation (2) above includes an R-ratio correction (l-0.82Rr 22 that applies to both positive and negative R-ratios [7, 8]. For negative R-ratios, the factor is less than 1, yielding a corresponding decrease in crack growth rate when Kmin is negative. Therefore, Equation (2) is used directly for crack growth in the SMR when K max is greater than zero, for both positive and negative K min . As noted, a factor of 5 [7] is applied for PWR environmental effects.

2. Ifboth K max and K min in a load cycle are negative, zero fatigue crack growth is assumed. This is a reasonable assumption based on the work on compression fatigue crack growth (FCG for which K max and K min are negative) in Reference 10, in which it was shown that although FCG is present, the values ofFCG are several orders of magnitude smaller than for FCG for which K max and K min are not negative.

4.0 ASSUl\lPTIONS Basic assumptions for the analysis are listed below:

• Through-wall stresses are corve fitted with a third order polynomiaL The use of the third order polynomial is limited by the crack modeL In most cases the curve fit is more conservative than the actual plot (e.g., axial stress for Path 2 in Figure 4). The residual stresses provide enough compression in the steady state operating condition that any slight changes in the curve fit will have negligible effect on the overall crack growth.

• See Section 3.2 for assumed flaw models.

• See Section 3.4 for the crack growth laws and related assumptions.

• For FCG, there is no requirement to bound the load pairing between transients per the ASME Code Section XI [9]. Each thermal transient that was analyzed in the thermal and mechanical stress analysis calculation package [4] is analyzed sequentially per Table 4, and the cumulative effect of all transients is summed. In addition, incremental growth due to PWSCC (if active) is computed and added to incremental growth due to FCG. This approach is consistent with the ASME Code Section XI, C-3200.

5.0 CALCULATIONS A Microsoft Excel VBA routine is implemented to calculate combined FCG and PWSCC growth. This combined FCG and PWSCC calculation is based on a yearly basis: one year ofPWSCC followed by one year ofFCG. Note that for more accuracy, crack growth may also be based on a monthly basis.

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Combined FCG and PWSCC growth is calculated until the flaw depth reaches the interface of the base metal and overlay. The number of cycles and sequence of events for FCG are shown in Table 4 and are given for 60 years of operation. The PWSCC portion of the crack growth is only active if, during the growth of the flaw, the stress intensity factor at steady state normal operating conditions (NOC) is a positive value, which is in accordance with MRP-115 [II, page 4-2], which states that "there is insufficient data to justify a stress intensity factor threshold other than zero." The Ks from spreadsheet "StresslntensityFactors I.xls" are imported into spreadsheet "CG~SPRAY~DMWI.xls" to compute crack growth for DMWI; similarly, "StresslntensityFactorsSE.xls" are imported into spreadsheet "CG~SPRAY~DMWSE.xls" and "StresslntensityFactors2.xls" are imported into spreadsheet "CG SPRAY DMW2.xls".

For FCG, the individual terms that constitute nominal K max and K min for the ealculation of ilK in Equations (2) and (4) are summarized in the tabulations below. The individual Ks for nominal K max are combined (summed) with all appropriate scale factors applied. Similarly, the individual Ks for nominal Kmin are combined (summed) with all appropriate scale factors applied. ilK is computed by taking the difference of the resulting summed K max and K min . Note that Kresidual and ~eadweight are constant loads.

Kresidual Kresidual Kpressure Kpressure

~ead weight ~ead weight Kthermal piping load*state 1 Kthermal piping load*state 2 Kthermal state I Kthermal state 2 Kstratitication state I Kstratification state 2 Thermal K max and K min values of Transient lA at time 0 seconds are used for Hydrostatic Test. Similarly thermal K max and Kmin values of Transient IB at time 0 seconds are used for Stratification event.

OBE is treated as a separate event that combines with the highest K value for the K max side and the lowest K value for the Kmin side for 660 cycles [5].

PWSCC The K-vs-a curves for both circumferential and axial flaws at normal operating conditions (NOC) are calculated in the PWSCC worksheets of spreadsheet "CG_SPRAY_DMW1.xls" and shown in Figure 9 for DMWI; "CG_SPRAY DMWSE.xls (Figure 10) for safe end; "CG_SPRAY DMW2.xls" (Figure

11) for DMW2. The stresses at steady state NOC include internal pressure stresses (P 2155 psig [5]),

residual stresses (at 70°F), steady state thermal stresses at 650°F, and stresses due to the non-cyclic piping loads (including deadweight). Note that an additional 2155 psig [5] was added to the Co coefficients for the stresses labeled "Resid+press+temp" in Table I to account for crack face pressure.

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This crack pressure is only applicable for the PWSCC evaluation. For both axial and circumferential flaws, the models described in Section were used to calculate the stress intensity factors, but with an aspect ratio of 0.5 for the axial flaw.

PWSCC is a time dependent phenomenon and occurs at a sustained loading condition. Given that the great majority of plant operation is at nonnal steady state operating conditions (NOC), PWSCC is defined by stress conditions at NOC. The 75% initial crack is grown using the cyclical loadings described previously. At steps in the evaluation, the value of the crack tip K is continuously checked. If the K is less than zero no PWSCC is assumed for the next step of fatigue crack growth. If the crack tip K becomes greater than zero, then the PWSCC crack growth rate in Section 3.4.2 is used to calculate an incremental PWSCC crack growth, which is added to the total crack growth.

In this case the Ks at steady state NOC were negative for all flaw depths for both the circumferential and axial flaws except for axial flaw at safe end. Thus, PWSCC is not active for all flaw depths between 75% and 100% of the original base metal except for axial flaw at safe end.

6.0 RESULTS AND CONCLUSIONS The crack growth results are shown in Table 5. At the susceptible material region, it takes greater than 60 years for an initial flaw of75% of the original base metal thickness at the analyzed section to reach the overlay for the circumferential flaw and the axial flaw.

It was also shown that the Ks at steady state NOC were negative for all flaw depths for both the circumferential and axial flaws except for axial flaw at Safe End. Thus, PWSCC is not active for all flaw depths between 75% and 100% of the original base metal except for axial flaw at Safe End.

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7.0 REFERENCES

I. SI Calculation No. 1000320.312, "Material Properties and Finite Element Models for Pressurizer Spray Nozzle with Weld Overlay Repair," (for revision number refer to SI Project Revision Log, latest revision).

2. Zahoor, A., EPRI Report No. NP-630 1-0, Ductile Fracture Handbook, Volumes 1, 2, and 3, (N 14-1), Research Project 1757-69, June 1989.

3. Hiroshi Tada, Paul C. Paris, and George R. Irwin, "The Stress Analysis of Cracks Handbook," Third Edition, ASME Press, 2000.

4. SI Calculation No.1 000320.313, "Pressurizer Spray Nozzle Thermal and Mechanical Stress Analysis Calculation," (for revision number refer to SI Project Revision Log, latest revision).

5. SI Calculation No. 1000320.311, "Design Loads for Pressurizer Spray Nozzle with Weld Overlay Repair," (for revision number refer to SI Project Revision Log, latest revision).

6. SI Calculation No.1 000320.314, "Residual Stress Analysis of Pressurizer Spray Nozzle with Weld Overlay Repair," (for revision number refer to SI Project Revision Log, latest revision).

7. NUREG/CR-6907, "Crack Growth Rates of Nickel Alloy Welds in a PWR Environment,"

U.S. Nuclear Regulatory Commission (Argonne National Laboratory), May 2006.

8. NUREG/CR-6721 , "Effects of Alloy Chemistry, Cold Work, and Water Chemistry on Corrosion Fatigue and Stress Corrosion Cracking of Nickel Alloys and Welds," U.S. Nuclear Regulatory Commission (Argonne National Laboratory), April 2001.

9. ASME Boiler and Pressure Vessel Code,Section XI, Rules for Inservice Inspection of Nuclear Power Plant Components, 2004 Edition.

10. Lenets, Y. N., "Compression Fatigue Crack Growth Behavior of Metallic Alloys: Effect of Environment," Engineering Fracture Mechanics, Vol. 52, No.5, 1997.

11. Materials Reliability Program Report MRP-115, "Materials Reliability Program: Crack Growth Rates for Evaluating Primary Water Stress Corrosion Cracking (PWSCC) of Alloy 82, 182, and 132 Welds," September 2004.

12. SI Calculation 0800554.301, "Computation ofInfluence Coefficients for Two Stress Intensity Factor Solutions From EPRI Ductile Fracture Handbook," Rev. O.

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Table 1: Stress Coefficients for Various Loadings at tbe DMWI Transient/Load Time Path Axial Stress Coefficients, psi Time Path Hoop Stress Coefficients, psi Name (sec) Used Co C1 I Cz C3 (sec) Used Co C1 C2 C3 Unit Pressure ( ) ~- 2 1661.9 I ~67 1500A ~286.1 ~~

7 3341.9 ~1912.3 1470.5 ~557,3 Unit Axial ~~

2 66,7 I 7.6 I ~ 122 5,2 ~~

7 ~0.3 2.1 -2.1 -0.2 Unit Moment -~ 2 69 8 1.1 ~.

0.3 ~0.2 Resid70, Path 1 ~~ 1 7465 ~322363 548734.7 ~200332 I ~67148 *228622 719933 ~365302 Resid70, Path 2 ~~ 2 ~1861 ~295011 657991 ~304460

~ 2 *65863 *378679 1327293 -812842 Resid+press+temp, Path 1 (2l 1 I

-18011 -92847 199113 *57561 *57281 ~145167 552343 ~293078 Resid+press+temp, ~~

2 2 Path 2 t1l ~12593.2 -175063 461698.7 *222685 -51924.3 -319382 1159462 ~ 723909 10956 1 -4687.3 65569.9 ~ I04067 40865.5 6794.9 -41470.2

~81 Transient 1HIGH 10908 8 22438A Transient ILOW 360 2 ~534.6 2500.2 ~3219.2 1271.2 2 )6 2169.1 *2812.2 1121.4 t2HIGH 21681 I ~1343.9 44385.2 ~75139.5 29820A 2 46.7 ~227.6 -27767.3 16212.2 Transient2LOW 0 2 -11134.5 83554.2 -132163 56576.1 I 313.8 7906.6 -15394.3 6399.7 Transient3HIGH 308A 13 -19tl20.6 136265.1 -193031 74191.1 316.8 7 -1263.1 52219.8 ~88485.8 35445.1 Transient3LOW 51.6 2 *14590.1 92995.3 -139203 58413 51.6 14 -2363.4 45917 -81903.7 36658.2 Transient4HIGH 99 13 -2t1370 139017.8 ~ 194329 74416.1 99 7 ~2863.1 56302.1 -91277.3 36191.1 Transient4LOW 773.4 2 -19096.1 105262A -148215 60734.5 773.4 14 -6886.5 60933 -94304.2 40207.8 Transient5HIGH 692 13 -20100.9 138753.1 -194619 74624.2 692 7 -2639.6 55655A -90812.1 36091.1 Transient5LOW 22 2 -13457.3 91128.6 ~139137 58722.9 22 14 I -139Ll 43795.9 -81671.6 36961.9 Transient6HIGH 900 13 -17173.7 126330 -180092 69353.5 915 8 I 3807.5 27050.5 -68106.9 33221.1 Transient6LOW 0 2 -11478 86207.3 -136460 58464.1 0 14 I 973 36027.9 -75495.6 35363.2 Notes:

(I) 1000 psig was added to the Co coefficient tenn of the Unit Pressure case to account for crack face pressure.

(2) 2l55psig was conservatively added to the Co coeflicient term of the residual stresses and nonnal operating pressure and temperature cases to account for crack face pressure. This is needed for calculating stress intensity factors at nonnal steady state operating conditions.

File No.: 1000320.316 Page 1401'34 Revision: 0 F0306-01RI

Table 1: Stress Coefficients for Various Loadings at the Safe End (continued)

Transient! Axial Stress Coefficients, psi Hoop Stress Coefficients, si Time Path Time Path Load (sec) Used Co C, C1 C3 (sec) Used Co C, C1 C3 Name Unit Pressure ( J ~- 3 1815,7 -656,9 980,1 A94,2 -- 4 3357 -1791.2 722,1 -117,8 Unit Axial -- 15 68.4 ~4,5 7, I -3.6 -- 16 47.7 -88.3 45.1 -14.2 Unit Moment ~.

10 71 16 1.1 -- 0.3 64.9 Resid70. Path 3 ~~

3 -20214 I *232640 624324 -317131 .. 3 -66326 21866 454429 -344296 Resid70. Path 4 -~ 4 -19821 *208347 559612 -294667 -- 4 -51795 -305618 909433 A79645 Resid+press+temp, Path 3 ,,) 3 8952,2 *207804 538416 -276023 *49674.4 33872.59 363978.6 *289718 Resid+press+temp, 4

-152053 409736.6 -214737

-- 4 Path 4 '" 12181.3 *32756.2 *261646 753794.8 -392351 rransientl HIGH 10932 3 I 10594.2 A386.8 -27209.4 16065,7 10932 16 13909.9 -14666 -15256,8 12242.7 Transientl LOW 360 16 A83.3 1344,6 *1169.3 401.5 360 16 *584,8 1926.6 *1865 655,7 Transient2HIGH 21654 16 12179.8 -9985,9 -19986.3 13406.9 21654 16 17218,4 -28839.8 404.4 6670.3 Transient2LOW 39600 4 938.4 3301.9 -9813.3 4835.5 39600 4 1639.1 1157,6 -7631.8 4167,2 Transient3HIGH 299.8 15 8189.6 5471.6 -38123.8 20063.8 286.5 10 13687,7 -9626.6 *21915.1 14826.6

=

Transient3 LO W 3l.8 16 A149.4 268088 -38059 15537,5 31.8 16 -3511.5 31334.3 A6023.7 19138.3 Transient4HIGH 99 15 5257,6 12655 -42512.8 21052.7 99 10 9661.5 -587.5 -26310.4 15285.5 Transient4LOW 753.4 16 11382,9 41376.2 -43413.1 15482.9 753.4 16 -12818.8 55509.2 -62616.4 23008.8 Transient5B1GH 712 15 5711.5 12030,5 -42473.7 21022.9 692 10 9633.6 -401.9 -26534.7 15349.2 TransientS LOW 19 10 -2065,5 27765.9 -47520.9 20240.8 20 16 149.9 26957.4 -48768 21438.3 Transient6HIGH 900 15 8475,2 5921,8 -40356.8 21325.4 900 10 12617.7 -6175.7 -26074.6 16461 Transient6LOW 0 10 4410.4 10815.8 -36157.4 17955.3 0 3 4456.3 15949.2 -48490.3 24342.3 Notes:

(I) 1000 psig was added to the Co coefficient term of the Unit Pressure case to account for crack face pressure.

(2) 2 I 55psig was conservatively added to the Co coefficient term of the residual stresses and normal operating pressure and temperature cases to account for crack face pressure. 111is is needed for calculating stress intensity factors at normal steady state operating conditions.

File No.: 1000320.316 Page 15 of34 Revision: 0 F0306-0IRI

Table 1: Stress Coefficients for Various Loadings at the DMW2 (continued)

Transient/ Axial Stress Coefficients, Dsi Hoop Stress Coefficients, si Time Path Time Path Load (sec) Used Co C, C, C, (sec) Used Co C, C, C, Name Unit Pressure I -, 5 1414.7 -316.1 -1777 1511.8 .. 18 3739.4 -1153.5 1113.9 -400.9 Unit Axial .. 6 96 14.4 -107.6 103.7 -- 18 82.4 -164.8 93.5 -40.2 Unit Moment .. 11 95 18 1.1 -- 0.3 85.6 Resid70, Path 5 .. 5 -9408.89 -342880 1142648 -802714 5 -50760.7 -504408 2185410 -1719147 Rcsid70, Path 6 .. 6 6829818 -461574 1115647 -75171 1 .. 6 -48219.1 ,416755 1560161 -1064928 Path 5 tij

.. 5

,17921.2

.. 5

-160571 689345.1 -513050 -32677.3 -405929 180'947 -1442403

.. 6 .. 6 Path 6 <'I -907146 -271521 848615.8 -545389 -31883.7 -290259 1194809 -816460 Transientl HIGH 10908 5 532.8 37006.3 -95131.6 54415.1 10884 12 9500.6 6657.8 -46437.5 18134.2 Transient 1LOW 28092 6 -16624.3 80523.4 -85666 2610L5 360 6 -377.9 16936 -2109 1147.3 TransicntlHIGH 21654 5 7208.5 4916.6 -52111.9 35713.8 11654 12 13225.1 -14969.4 -19240.2 17075.7 Transicnt21.0W 0 6 -15574.7 78097.9 -84573.7 26002.9 39600 6 1377 -105.8 -75473 6160 Transient3HlGH 282 5 -7508.8 81611.5 -159843.1 82195.8 282 12 8305.9 29280.1 -78884.1 41625.7 Transicnt3LOW 18.6 6 -15193 107206.1 -1056073 30333.5 18.6 6 -3084.1 8844.3 -2051.7 4800.2 rransient4HIG11 64.5 5 -10314.3 88758.5 -162540,1 81514.5 64.5 12 5118.9 37375.3 -82733.5 41781.5 Transient4LOW 743.4 6 -31923.5 119912.5 -123085.4 35037.3 743.4 6 -10435.1 31226.1 -17714.2 8606.7 Transient5HIGH 671 5 -10516.9 90022 -164417.3 81304.6 652 11 4883.7 38531.1 -84251 42366.4 Transicnt5LOW 18 6 -21172 1008071 -105731.4 31555.4 19 6 5216 -533.1 2008.5 4964.7 Transicnt6HIGH 900 5 -6493.5 76835.8 -154357 80489.8 900 11 7576 28710.8 -77505.3 41178.8 Transient6LOW 0 6 -15834.2 80348.1 -88159.6 17599.1 0 6 7433.8 -13141.1 19870.1 760.1 Notes:

(I) 1000 psig was added to tbe Co coefficienI term of the Unit Pressure case to account for crack face pressure.

(2) 2155psig was conservatively added to tbe Co coefficient term of the residual stresses and normal operating pressure and temperature cases to account for crack face pressure, This is needed for calculating stress intensity factors at normal steady state operating conditions.

File No,: 1000320,316 Page 16 of34 Revision: 0 F0306-01RI

Table 2: Spray Nozzle Piping Interface Loads Forces, Ib Moments, Ib-in Load Case Fx Fv Fz Mx Mv Mz ei ht 3 14 13 252 1222 781 629 493 10932 20102 15723 6 4 11 6468 704 348 imum Stratification Moment Ran e 173,263 Ib-in()

Notes: I) Transformed forces and moments are listed on an absolute basis.

2) Fy is oriented in the axial direction of the nozzle.

3) The number of cycles is 7200 per [5].

File No.: 1000320.316 Page 17 of34 Revision: 0 F0306-01RI

Table 3: Bounding Tbermal Transients for the Spray Nozzle 10800 370 370 173 1378 159 1.000 10800 370 70 22 21 216 1.000 12240 410 110 26 26 319 0.082 12240 410 110 141 427 319 0.082 27432 650 532 160 739 2242 0.953 27432 650 532 183 2149 2242 0.953 I

10800 430 300 177 1540 216 0.474 21600 400 280 177 1484 216 0.433 21600 400 210 152 578 159 0.289 25920 330 150 147 494 15 0.165 (bounds 0 650 566 182 2111 2242 1.023 2A, 3,4, 4 682 586 182 2088 2792 1.064 6,7,8C, 12 682 590 182 2084 2792 1.072 10,11, 144 682 590 182 2084 2792 1.072 14, and 144 639 575 182 2101 2092 1.041 20B) 192 633 562 182 2116 1997 1.014 282 625 546 183 2134 1897 0.981 I

8 662 570 174 1261 2442 1.031 12 662 570 174 1261 2442 1.031 20 650 555 164 844 2242 1.000 30 632 550 154 612 1992 0.990 720 643 550 154 612 1992 0.990 720 643 636 173 1405 2142 1.167 (also 1 651 544 183 2136 2260 0.977 bounds 4 654 562 182 2116 2312 1.014 17A) 6 657 564 182 2113 2348 1.019 10 661 560 182 2118 2419 1.010 15 666 576 182 2100 2507 1.043 22 653 568 182 2109 2297 1.027 22 653 568 41 44 2297 1.027 600 616 558 41 44 1757 1.006 (also 900 519 500 182 2093 842 0.887 182 Note: The above table is reproduced from File No.: Page 18 of34 Revision: 0 F0306-01 Rl

Table 4: Sequence of Events and Cycles for Fatigue Crack Growth No. of Events Event ID 60-yr Cycles peak/valley pairs Transient 1 (Heatnp) 1 240 1 Transient 2 (Cooldown) 2 240 1 Transient 3 (Step load reduction) 3 65938 1 Transient 4 (Reactor Trip with Loss of 4 1 Flow/Loss of Station Power) 80 Transient 5 ( Reactor Trip Due to High Reactor 5 1 Temperatnre) 180 Transient 6 (Rapid Depressurization) 6 1480 1 Transient 7 (Hydro test) 7 20 1 Stratification Moment 8 7200 1 OBE 9 660 1 File No.: 1000320.316 Page 19 of34 Revision: 0 F0306-01RI

Table 5: Crack Growth Results Flaw (I) Flaw Depth By Fatigue and Time to Reach Overlay P\VSCC at 60 years Circumferential Flaw at DMW 1 >60 0.5625" Axial Flaw at DMWI >60 0.5625" I Circumferential Flaw at Safe End >60 0.5367" Axial Flaw at Safe End >60 0.5367" Circumferential Flaw at DMW2 >60 0.3124" Axial Flaw at DMW2 >60 0.3124" Notes:

(1) Initial flaw depth 75% of original base metal thickness at the section analyzed 0.563" for DMW1; 0.537" for Safe End; 0.313" for DMW2.

File No.: 1000320.316 Page 20 of34 Revision: 0 F0306-01 Rl

SAFE END DMW1 WELD REGION Thk." 0.383" (Min)

Thk. - 0.633" (Max)

DMW2 WELD REGION Ro- 2.2625" Ro"2.56" RI-1.846" Ri=1.76" RI=1.81" TMI-l Spray Nozzle Residual Geanetry Figure 1: FEM Section Geometries Used For Crack Growth File No.: 1000320.316 Page 21 of34 Revision: 0 F0306-01 RI

Note: Only the model with minimum overlay dimensions is shown. The path numbering for the model with maximum overlay dimensions is similar.

Figure 2: Critical Paths Used for Linearized and Through-wall Mapped Stresses File No.: 1000320.316 Page 22 of34 Revision: 0 F0306-01 RI

Axial Stress

-P,lth 1 - Path 2 --- Path 7 - Path 8 "'.._~ Path 13 - Path 14 soO(:

7000 6000

---

*-~~--1\~-~-.----:-.-----------

Q.

5000 4000 r~

. . . . . . . . . . . . . . **********tl VI

~'" 3000 II \.'

\I'l  ! ~.

2000 1000 o I.

o 5000 10000 15000 20000 25000 30000 35000 i

-1000 TIme (sec)

Hoop Stress

-P:lth1 --P"th? --P.,th7 --'ilthR --P"th1i -P;Jth1-1 10000 8000 6000

~

'"VI 4000

~

Vl 2000 0

5000 10000 15WO 20000 2500:> 3000(* 35000

-2000 . ....---- " -

TIme (sec)

Figure 3: Time Histories of Linearized M + B Stress at the Inside Surface for Heatup Transient for DMWI File No.: 1000320.316 Page 23 of34 Revision: 0 F0306-01 R I

Path 1 Through-Wall Residual Stress Path 2 Through-Wall Residual Stress Y -304460x 3 + 657991 x 2 - 295011 x - 1861.4 40 ' ; ,Hoop 70"F Poly, (AxiaI70'F) 20 (Hoop 70"F)

'(i)

. 0 0 til VI VI

!20 li)

20 li)

-40 - . - Axial 70' F

-40

~Hoop70"F

-60

-60 Poly. (AxiaI70'F)

-80

~~ Poly. (Hoop 70' F)

-80

-100

-100 -'120 0 0.5 1 0 0.2 0.4 0.6 0.8 Di stance from ID Surface (in) Distance from 10 Surface (in)

Path 4 Through-Wall Residual Stress I 00 .,---'-~-'~-~-----~-........,

y= -2SI4667)(3+ 5596'12x 2 - 208347x-80 60 40

' ijj

'(i; 20 V) a +--------+,I--+~(J_------_i 0 II) til II)

Q,) III u?'20 +-~~--_7f-+7"--1 tiS -20 Hoop 70'F -40 Poly. (Axial 70 F)

-60 Poly. (Hoop 70"F)

-60 ~~El;"f'-_*~ '========~I

-80 y = -317131 x 3 + 624324x L 232640x - 20214

-100 o 0.2 0.4 0.6 0.8 0 0.2 0.4 0.6 0.8 1 1.2 Distance from 10 Surface Distance from 10 Surface (in)

Figure 4: Through-wall Residual Stress Distribution at 70°F and Curve Fits Note: jt1inimum overlay wall thickness is used to calculate residual stresses.

File No.: 1000320.316 Page 24 of34 Revision: 0 F0306*01 RI

Path 5 Through-Wall Residual Stress Path GThrough-Wall Residual Stress 80 ,-~~~~~~~~-~---~

802714)(3+ 1E+06x 2 342880x 9408.9

-.20

'~w 'Vi j(,

0 ~~~---~--ir.......-,f----------i

'0 til III til Q) ~20 +-~-----Hf-----------i

<it 20 Iii

• 40 :J Hoop70'F

-+-AxiaI70"F Poly.(AxiaI70'F)

• 60 s Hoop70'F Poly. (Hoop 70' F)

-80 Poly, (AxioI70"F) -80 +-~~:""""-_---'========r-'

Poly. (Hoop 70~ F) y = -751711x 3 + 1E+06x 2 - 462574>: + 6829.8 I

-100 -100 +------,----,...---,----r-------j 0 0.2 0.4 0.6 0.8 o 0.2 0.4 0.6 0.8 OistancefromlD Surface Figure 4: Through-wall Residual Stress Distribution at 70°F and Curve Fits (continued)

Note: Minimum overlay wall thickness is used to calculate residual stresses.

File No.: 1000320.316 Page 25 of34 Revision: 0 F0306-01Rl

Circumferential Flaw Axial Flaw Resid70, Path 1 - Resld70, Path 2 0

0.5 0.6 0.7 o.g 0.9 lfl *20 °

?c

'iii

• ,10 u

Q (ll

• 60

'Vi c

C1I *80 E

~

<lJ iii *100

• 1)0 Flaw Depth (in)

Figure 5: K-vs-a Plots for Paths 1 and 2 at DMWI for Residual Stresses at 70°F Note: Minimum overlay wall thickness is used to calculate residual stresses, File No.: 1000320.316 Page 26 of34 Revision: 0 F0306*01Rl

Circumferential Flaw

'-'~-f\'e5icI70, Jath -Resid70, Path 4 0

0.2 0,3 0.4 os 06 0.7 0.8 09 1 10 lJ'\

~

,5

.0; *20

()

It!

u. *30

'Vi c:

(II

.E .La

...~"'

V'l

• 50

• 60 Flaw Depth (in)

Axial Flaw

---Rcsid70, Path 3 --Rcsid70, Path 4 40 20 lJ'\

0<; 0

,~

'Vi 0.5 0,6 0.8 0.9 1

. *20

()

ti I1l *40 u.

'~

c: *60

~

c (II *80 lJ'\

• 100

• 120 Flaw Depth (in)

Figure 5: K-vs-a Plots for Paths 3 and 4 at Safe End for Residual Stresses at 70°F (cont'd.)

Note: .Minimum overlay wall thickness is used to calculate residual stresses.

File No.: 1000320.316 Page 27 of34 Revision: 0 F0306-01RI

Circumferential Flaw Resld70, Path 5 -Resid70, Path 6 10

~

(IJ

-70 Flaw Depth (in)

V'l Figure 5: K-vs-a Plots for Paths 5 and 6 at DMW2 for Residual Stresses at 70°F (cont'd.)

Note: Minimum overlay wall thickness is used to calculate residual stresses.

File No.: 1000320.316 Page 28 of34 Revision: 0 F0306-01RI

Source: Reference 2 (Zahoor model, 1 :S R/t:S 10)

Figure 6: Circumferential Flaw Model, Under Arbitrary Through-Wall Stress Distribution File No.: 1000320.316 Page 29 34 Revision: 0 F0306-0IR I

Source: Reference 3 is defined below (also from Reference 3)

-I".

,,;:J.t.1

~ --

Figure 7: Circumferential Flaw Model, Moment Loading File No.: 1000320.316 Page 30 of34 Revision: 0 F0306-01 RI

e SInIclllntll""""" Associatf1s. Inc.-

Source: Reference 2 (Zahoor model, 1 S R/t S 10)

Figure 8: Axial Flaw Model, Under Arbitrary Through-Wall Stress Distribution File No.: 1000320.316 Page 31 of34 Revision: 0 F0306-01RI

Circumferential Flaw 40 "1 30 Q

c t/)

20

.lI:

o 10 1;;

tl': 0

~

!-10

'II C

t/) -20

...~

U; -30

-40 Flaw Depth, Inches Axial Flaw o

"1 Q

1: -10

-60 Flaw Depth, Inches Note: Initial flaw depth 75% of original base metal thickness at the analyzed section = 0.5625" Figure 9: K-vs-a at Normal Steady State Operating Conditions for DMWI File No.: 1000320.316 Page 32 of34 Revision: 0 F0306-01RI

Circumferential Flaw o

~ -5 Q

.(

• i -10 o -15

'S

. -20

.f:'

-25 iUl

~

-30 Ul

~

en -35

-40 Flaw Depth, Inches Axial Flaw 30

~

Q 20

.(

c Ul

.:0:

10 0

t) u.. 0 E'Ul

- C Ql c

-10 Ul

~

-20

( /)

-30 Flaw Depth, Inches Note: Initial flaw depth = 75% of original base metal thickness at the analyzed section = 0.5367" Figure 10: K-vs-a at Normal Steady State Operating Conditions for Safe End File No.: 1000320.316 Page 33 of34 Revision: 0 F0306-01RI

Circumferential Flaw I.Il o

-5 c

~ -10

.:0::

<)

ti -15

~

u..

.::- -20 In c -25 In Xl

n -30

-35 Flaw Depth, Inches Axial Flaw

~

o c

ill

.:0::

<)

U

!Ill u..

£' -20 ill

-c<II C

ill Xl U)

Flaw Depth, Inches Note: Initial flaw depth = 75% of original base metal thickness at the analyzed section 0.3124" Figure 11: K-vs-a at Normal Steady State Operating Conditions for DMW2 File No.: 1000320.316 Page 34 34 Revision: 0 F0306-01RI

APPENDIX A COMPUTER FILE DESCRIPTIONS File No.: 1000320.316 Page A-I of A-3 Revision: 0 F0306-01Rl

Stress Output Filename Description Linearized stresses and mapped through-wall stresses for transient $ and path

• TMI STR~$~p* MAP.CSV~

=

• path no. 1 through 18 (Figure 2); $ 1A, 1B, 7 8C 11,8A 15, 8B, and 9 TMI- PR p* MAP.OUT (CSV file Through-wall stresses for unit pressure run for path

• version also included) = =

where

• path no. 1 through 18 (Figure 2)

TMI AXIAL p* MAP.OUT (CSV Through-wall stresses for unit axial load run for path *

~ file version also included) = =

where

• path no. 1 through 18 (Figure 2)

TMI MOM$ p* MAP.OUT (CSV file Through-wall stresses for unit moment load run for path

• version also inclUded) = = =

where

• path no. 1 through 18 (Figure 2); $ X and Z.

File No.: 1000320.316 Page A-2 of A-3 Revision: 0 F0306-01RI

Spreadsheets Filename Description Spreadsheets that contain Microsoft Visual Basic macros for calculating stress intensity StresslntensityFactors$.xls factors at the susceptible material regions DMW1, Safe End and DMW2, where $ =1, SE, and 2.

Spreadsheets that contain extraction of TMI~ Spray~ThennMechStressCoefCDMW$.xlsm linearized stresses; source of Table 1.

Where, $ = 1, SE, and 2 Spreadsheet that contains stress coefficients for thermal transient stresses at TableOfCoefficients.xls the SMR; used as input to "StresslntensityFactors$" spreadsheets, where $ =1, SE, and 2.

Spreadsheet that contains through-wall IOOO320-314.xls residual stresses from Reference 6.

Spreadsheet calculation of crack growth at the susceptible material regions for the CO SPRAY DMW$.xls circumferential and axial flaws at DMW1 ,

Safe End and DMW2, where $ =1, SE, and 2.

File No.: 1000320.316 Page A-3 of A-3 Revision: 0 F0306-01RI

ATTACHMENT 5 Non-Proprietary Version Calculation No. 1000320.310 and Calculation No. 1000320.314

I, No.: 1000320.310

>liT Associates, Inc.1JlJ

1. % if $  % ,ffi uU UCiUlI UI '4 Project No.: 1000320 CALCULATION PACKAGE Quality Program: [81 Nuclear o Commercial PROJECT NAME:

TMI Pressurizer Spnly Nozzle WOL CONTRACT NO.:

59091 CLIENT: PLANT:

Welding Services Inc. (WSI) Three Mile Island Nuclear Generating Station, Unit 1 CALCULATION TITLE:

Pressurizer Spray Nozzle Weld Overlay Sizing Calculation Project Manager Preparer(s) &

Document Affected Revision Description Approval Checker(s)

Revision Pages

._~,§jllnature & Date Sif!natures & Date

--~--~

I ~ J8 JnitiaJ Issue PL £j 0

rJ a..U-1. --

~> (Qa ~......

R. L. Bax N. Sadeghi RLB 03/1 J/J J NS 03/JO/11 V

A. Alleshwaram AA 03/JOIII Page J of 18 FOJO(i*OIRI

Table of Contents

1.0 INTRODUCTION

3

2.0 DESCRIPTION

OF CONFIGURATION AND REPAIR PROCESS 3 3.0 ASME CODE CRITERIA 3 4.0 LOADS AND DESIGN lNPUTS 7 5.0 WELD OVERLAY THICKNESS SIZING 10 6.0 WELD OVERLAY LENGTH REQUIREMENTS 12 6.1 Structural Reinforcement 12 6.2 Preservice Examination 14 6.3 Area Limitation on Nozzle 14 6.4 Maximuln Overlay Sizing 15 7.0 DISCUSSIONS AND CONCLUSIONS 15

8.0 REFERENCES

17 List of Tables Table I: Safety Factors for Sizing - Circumferential Flaw 5 Table 2: Specified Forces and Moments at the Safe End-to-Elbow Weld Location 8 Table 3: Forces and Moments at Weld Locations 9 Table 4: Dimensions for Overlay Sizing 10 Table 5: Calculated Stresses 11 Table 6: Allowable Stresses and Calculated Stress to Allowable Stress Ratios 12 Table 7: Minimum Required Overlay Length 14 Table 8: Minimum Required Overlay Dimensions 15 List of Figures Figure I: Locations Examined for FSWOL Sizing 10 Figure 2: Full Structural Weld Overlay Geometry, Minimum Dimensions (Schematic Representation) 16 File No.: 1000320.310 Page 2 of 18 Revision: 0 FO\()(>*OIRI

1.0 INTRODUCTION

A weld overlay repair is being designed for the 4" nominal diameter pressurizer spray norLle-to-safe end dissimilar metal weld (DMWI) and the safe end-ta-elbow dissimilar metal weld (DMW2) at Three Mile Island Nuclear Generating Station, Unit I (TMI-I). This calculation documents the required structural sizing calculations for a full structural weld overlay (FSWOL) repair of these welds, based on plant-specific geometry and loadings, and the design requirements of ASME Code,Section XI, Code Cases N-504-3 [4] and N-638-1 [5] (Note: A relief request will be prepared to allow the use of these two Code Cases).

2.0 DESCRIPTION

OF CONFIGURATION AND REPAIR PROCESS The pressurizer spray nozzle is the attached spray line elbow is The FSWOL repair will be performed using primary water stress corrosion cracking (PW8CC) resistant Alloy 52M material deposited around the circumference of the configuration. The overlay material will be deposited using the machine gas tungsten arc welding (GTA W) process. For the Alloy 52M weld overlay filler metal, the selected material is 8B-166, Rod & Bar [3], corresponding to Alloy 690 (58Ni-29Cr-9Fe).

3.0 ASME CODE CRITERIA The applicable ASME Code of repair and replacement for TMI-I is the 2004 Edition of ASME Code, Section Xlll] per Reference 6. The basis for FSWOL sizing is the ASME Code,Section XI, Code Case N-504-J [4] and the ASME Code,Section XI, Division I, Class I [I) rules for allowable flaw sizes in austenitic and ferritic piping (lWB-3640). The ASME Code,Section XI. Code Case N-504-3 [4], and the temper bead welding approach documented in Code Case N-638-1 [5], are used herein and are applied to dissimilar metal welds using nickel alloy filler, Alloy 52M. To detennine the overlay thickness, Code Case N-504-3 refers to the requirements of ASME Code,Section XI, IWB-J640. IWB-3640 of the 2004 edition of the Code refers to Appendix C, which contains the specific methodology lor meeting the allowable flaw sizes. The overlays are to be applied using the CiT AW process, which is a non-flux process. Therefore, for circumferential /laws, the soun:c equations in Reference I, Appendix C, Section C -5320 (limit load criteria) are the controlling allowable flaw size equations for combined loading (membrane plus bending) and membrane-only loading. These equations are valid for flaw depth-to-thickness ratios for flaw lengths ranging from () to 100% ofthe circumference as defined in Reference I, File No.: 1000320.310 Page J of IX Revision: 0 FOJOf>..OIRI

Section C-5320 of Appendix For purposes of designing the overlay, a circumferential flaw is assumed to be 100% through the original wall thickness for the entire circumference of the item being overlaid.

The overlay is sized by using the source equations in Section C-5320 [I].

The allowable bending stress under combined membrane plus bending loads is given by the equation:

Sf' =

SF" (j

Ifl [ .5'F1]

I~--

m Reference I, C-5321, where, (j~ := -:!-l2 20'*

I< t

)

!!. sin 13, for (8 + 13) > 1i The allowable membrane stress is given by the equation:

Reference I, C-5322, where,

~ = arcsin[ 0.5(7)sin e).

and Sc allowable bending stress for a circumferentially flawed pipe

(

(Jb bending stress at incipient plastic collapse SFI1l safety factor for membrane stress based on Service Level as shown in Table I [I, C-2621]

SF/> safety factor for bending stress based on Service Level as shown in Table I [I, C-2621]

a flaw depth t total wall thickness (includes overlay thickness, in this case)

Sf allowable membrane stress for a circumferentially flawed pipe al. m membrane stress at incipient plastic collapse

() half flaw angle [I, Figure C-43 I0-1], 180" or 'Ii tur a 100% full circumferential flaw IJ -- angle to neutral axis of flawed pipe in radians File No.: 1000320.310 Page 4 of 18 Revision: 0 n!306..0IRI

Gi" unintensified primary membrane stress at the flaw location flow stress (Sl' + Su)/2 [I, C-8200(a)J specified value for material yield strength taken at the evaluation (operating) temperature from Reference 3 S/I specified value for material ultimate strength taken at the evaluation (operating) temperature from Reference 3 Safety factors are provided in Appendix C of Section XI for evaluation of flaws in austenitic stainless steel piping. The safety factors used for the weld overlay sizing are shown in Table I and are taken from C-2621 [1].

Table 1: Safety Factors for Sizing Circumferential Flaw Service Membrane Stress Bending Stress Level Safety Factor, SF", Safety Factor, SFb A 2.7 2.3 B 2.4 2.0 ,,_.,~--

C 1.8 1.6 D 1.3 1.4 The overlay thickness must be established so that the flaw assumption herein meets the allowable flaw depth-to-thickness ratio requirement of the source equations [I, C-5320], thr the thickness of the weld-overlaid item, considering combined primary membrane-plus-bending stresses and membrane-only stresses, per the source equations defined previously. Since the weld overlay is an austenitic material and applied with a non-flux welding process, which has high fracture toughness, the limit load failure mode is applicable [1, Figure C-4210-1 for non-flux welds] and hence, limit load evaluation techniques are used here.

The non-overlaid piping stresses for use in the equations are usually obtained from the applicable stress reports for the items to be overlaid. However, in this calculation, they are calculated based 011 forces and moments at the welds using equations from C-2500 of Section Xl, Appendix C as described below.

Primary membrane stress (Gi,,) is given by:

Gi" = pD/( 4t), where:

P operating pressure for the Service Level being considered D outside diameter of the component including the overlay t thickness, consistent with the location at which the outside diameter is taken including the overlay (note that the inside diameter (lD) cladding is not counted toward wall thickness)

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Primary bending stress (O"j,) is given by:

(Jj, DMt/(21), where:

D outside diameter of the component including the overlay d ~

inside diameter, consistent with the point at which the outside diameter is taken (note that the lD cladding is not counted in the inside diameter)

MJ, resultant moment for the appropriate primary load combination for each Service Level (square root of the sum of the squares (SRSS) of three moment components in X, Y, and Z directions)

I moment of inertia, (n:/64) (D 4 et)

The contribution of axial and shear forces to piping stress (other than force couples contributing to moments) is not included based on C-2500 of Section XI, Appendix C [IJ.

The following load combinations are used for the full structural weld overlay.

Service Level A (Normal): Pressure (P) + Deadweight (OW)

Service Level B (Upset): P + DW + Operational Basis Earthquake (OBE)

Service Level C (Emergency): P + OW + Safe Shutdown Earthquake (SSE)

Service Level 0 (Faulted): P +OW +SSE Service Levels A, B, C, and D in the ASME Code [I] are alternatively referred as Normal, Upset, Emergency, and Faulted conditions, respectively, in this evaluation. Per ASME Code, Section Xl C-5311 for the Combined Loading case, test conditions shall be included with the Service Level B load Combination. However, the hydrostatic pressure test is not applicable to the weld overlay repair and is not included in the FSWOL design.

The weld overlay sizing is an iterative process, in which the allowable stresses are calculated and then compared to the stresses in the overlaid component. If the stresses in the component are larger than the allowable stresses in the component then the overlay thickness is increased, and the process is repeated until it converges to an overlay thickness which meets the allowable stresses.

The thickness of the weld overlay is detennined through an iterative process. The thickness of the overlay (to/) is assumed resulting in a total thickness of Up + to/) where tp is the original pipe thickness.

The applied flaw size-ta-thickness ratio based on a FSWOL (flawed through the original pipe wall thickness, II') is ["lUI' + to/). The allowable stresses are then detennined from the source equations (see the beginning of Section 3.0). If this allowable stress value is greater than the calculated stress for the overlaid component, the overlay thickness (t,,,) is reduced. On the other hand, if the allowable stress value is less than the calculated stress for the overlaid component, the overlay thickness (lo/) is increased.

The process is repeated until the assumed overlay thickness results in a stress ratio of the calculated File No.: 1000320.310 Page 6 of 18 Revision: 0 FH30o-(}I RI

Structurlllllt!tJrIty .4ssocia$s, Inc.-

stress to the allowable stress that is equal or less than 1.0. As the maximum allowed value for alt is 0.75

[1, C-5320], t,,! is initial1y set as t,)3. If the overlay thickness of t,l3 meets the allowable stresses for pure membrane and combined membrdne plus bending stresses, then no more iterations are perfonned. If the allowable stresses are not met, then the overlay thickness is increased until the ratio of the computed stress to the allowable stress is less than or equal to 1.0.

In this process, the allowable stresses and adjusted stresses due to overlay thickness iterations are calculated for all applicable Service Levels (A, B, C, and D) and compared. The service level with the maximum ratio of the calculated stress 10 the allowable stress will control the overlay thickness.

The axial length and end slope of the FSWOL are sized to be sufficient to provide for load redistribution (considering both axial force due to pressure and bending loads) from the overlaid component to the weld overlay and back, such that applicable stress limits of the ASME Code,Section III, NB-3200 [2]

are satisfied. Shear stress calculations are performed to assure that the weld overlay length meets these requirements.

4.0 LOADS AND DESIGN INPUTS In order to determine the loads at the nozzle-to-safe end weld, the forces and moments must be transformed such that the revised coordinate system is aligned with the nozzle axis. After the transfonnation performed in the spreadsheet TMl.xlsx, the loads are in a local coordinate system with local-y axial to the nozzle. See Table 2 for the transformed results.

Tables 2 and 3 do not include forces and moments due to thennal expansion of the piping attached to the nozzle. For designing FSWOLs, only primary loads are considered and the secondary loads, such as thennal expansion, need not be included in Ihe design calculations. For the transformed result. all forces and moments are taken on an absolute basis. That is, in Table 2 (Post-Transfonnation), all torces and moments are taken as positive.

The loads shown in Table 2 are assumed to he applied at the safe end-to-elbow weld. These moments are adjusted for the nozzle-to-safe end weld to accounl for the eccentricity between the shear forces (transformed Fvand F:) at the safe end-to-elbow weld and the nozzle-lo-safe end weld centerlines.

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Structurlllnttgrttr .4ssocfaftls. Inc.-

Table 2: Specified Forces and Moments at the Safe End-to-Elbow Weld Location File No.: 1000320.310 Page 8 of 18 Revision: 0 F!l30ft-OIR I

AssaciBtlJs. Inc.-

Table 3: Forces and Moments at Weld Locations Forces and Mnments, Nozzle-tn-Safe End Weld (OMWI)II>>11 Forces and Moments, Sufe End-to-Elbow Weld (OMW2)

F... Fy Fz M:" MJ' Mz MRSS F... Fz M... My ,if; MRSS (lbs) (Ibsl {lbsl tin-Ibs' lin-Ibsl lin-Ibs' (in-Ibsl (Ibs' Ilbs' lio-Ibs' tin-Ibs' tin-Ibs) lin-Ibs)

OW 3 14 IJ 347 1222 803 ---- 3 13 252 1222 781 -

OBE 422 629 493 14628 20102 18888 ---- 422 629 493 10932 20102 15723 -~--

SSE 844 1257 986 29257 40203 37776 *--- 844 1257 986 21864 40203 31446 ,--

Service Level A 3 14 13 347 1,222 803 1,503 J 14 13 252 1.222 781 1,472

{Normall Service Level B 425 643 506 14,976 21,.324 19,691 32,661 425 643 506 11.184 21,324 16,504 29,192 (Unsell Service Level C 847 l,171 998 29,604 41,425 38,580 63,881 847 1,271 998 22,116 41,425 12,227 56,954 (Emen/enev)

Service Level D 847 1,271 998 29,604 41,425 38.580 63,881 847 1,271 998 22,116 41,425 32,227 56,954 (Faulted)

, , ~

Notes: I) Thc nozzle-In-safe end weld moments necoonl for the ecccntncily of 7,5 bctween the eenterhnes of DMW I and DMW2

2) Forces at the nozzle'f<>-sufe end weld nrc assumed equivalent to lhe forces at the safe end-lo-elbow weld, File No.: 1000320..310 Page 9 of 18 Revision: 0 rOlflh-01RJ

5.0 WELD OVERLAY THICKNESS SIZING The normal operating pressure [6], and overlay thickness are shown in Table 4. At the nozzle side of the DMW', Location IB includes the thickness of the nozzle plus the thickness of the 10 cladding, while Location' A considers only the thickness of the nozzle (excluding the thickness of the ID cladding) (see Figure I). An initial alt value of 0.75 (the limiting value as stated in C-5322 of Appendix C of Section Xl [1]) was the initial input to the iteration. The assumed 360 0 flaw results in a flaw length to circumference ratio of '.0. Figure I shows the locations for FSWOL sizing.

IA/IB 2 PRESSURIZER SAFE ELBOW SPRAY NOZZLE END Figure I: Locations Examined for FSWOL Sizing Table 4: Dimensions for Overlay Sizing Location IA Location IB Location 2 Location 3 Location 4 Nozzle Side Nozzle Side Safe End Safe End ofDMWI ofDMWI Side of Side of WiD Clad wi Clad DMWI DMW2 2155 2155 2155 2155 (5) Normal operating pressure of 21 55 psig [6 J is used.

The final calculated membrane stresses ((Jm) and bending stresses ((Jh) at each service level for the pipe +

overlay configumtion are shown in Table 5. This table also shows the ratio of the membrane stress (OJ,,)

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to the flow stress (a/) at the selected locations. The material properties are evaluated at the normal operating temperature of 6500 F [6] using Section n, Part D of the ASME Code [3].

Table 5: Calculated Stresses Location IA Location IB Location 2 Location 3 Location 4 Nozzle Side Nozzle Side Safe End Safe End Elbow ofDMWI ofDMWI Side of Side of Side of wlo Clad wi Clad DMW1 DMW2 DMW2 Service Level Orr' psi 3941 3030 3030 4476 4734 S" psi 27,500 27,500 27,500 27,500 27,500 S.., psi 80.000 80,000 80,000 80,000 80,000 Or, psi 53,750 53,750 53,750 53,750 53,750 OmlOr 0.0733 0.0564 0.0564 0.0833 0,0881 A Normal 0b, psi 127 104 104 195 213 B Upset 0b, psi 2765 2259 2259 3874 4227 C Emergencv 0b, psi 5408 4418 4418 7558 8247 D Faulted 0b, psi 5408 4418 4418 7558 8247 Table 6 shows the allowable stresses as determined from the source equations discussed in Section 3.0.

The membrane and bending stresses from Table 5 are compared to the allowable stresses as shown by the ratios in Table 6. The limiting cases for the membrane and bending stresses are shown in bold. In the limit load analyses, the flow stress of the Alloy 52M weld overlay material is used, consistent with the assumption of a full 360 0 flaw through the original pipe wall for the design of the full structural weld overlay.

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LSlnfctul7l11ffiflfltlrv Associates, Inc.-

Table 6: Allowable Stresses and Calculated Stress to Allowable Stress Ratios Location IA Location 18 Location 2 Location 3 Location 4 Nozzle Side of Nozzle Side of Safe End Safe End Elbow Service DMWI DMWI Side of Side of Side of Level w/o Clad wi Clad DMWI DMW2 DMW2 f:l in rad ians 0.4494 0.4866 0.4866 0.4327 0.4230 0\, psi 18619 20002 20002 18025 17660 Level A Nonnal So, psi 5613 6789 6789 5019 4698

• Level B Upset St, psi 7010 8233 8233 6401 6069 Level C Emergency So, psi 9885 11155 11155 9276 8934 Level D Faulted St, psi 12389 13588 13588 11842 11522 Level A Nonnal otiS . 0.0227 0.0153 0.0\53 0.0389 0.0454 Level B Upset crJSc 0.3944 0.2743 0.2743 0.605\ 0.6965 Level C Emergency crt/Sc 0.5471 0.3961 0.3961 0.8148 0.9231 Level D Faulted crtiSc 0.4365 0.3251 0.3251 0.6382 0.7\58 O

a m, psi 13571 \3438 13438 13776 13833 Level A Nonnal S" psi 5026 4977 4977 5102 5\23 Level B Upset S" psi 5655 5599 5599 5740 5764 Level C Emergency S" psi 7540 7465 7465 7653 7685 LevelD Faulted S" psi 10440 10337 10337 10597 10641 Level A Nonnal crhJS 0.7841 0.6089 0.6089 0.8773 0.9240 Level B Upset cr,nlS, 0.6970 0.5413 0.5413 0.7798 0.8213 Level C Emergency crrr/S, 0.5227 0.4059 0.4059 0.5849 0.6160 Level D Faulted om/S, 0.3775 0.2932 0.2932 0.4224 0.4449 Notes: if" r Bendmg stress at metptent plastIC collapse I, C-5320]

S: Allowable bending stress r\, C-5320]

S, Allowable membrane stress [I, C-5320J dm Membrane stress at incipient plastic collapse rI, C-5320J (All tenns defined in Section 3.0) 6.0 WELD OVERLAY LENGTH REQUIREMENTS The weld overlay length must consider three requirements: (I) length required for structural reinforcement, (2) length required for preservice examination access of the overlaid weld, and (3) limitation on the area of the nozzle surface that can be overlaid.

6.1 Structural Reinforcement Structural reinforcement requirements are expected to be satisfied if the weld overlay length is O.7S.JRi on either side of the susceptible weld being overlaid [4], where R is outside radius of the item and t is the File No.: 1000320.310 Page 120fl8 Revision: ()

F0306-01RI

nominal the at the applicable side of the overlay. However, to assure ASME Code, Sel~tjcm III, l'iIJ'-.h~V\J [2] cornpliance, stress calculations are instead performed to mnmnum reQU1rE~d structural length.

section the length of the overlay is evaluated for axial shear due to transfer of axial load and moment from the overlaid item to the overlay. Subparagraph NB-3227.2 [2] limits pure shear due to Loaul,ng or any Service Level loadings except Service Level D to 0.6Sm. For

......'tH*.. Level D (Faulted) conditions, the stress intensity limit is the lesser of 2.4Sm and O. 7S/I [2, NB-3225 and Appendix F], equiva.lent to the of 1.2S/i1 and 0.3581/ for shear stress, since stress intensity is equal to twice the shear stress. These values are shown in Table 7 for the spray IJU,LLl', attached elbow, and weld overlay materials.

Shear stress around the circumference at the overlay-base material interface due to axial force and moment loading equals:

where, outside radius of overlaid item at crack length of overlay at outside surface of overlaid item on one side of crack shear area, 2 trRaL n:Ro2r P = pressure M resultant moment from piping interface loads at crack Thus Solving for L and equating rwith the allowable shear stress (Sallow) yields:

L [pRoll + MI(n:R/)JISallOlv, where, c\"allllw = O.6Sm (Service Levels A, B, and C)

Sallow Lesser of 1.2S", and 0.35S" (Service Level D)

The evaluation for required length is documented in Table 7 for the pressurizer spray nozzle and elbow.

The overlay weld metal is also evaluated as it may control if the base metal has a higher value of Sm or Suo The greater value of the required overlay length will be taken. The material properties are evaluated at the normal operating temperature of 650°F [6] using Section n, Part D of the ASME Code [3].

Since the overlay ends on the pressurizer spray nozzle at one end and the elbow at the other end, and extends over the safe end, the surface shear transfer into the base metal occurs onto the nozzle and elbow only. In this configuration, the requirements for shear lengths at intennediate locations (safe end) are not relevant and would have no influence on the required overlay. Therefore, they are 110t included herein.

The required overlay length is calculated at Locations 1 and 4 along the nozzle and elbow configuration (both sides ofthe DMWI and DMW2). The evaluation results are presented in Table 7. The design File No.: 1000320.310 Page 13 of 18 Revision: 0 nnn6-OlRI

drawing implements a configuration that meets all the designed FSWOL thickness and length requirements. The lengths shown in Table 7 ensure adequate shear stress transfer along the length of the weld overlay. Service Level C is the most limiting of all cases. This length is sufficient to transfer the imposed loads and maintain stresses (shear) within the appropriate ASME Code allowable limits [2].

Table 7: Minimum Required Overlay Length Location INIB Location IAIIB Location 4 Location 4 Nozzle Side of Nozzle Side of Elbow Side of Elbow Side of DMWI DMWI DMW2 DMW2 R(\,in 2.56.3 2.563 2.25 2.25 Material Alloy 52M SA-508 Class 1 Alloy 52M SA-40.3 WP316 SO" ksi 23.30 17.80 23.30 16.60 I Service Level A 0.6Sm, ksi 13.980 10.680 13.980 9.960 ServLce Level B 0.6Sm , ~si 13.980 10.680

---_._-- 13.980..

---

_ 9.960 Service Level C 0.6S m, ksi 13.980 10.680 13.980 9.960 Service Level 0 1.250 " ksi 27.960 21.360 27.960 19.920 So, ksi 80.00 70.00 80.00 71.80 Service Level 00.3580 , ksi 28.000 24.500 28.000 25.130 Service Level A L, in 0.2027 0.2654 O.ISOO 0.2527 Service Level B L, in 0.3108 0.4068 0.3047 0.4277 Ser"ice Level C L, in 0.4190 0.5485 0.4296 0.6030 SClvice Level 0 L, in 0.2095 0.2742 0.2148 0.3015 6.2 Preservice Examination Weld overlay access for preservice examination requires that the overlay length and profile be such that the overlaid weld and any adjacent welds can be inspected using the required NDE techniques. This requirement could cause the overlay length to be longer than required for structural reinforcement. The specific overlay length required for preservice examination is determined based on the examination techniques and proximity of adjacent welds to be inspected.

6.3 Area Limitation on Nozzle 2

The total weld overlay surface area is limited to 500 in (this value will be speCIfied in the relief request) on the nozzle (carbon steel base material) when using ambient temperature temper bead welding to apply the overlay. Using an outside diameter of 5.125", the maximum length is limited to 500/(1tDo ) = 31.0" on the carbon steel nozzle material. The required overlay length on the nozzle will be less than this limit (see Table 7).

File No.: 1000320.310 Page 14 of 18 Revision: 0

6.4 Maximum Overlay Sizing This calculation documents the minimum overlay thickness and length necessary for stmctural requirements. Additional thickness and length may be added to address inspectability and crack growth concerns. In addition, a maximum overlay thickness (typically an additional 0.25") and a maximum overlay length will be determined. The determination of the maximum length is based on implementation factors and is intended to be large enough so as to not unnecessarily constrain the overlay process. These dimensions will be indicated on a subsequent design drawing to create a "box" within which the overlay is analyzed. In the subsequent analyses, the finite element models use the geometry (minimum or maximum) that will produce conservative results.

7.0 DISCUSSIONS AND CONCLUSIONS Table 8 and Figure 2 summarize the minimum required overlay dimensions. This calculation documents the development ofa weld overlay design for the 4" nominal diameter pressurizer spray nozzle-to-safe end dissimilar metal weld and the safe end-to-elbow dissimilar metal weld at TMI-l. The design meets the requirements of the ASME Code,Section XI, Code Case N-504-3 [4] and ASME Code,Section XI, Appendix C [I] for a full stmctural weld overlay.

The weld overlay sizing presented in Table 8 is based upon the primary loadings documented in Section 4.0 and using the criteria from the ASME Code,Section XI, Appendix C. The overlay thicknesses and lengths listed in Table 8 meet ASME Code stress criteria.

Table 8: Minimum Required Overlay Dimensions Location Thickness, in. Length, in.

Nozzle Side of lA/IB 0.19/0.25 055 DMWI Safe End Side ofDMWl 2 0.25 NA Safe End Side ofDMW2 J 0, I5 NA Elbow Side of 4 0,14 0.61 DMW2 File No.: 1000320.310 Page 15 of 18 Revigion: 0 F0106-01 R I

0.19"/0.25" 0.25" 0.15" 0.14" PRESSURIZER SPRAY NOZZLE ELBOW Figure 2: Full Structural Weld Overlay Geometry, Minimum Dimensions (Schematic Representation)

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8.0 REFERENCES

1. ASME Boiler and Pressure Vessel Code,Section XI, Rules for Inservice Inspection of Nuclear Power Plant Components, 2004 Edition.

2. ASME Boiler and Pressure Vessel Code,Section III, Rules for Construction of Nuclear Facility Components, 2004 Edition.

3. ASME Boiler and Pressure Vessel Code,Section II, Part D, Material Properties, 2004 Edition.

4. ASME Boiler and Pressure Vessel Code, Code Case N-504-3, "Alternative Rules for Repair of Classes 1,2, and 3 Austenitic Stainless Steel Piping,Section XI, Division I."

5. ASME Boiler and Pressure Vessel Code, Code Case N-638-I, "Similar and Dissimilar Metal Welding Using Ambient Temperature Machine GTAW Temper Bead Technique, Section Xl, Division I."

6. Email from William McSorley (Exelon) to Nonnan Eng (SI), dated March 02, 2011,

Subject:

"Status ofpzr Spray SWOL Analysis," includes attached file "GDeBoo Review of 310, 314, 3 15 & 316.doc", SI File No. 1000320.212.

File No.: 1000320.310 Page 17 of 18 Revision: 0 F1l306-01Rl

15. Crane Company, Technical Paper No. 410, "Flow of Fluids through Valves, Fittings, and Pipe," 1976.

16. GPU Nuclear Drawing No. ID-2l2-23-028, Sheet 2 of 6, Rev. I, "LPSIIDecay Heat Removal, Piping Analysis," SI File No. 1000320.204.

File No.; 1000320.310 Page 18 of 18 Revision: 0 IU'(lll*OIRI

R  % k lJUUuuua, uU""yJuy

JL ASSOCI8. tes, Inc. ill) File No.: 1000320.314 Project No.: 1000320 CALCULATION PACKAGE Quality Program: [gJ Nuclear 0 Commercial PROJECT NAME:

TMI-l Pressurizer Spray Nozzle WOL CONTRACT NO.:

59091 CLIENT: PLANT:

Welding Services Inc. (WSI) Three Mile Island Nuclear Generating Station, Unit I CALCULATION TITLE:

Residual Stress Analysis of Pressurizer Spray Nozzle with Weld Overlay Repair Project Manager Preparer(s) &

Document Affected Revision Description Approval Checker(s)

Revision Pages f-.---.

Signature & Date f-- Signatures & Date o 1-32 Illitiallssue Computer Files Apal11a Alleshwaram G. Mukhim AA 03/17/11 GSM 03117111 Craig Jenson CEJ 03/17111 Page I of 32 FIHllldli R 1

Stfl'lctllfallfltl:mritll Associates, Inc.'

Table of Contents 1.0 OBJECTIVE 4 2.0 DESIGN INPUTS 4 2.1 Finite Element Model 4 Material Properties '" 5 3.0 ASSUMPTIONS 6 4.0 METHODOLOGY 7 4.1 Weld Bead Simulation 7 4.2 Welding SitTIulation 8 4.2.1 Internal Pressure Loading 8 5.0 WELDMENT TEMPERATURE GUIDELINE 8 6.0 RESlJLTS OF ANALYSIS 9

7.0 REFERENCES

10 List of Tables Table I: ANSYS Input and Output Files 11 File No.: 1000320.314 Page 2 of32 Revision: 0 FO 106-0 ! R!

Sffl'lcillftli Il1tl~nrilll Associates, Inc;J9 List of Figures Figure I: Applied Structural Boundary Conditions to the Pressurizer Spray Nozzle Finite EJelnent Model 12 Figure 2: As-Modeled Components for the Pressurizer Spray Nozzle 13 Figure 3: As-Modeled Nuggets for ID Repairl (3), ID Repair2 (2) and WOL (J 33) 14 Figure 4: Internal Pressure Loading IS Figure 5: Predicted Fusion Boundary for ID Repair I 16 Figure 6: Predicted Fusion Boundary for ID Repair2 17 Figure 7: Predicted Fusion Boundary for Buffer Layer.. 18 Figure 8: Predicted Fusion Boundary tor Weld Overlay 19 Figure 9: Post ID Repairl Axial Stress at 70°F 20 Figure 10: Post ID Repairl Hoop Stress at 70°F 2J Figure 11: Post ID Repair2 Axial Stress at 70°F 22 Figure 12: Post ID Repair2 Hoop Stress at 70°F 23 Figure 13: Post Weld Overlay Axial Stress at 70°F 24 Figure 14: Post Weld Overlay Hoop Stress at 70°F 25 Figure 15: Post Weld Overlay Axial Stress at 650°F and 2155 psig 26 Figure 16: Post Weld Overlay Hoop Stress at 65()OP and 2155 psig 27 Figure 17: ID Surface Axial Residual Stresses 28 Figure 18: ID Surface Hoop Residual Stresses 29 Figure 19: Path Definitions 30 Figure 20: Post WOL Through-Wall Stress Plots, Paths I through 4 3I Figure 21: Post WOL Through-Wall Stress Plots, Paths 5 and 6 32 File No.: 1000320.314 Page 3 of32 Revision: 0

Structural Integrity Associates. Inc.Jf>

1.0 OBJECTIVE TIle objective of this evaluation is to perfonn a weld residual stress analysis using the ANSY S finite element software [I] on the pressurizer spray nozzle due to a weld overlay (WOL) repair for Three Mile Island Nuclear Generating Station, Unit I (TMI-I). The WOL is applied on the spray nozzle-to-safe end dissimilar metal weld (DMW) and safe end-to-attached pipe dissimilar metal weld (DMW). This analysis includes simulating weld repairs at the inner diameter (10) surface for postulated flaws within the nozzle-to-safe end weld and the safe end-to-attached pipe weld. The ID weld repairs are simulated to provide an unfavorable stress condition (prior to applying the weld overlay) due to the original fabrication of these welds, which are used as the initial condition for the WOL evaluation.

The results will be evaluated to demonstrate that the weld overlay repair has indeed generated a favorable stress condition for the pressurizer spray nozzle, safe end, and the local attached spray pipe by inducing a compressive stress condition on the ID sur1ace. The favorable stress condition minimizes and/or arrests crack initiation/propagation caused by Primary Water Stress Corrosion Cracking (PWSCC) in the susceptible DMW material.

Furthennore, the as-modeled weld overlay repair corresponds to the minimum design dimensions of the pressurizer spray nozzle overlay [2], and thus is considered to bound the case for an as-built weld overlay (typically longer and thicker upon installation).

2.0 DESIGN INPUTS 2.1 Finite Element Model The finite element model (FEM) of the pressurizer spray nozzle, including material properties, is obtained from Reference 3 (input file TMI-SPRAY-RES_64biUNP). The FEM is modeled as an axisymmetric model even though the elbow is present after the safe end-to-pipe weld. This is because the results from the residual analyses are extracted through the nozzle-to-safe end weld and safe end-to-pipe weld, which are sufficiently far away from the elbow. Two 10 weld repairs, nozzle-to-safe end weld repair (10 repairl) and the safe end-to-attached pipe weld repair (10 repair2) are simulated in this analysis to show that the weld overlay repair overcomes the tensile stresses generated by these postulated JD repairs.

Figure I shows the applied structural boundary conditions on the axisymmetric tlnite element model.

while Figure 2 identifies the different components ufthe pressurizer spray nozzle. The weld overla nugget layout used for the residual stress evaluation is shown in Figure 3.

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.t:::lr1lf'tfIY:'lllnttlfJrifu Associates, Inc.

Axisymmetric PLANE55 elements are used in the thennal analysis, while axisymmetric PLANE 182 elements are used in the stress analysis. The weld bead depositions are simulated using the element "birth and death" feature in ANSYS.

The element "birth and death" feature in ANSYS allows for the deactivation (death) and reactivation (birth) of the elements' stilIness contribution when necessary. It is used such that elements that have no contribution to a particular phase of the weld simulation process are deactivated (via EKILL command) because they have not been deposited. The deactivated elements have near-zero conductivity and stiffness contribution to the stmcture. When those elements are required in a later welding phase, they are then reactivated (via EALIVE command).

The analyses consist of a thennal pass to detennine the temperature distribution due to the welding process, and an elastic-plastic stress pass to calculate the residual stresses through the thermal history induced by each weld pass. Appropriate weld heat efficiency, along with sufficient cooling time, are utilized in the thennal pass to ensure that the temperature between weld layer nuggets meets the required interpass temperature as well as obtain acceptable overall temperature distribution within the FEM (i.e.,

peak temperature, sufficient resolution of results, etc.).

2.2 Material Properties The materials of the various components of the model are listed below per Table I of Reference 3.

• Pressurizer Hemispherical Head: SA-516 Grade 70

• Pressurizer Hemispherical Head and Nozzle Cladding: SA-240 TP304

• Pressurizer Spray Nozzle: SA-508 Class I

• Nozzle-to-Safe End Weld: Alloy 82/I 82

• Safe End: SB-166 (Taken as Alloy 6(0)

• Safe End-to-attached Pipe Weld: Alloy 82/182

• Spray Piping: SA-403 WP316 (Elbow)

• Buffer Layer: ER-308L

• Weld Overlay: Alloy 52M

• Nozzle-to-Safe End Weld ID Repair: Alloy 82/182

• Safe End-to- attached Pipe Weld If) Repair: Alloy 82/182 rhe temperature dependent nonlinear material property values are obtained from Reference .3 (input file MProp ~MIS()~NLinear~ TMUNP). This analysis applies the multilinear isotropic hardening material behavior available within the ANSYS f1nite element program.

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SlfllclulallllJtecmitv Associates, Inc.(~'

3.0 ASSUMPTIONS The following assumptions are used in the residual stress evaluation:

1 Assumptions in Reference 3 are applicable in this calculation.

2. The nozzle-to-safe end weld is not included in the residual stress determination as the resulting stress state will be conservatively assumed to be zero, as studies have shown [12] that the as-welded (butt weld) stress state is typically compressive at the ID surface. Imposing the residual effect of the ID repair on a zero stress state is conservative, as this increases the tensile stresses at the ID. Reference 12 documents that even with the significant compressive stresses of the as-welded butt weld, the residual stress state of the final ID weld repair is adverse, with significant axial and hoop tensile stresses. The same assumption holds true for the safe end-to-attached pipe weld.

3. A convection heat tmnsfer coefficient of 5.0 Btu/hr-ftl-op at 70°F bulk ambient temperature is applied to simulate an air environment at the inside surface during the application of both the ID weld repairs.

4. The outside surfaces have a heat transfer coefficient of 5.0 Btulhr-ffl-oP at 70°F bulk ambient temperature during the application of the ID weld repairs to simulate an air environment.

5. During the weld overlay process (for the buffer (stainless steel) layer), the applied heat transfer boundary condition of 5.0 Btu/hr-ft2_oF at 70 0 P bulk ambient temperature was used on the inside surface to simulate an air environment.

6. It is assumed that during the weld overlay process, for the transition from the buffer layer (stainless steel) to the weld overlay (Alloy 52M), the buffer layer cools down to the ambient tempemture of 70°F.

7. During the weld overlay process (for the weld overlay (Alloy 52M) material), the applied heat transfer boundary condition of 5.0 Btulhr-ft2_oF at 70°F bulk ambient temperature was used on the inside surface to simulate an air environment.

8. The outside surfaces have a heat transfer coefficient of 5.0 Btu/hr-ft2 - O P at 70°F bulk ambient temperature during the WOL process (for both the buffer layer (stainless steel) and weld overlay (Alloy 52M) layers) to simulate an air environment.

9. A maximum interpass temperature of 350°F between the depositions of weld nuggets is assumed for all welding processes [5].

10. Additional assumptions including details 011 the heat source and heal efficiency values can be obtained from Reference 6.

11. For the boundary conditions, symmetry is applied at the free end of the vessel, and the nodes at the free end of the modeled pipe arc coupled in the axial direction to simulate the attached piping.

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4.0 METHODOLOGY The residual stresses due to welding are controlled by various welding parameters, thern1al transients due to application of the welding process, temperature dependent material properties, and elastic-plastic stress reversals. TIle analytical technique uses finite element analysis to simulate the multi-pass ID weld repairs, and weld overlay processes.

A residual stress evaluation process was previously developed in an internal Structural Integrity Associates (SI) project. Details of the process and its comparison to actual test data are provided in Reference 6. The same process is used herein. The finite element model of the pressurizer spray nozzle was developed in Reference 3. This model consists of a local portion of the pressurizer top head, pressurizer top head and nozzle cladding, the pressurizer spray nozzle, the nozzle-to-safe end weld, a postulated ID weld repair for the nozzle-to-safe end weld, the safe end, the safe end-to-attached pipe weld, a postulated ID weld repair for the safe end-to-attached pipe weld, a local portion ofthe spray attached piping, and the weld overlay repair (including the buffer layer). The as-modeled weld overlay repair meets minimum structural requirements as well as nondestructive examination (NDE) requirements [13].

401 Weld Bead Simulation In order to reduce computational time, but yet obtain a valid solution, individual weld beads or passes can be lumped together into weld nuggets. This methodology is based on the approach presented in References 7, 8, 9 and 10.

The number of equivalent bead passes is estimated by dividing each nugget area by the area of an individual bead. The resulting number of equivalent bead passes per nugget is used as a multiplier to the heat generation rate. The welding direction is taken to be from the nozzle to the attached pipe. A summary of nuggets for the welds are as follows (see Figure 3):

• The 10 weld repair for the nozzle-to-safe end weld is performed in three layers, with one nugget for each layer. A total of three nuggets are defined for this 10 weld repair.

• The ID weld repair for the safe end-to-attached pipe weld is perfonned in two layers, with one nugget for each layer. A total of two nuggets are defined for this 10 weld repair

• The weld overlay is perfonned in six layers. A total of one hundred thirty three nuggets are defined for the weld overlay:

o Layer one is comprised of nineleen nuggets (including six nuggets for the buffer layer) o Layer two is comprised of eighteen nuggets o Layer three is comprised of nineteen nuggets o Layer four is comprised of thirty nuggets o Layer five is comprised of twenty seven nuggets o Layer six is comprised of twenty six nuggets File No.: 1000320.314 Page 7 of 32 Revision: 0 FOJ06..(IlRI

StrlJClllfJlllnlltlnlFiitf Associatesl 4.2 Welding Simulation The ID repair I is applied first. The end and the attached pipe are in place when the repair is applied, however the ID repair2 elements are deactivated. After the ID repairl is completed, the model is cooled down to a uniform ambient temperature of 70°F. The ID repair2 simulation is then applied. After the ID repair2 is completed, the model is again cooled to a unifonn ambient temperature of70°F. This is followed by the application of the buffer layer, cooling it to an ambient temperature of70°F, and finally followed by the weld overlay simulation. After the weld overlay is completed, the model is cooled to a unifonn ambient temperature of 70°F to obtain residual stresses at room temperature. Then it is heated to a unifonn operating temperature of 650°F [14], and an operating pressure of 2155 psig [II] to obtain the combined residual stresses at operating temperature and pressure.

4.2.1 Internal Pressure Loading The operating pressure of2155 psig is applied to the interior surfaces of the model. An end-cap load is applied to the free end of the attached piping in the fonn of tensile axial pressure, and the value is calculated below. See Figure 4 for the applied pressure loading. Symmetric boundary conditions are applied at the circumferential free end of the pressurizer top head, and the nodes at the free end of the attached piping are coupled in the axial direction as shown in Figure I, to simulate continuity.

2155.1.8007 1 3900 psig where, Pend.cap End cap pressure on attached pipe (psig)

P Internal pressure (psig) fins ide '" Inside radius of attached pipe (in) [3, as modeled]

rout~lde Outside radius of attached pipe (in) [3, as modeled]

The ANSYS input and output files for the analysis are listed in Table I.

5.0 WELDMENT TEMPERATURE GUIDELINE The analytical procedure described in Section 4.0 has provided reasonable results as seen in previous similar analyses [6] when compared to results from test data. This can be demonstrated by observing the fusion boundary prediction of the welds. Figures 5 through 8 show the predicted fusion boundaries for all the welding processes as generated by ANSYS for this specific overlay. The fusion boundaries represent the predicted maximum temperature contour mapping that the weld nugget elements will reach during each welding process. Note that the figures are composiles showing the maximum temperature among all nuggets of each weld. This is made possible by an ANSYS macro (MapTemp.MAC) that File No.: 1000320.314 Page 8 of32 Revision: 0 HH06-HIRI

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reads in the maximum predicted temperatures across the different weld nugget elements during the welding process, and displays them as a temperature contour plot.

The figures show that all weld elements have reached temperatures between 2,674°F and 3,aOO°F. It also shows that the heat penetration depth, where temperatures are above I ,300°F, is similar in size to the heat affected zone (HAl) of between 1/8" and 1/4".

6.0 RESULTS OF ANALYSIS Figures 9 and 10 depict the axial and hoop residual stress distribution for the post ID weld repair of the nozzle-to-safe end weld condition at 70°F, respectively. The axial direction and the hoop direction are with respect to the global coordinate system of the finite element modeL The axial stress is SY and the hoop stress is Sl. It is shown that extensive tensile axial and hoop residual stresses occur along the inside surface of the nozzle in the vicinity of the nozzle-to-safe end ID weld repair.

Figures II and 12 depict the axial and hoop residual stress distribution for the post 10 weld repair of the safe end-to-attached pipe weld condition at 70°F, respectively. Figures 13 and 14 depict the axial and hoop residual stress distribution for the post WOL condition at 70°F, respectively. Figures 15 and 16 depict the resultant residual plus operating condition stress distributions for the post WOL configuration at the operating temperature of 650°F and operating pressure of 2155 psig in the axial and hoop directions, respectively.

Figures 17 and 18 are ID surface stress plots for the axial and hoop directions as a function of distance from the ID weld repair centerline, respectively. The results are plotted for post ID weld repair of nozzle-to-safe end weld, post ID weld repair of safe end-to-attached pipe weld, post WOL at 70°F, and post WOL at 650°F and 2155 psig.

Furthermore, Figures 17 and 18 show that post weld overlay compressive stresses for both the 70°F and operating conditions (650°F/2155 psig) are largely present at the 10 surfaces of the susceptible material.

This would indicate that at any intennediate steady state operating condition (i.e., temperature and pressure) the residual stresses would remain compressive. Any additive loads (I.e., thermal transients) are short tenn in nature and are not relevant to PWSCC concerns. The results suggest that the weld overlay has indeed mitigated the susceptible material against PWSCc.

In addition, through-wall axial and hoop stress results are extracted for various paths defined in Figure 19. Three stress paths are defined through the DMW of nOlzle-lo-safe end and three paths are defined through the OMW of safe end-to-attached pipe. The stress path results are shown in Figures 20 and 21. The results will be used for a subsequent crack growth evaluation in a separate calculation package. Two sets of data are extracted, which are for post WOL at 70°F and for post WOL at 650°F/2155 psig.

The post-processing outputs are listed in Table I. They are further processed in Excel spreadsheet l000320-3/4.xls.

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7.0 REFElU:NCES

l. ANSYS Mechanical and PrepPost, Release 11.0 (w/Service Pack I), ANSYS, Inc., August 2007.

2. SI Drawing No. 1000320.510 "Pressurizer Spray Nozzle Full Structural (FSWOL) Weld Overlay Design Drawing," (for revision number refer to SI Project Revision Log, latest revision).

3. 51 Calculation No. 1000320.312, "Material Properties and Finite Element Models for Pressurizer Spray Nozzle with Weld Overlay Repair," (for revision number refer to SI Project Revision Log, latest revision).

5. ASME Boiler and Pressure Vessel Code, Code Case N-740-2, "Full Structural Dissimilar Metal Weld Overlay for Repair or Mitigation of Class 1,2 and 3 Items,"Section XI, Division I.

6. SI Calculation No. 0800777.304, Rev. I, "Residual Stress Methodology Development and Benchmarking ofa Small Diameter Pipe Weld Overlay, Using MISO Properties."

7. Dong, P., "Residual Stress Analyses of a Multi-Pass Girth Weld: 3-D Special Shell Versus Axisymmetric Models," Journal of Pressure Vessel Technology, Vol. 123, May 2001.

8. Rybicki, E. F., et aL, "Residual Stresses at Girth-Butt Welds in Pipes and Pressure Vessels," U.S.

Nuclear Regulatory Commission RepOlt NUREG-0376, R5, November 1977.

9. Rybicki, E. F., and Stonesifer, R. B., "Computation of Residual Stresses due to Multipass Welds in Piping Systems," Journal of Pressure Vessel Technology, VoL 101, May 1979.

10. Materials Reliability Program: Technical Basis/or Preemptive Weld Overlayslor Alloy 82/182 Butt Weld~' in PWRs (l-v/RP-169), Revision I, EPRI, Palo Alto, CA: 2008.1016602.

II. E-mail attachment, Document 5971-2010-015, Rev. 0, "Design Input for TMI- I Pressurizer Spray Nozzle SWOL," from Bill McSorley (Exelon) to Norman Eng (SI), "Three Mile [sland Transmittal of Design Information," September 17,2010, SI File No. 1000320.210.

12. fdaterials Reliability Program: Welding Residual and Operating Stresses in PWR Alloy 182 Butt Weld~ (MRP-I06), EPRI, Palo Alto, CA: 2004. 1009378.

13. S[ Calculation No. 1000320.310, "Pressurizer Spray Nozzle Weld Overlay Sizing Calculation,"

(for revision number refer to SI Project Revision Log, latesl revision).

14. Email from William McSorley (Exdon) 10 Norn1an Eng (SI), dated March 02,2011,

Subject:

"RE: Status of pzr Spray SWOL Analysis," SI File No. 1000320.212.

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Table I: ANSYS Input and Output Files Input "'lie l)eseriptionJComment TMI~SPRAY*RES ..64biUNP Structural geometry for 2.0 axisymmetric geometry [3]

Material Property data of E, alpha, conductivity, specific heat, and stress strain curves f3]

BCNUGOETZDJNP Weld nuggets definition and boundary conditions file Writes boundary conditions and nugget definitions into PICK1DJNP BCNUOGETZ.o.INP file THERMAL2D.lNP Thennal pass for simulating weld processes STRESS2.oJNP Stress pass for simulating weld processes WELD I"mntrJNP Contains L.oREAD commands for ID Repairl portion of the stress pass WEL.o2 mntrJNP Contains L.oREA.o commands for ID Repaid portion of the stress pass WELD3_mntrJNP Contains LDREAD commands for buffer layer portion of the stress pass Contains LDREAD commands for weld overlay portion of the stress pass POST]ATHJNP Post processing file to extract path stresses POST J.oJNP Post processing file to extract ID surface stresses OntputFlle Description/('omment PATH,T70.0UT Path stress outputs tor post WOL at 70"F PATH,T650 .P2 155.0UT Path stress outputs tor post WOL at 650"F and 2155 psig lD NUST.oUT ID surface nodal coordinate outputs ID WELDLOUT ID surface stress outputs for post I.D Repair I al 70"F

---

~~

ID ~ WELD2.0UT I.D surface stress outputs for 1'051 I.D Repair2 at 70"F II) T70.0UT I In surface stress outputs for post WOL at 70"F If) '1'650 P2155.0lJT ID surface stress for WOLat 1000320~3! 4.xls Excel spreadsheet containing all output data File No.: 1000320.314 Page 11 of32 Revision: 0

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/

Residual stress aDalysis Figure 1: Applied Structural Boundary Conditions to the Pressurizer Spray Nozzle Finite Element Model File No.: 1000320.314 Page 12 of32 Revision: 0 Hl3()f>*UIRI

Structurallntegrdy Associates, Inc:'fI EI..EMENI'S ID Repair for MAT NOM 10 Repair for Safe Nozzle-to-Safe End-to-Pipe Weld End Weld Buffer Layer Safe End Attached Pipe

/, .......................

.-----" IN---- Nozzle Cladding Spray Nozzle Residual stress analysis Figure 2: As-Modeled Components for the Pressurizer Spray Nozzle File No.: I000320.31-t Page 13 of32 Revision: 0

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Resjdual stress 'll"*J.J.\!.;:JC.'

Figure 3: As-Modeled Nuggets for 10 Repairl (3), ID Repair2 (2) and WOL (133)

(Numbers i/1 parenthesis indicate the number olnuggets used /iJl' the corresponding simulation)

(In all instances. ID Repair! corresponds 10 the nozzle-to-safe end IVt!ld and ID Repair2 corresponds to the slife end-to-pipe

.veld)

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Structurallntegrily Associates, Inc.19 ELEM1'J'I'S MAT NUM I

II _ -3900

-3227

-2554

-1881

-1209

'535.9 136.825 809.55 1482 2155

~sidual stress ;ma1.ysis Figure 4: Internal Pressure Loading (Units are in terms o(psi)

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IIlDAL SOUJTICN S~99 SUB -1 TIME=132 TEMP St-N =70.001 SMK =3000 70.001 721.112 1372 2023 2674 395.556 1047 1698 2349 3000 Pre.:1ictro fusion txlun:ia .lot Figure 5: Predicted Fusion Boundary for ID Repairl (Units tire ill terms ofofJ File No.: 1000320.314 Page 16 of 32 Revision: 0 FO.~06.(J I R I

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l'OJAL SOI.1JTICN STEP=65 SUB =1 TTME-262 TEMP Sl-N ""70 SMX -:'1000 70 721.111 13'12 2023 2674 395.556 1047 1698 2349 3000 Predict~l fusion h:mn:hry plot . * * -J Figure 6: Predicted Fusion Boundary for ID Repair2 (Units ure in terms or°F)

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Structural Integrity Associates, Inc. 1P trrJAL SOWI'ICN S'IEP=221 SUB -1 T.lMFF432 TEMP S~ ~70.001 SMX -3000 70.001 721.112 1372 2023 2674 395.556 1047 1698 2349 3000 L Pre::iicta:i fusion murd_a~ry,--"p_l_o_t -,

Figure 7: Predicted Fusion Boundary for Buffer Layer (Units (Ire in terms of OJ-)

(Note: 77te bt(ffer layer is located in the safe end-To-pipe DMW regIOn)

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l\UlAL mI1JTIQIl STEP=4466 SUB =1 TIME'F1334 TEMP SM\1 ='ll. 311 st-lX =3000

• • ,. ' , ' ' ***,':C .... ' . ' , -"".--:" ~~

71.30 722.169 1373 2024 ?675 396.764 1048 1698 2349 3000 Pre:iictEd ttl;; ion h:n.J.rdn Jlot Figure 8: Predicted Fusion Boundary for Weld Overlay (UI/irs are iI/terms n.f OF)

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Structural Integrity Associates. Inc.!f1 NODAL SOLUTION STEP*~7

~U8 el 1'IME~13;>

S1 (AVO.

RS1S-,)

DMX *. 031<12]

SMN '-"65,)4

~;MX -9111'/

.----.,~-

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---~-----,.--,-...,.

~---'---'---.:.-...;.----

x J

-66504 -31344 3816 38916 '74136

• 48924 -13764 21396 565..,6 91111 Residual stress analysis Figure 9: Post 10 Repair I Axial Stress at 70°F (Units are illierms ofpsiJ File No.: 1000320.314 Page 20 of 32 Revision: 0

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NODAL SOLUTION STEP=57 SUB =1 TlME=132 n (AVGl DMX =.031923 SMN =-27987 SMl< =119450

-27987 4777 37540 70304 103068

-11605 21158 53922 86686 119450 Residual stcess analysis Figure 10: Post ID RepairJ Hoop Stress at 70°F (Units lire in terms c?lpsj)

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NODAL SOLUTION STEP=9J sua =1 TIMI';=262 SY (AVG)

RSYS~O OMX ~.031929 SMN =-1J8n SMX =914'17

• 73832 -31092 -352.111 36181 13121

-55462 -18123 18011 54151 91491 Residual stress ~nalysi3 Figure 11: Post 10 Repair2 Axial Stress at 70°F (Units are in terms o{psi)

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NODAL SOLUTION STEP~9J SUB ~1 TlME=262 5:': (AVG)

RSYS~O OM){ ~.031929 SMN =-29346 SM){ =119285

-29346 3683 36712 69741 102770

-12832 20197 51226 86~56 119285 Residual ~ttes8 analysis Figure 12: Post ID Repair2 Hoop Stress at 70°F (Units are in terms o{psi)

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NODAL SOLUTION STElP=3291 SUB TIMlil=l'!7g (AVG)

RltlYS=O Ill'll{ .062.323 SMN =-66352 3M:!: =75283

-50615 -19140 12334 43808 75283 Residual stress analysis Figure 13: Post Weld Overlay Axial Stress at 70°F (Units are in terms (Ifpsi)

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StrllcflJr;ill Intl"mitv Associates, Inc. E' NODAL SOLU'1'ION STEP"'3291 SUB =2 TIME=1474 sz (,h,.VG)

R3Y3"'O DMZ =.062323 314M =-123285 3MX =80733

-100616 -55279 -9941 35396 Residual stress analysis Figure 14: Post Weld Overlay Hoop Stress at 70°F (Units are in terms o(psi)

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NODAL SOLUTIUN 3TEP=J293 SUB =2 T:rME:l494 SY (AVG)

R3YS:O DMX =.302047 3HN =-59120 3HX =65204

,~

--~---.....,.-------

MX

---

~,.

-59120 -31493 23762 51390

- 45306 -17679 9949 37576 65204 Residual st~ess analysis Figure 15: Post Weld Overlay Axial Stress at 650°F and 2155 psig (Units £Ire in terms ofpsi)

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Slfl'1ctllfallntlJlnrlttJ Associates, Inc.1!J rWDAL SOLUTION STErr-3293 3UB =2 TIME=1494 3Z U\.VG)

R3YS=0 DMX =.302047 31'11'1 =-103219 ill'll[ =76952 83200 43162 -3124 36914 76952 stress analysis Figure 16: Post Weld Overlay Hoop Stress at 650°F and 2155 psig (Units are in terms a/psi)

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10 Surface AXial Residual Stress 00 80 60 40

'0

. 20 Safe End

(/)

(/)

(1) 0

.b

(/)

-20

-40

-60

-80

-9 *8 -7 -6 -5 -4 -3 -2 -1 0 2 3 Distancefrom ID Weld Repair Centerline (In)

Figure 17: ID Surface Axial Residual Stresses File No.: 1000320.314 Page 28 of 32 Revision: 0 fO.106-01 R I

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10 Surface Hoop Residual Stress

- - Post 10 weld repatr1 at 70°F  ::lost 10 weld repalr2 at 70"F

& Post ,;veld overiayat 70°F -~ Post weid overlay at 650°F/2155 PStg Figure 18: ID Surface Hoop Residual Stresses File No.: 1000320.314 Page 29 of 32 Revision: 0 Ffl306-01 R I

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,.I F:*n!"..:

'-'~_- In' p,/,

___ ".J Figure J9: Path Definitions (P I, P2. P.1, P4, P5 and P6 denote Stress Paths I, 2, 3. 4. 5 and 6)

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Path 1 Through-Wall Residual Stress Path 2 Through-Wall Residual Stress 80 1";=======::;----------1 100 r=======::::;------I 60 80 60 40 40 20

-40 4 - - - - - - - ' ' t - ' - - ( , ! ' - - - - - - - - - - j

-60 4---~'--~=----------__j

-60

-80

-100 +- - --------- --- r--------- ---- -- ___, J -100 4---,---..----..----,-------,r--l o 0_5 1 o 0_2 0.4 0.6 0,8 Distance from ID Surface lin) Distance from I0 SUrface (in)

Path 3 Through-Wall Residual Stress Path 4 Through-Wall Residual Stress 80 100

-+-Axial70F 80 -*=Hoop70 F 60 Axiel650 F12155psig 40 20

'jh eIII 0 III

<l>

.l::

(J) -20

-40

-40

-60 660F12155p(;jg

• 60 ,~;,,+--;,--- -----------i.

--Hoop650 -80

-100 o 02 0.4 06 08 0 0_2 0.4 0.6 0.8 1 1.2 Distance from 10 Surface (in 1 Distance from to Su.face (in)

Figure 20: Post WOL Through-Wall Stress Plots, Paths 1 through 4 (See Figure 19jor I-Ires\' path locatlOlls)

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Path 5 Through-Wall Residual Stress Path 6 Through-Wall Residual Stress 80 ' 1 ' - - ' - - ' - - -..-'--.~-~-**~- "--"---,

60 00 40 40 2Q

~

.:.I.

l/) 0 0' 0 ..... ,..*.,..,... .. - 1+*,,;'.. ...... _ ..

l/) III Q> ."

li?20 i. 20 f--'.~---'---i£'-Irf,-_._...- .._ - - - - j

-40

-60 Axial 050 F12155 psig Axial 65Cr Fr:215!5 PSiig Hoop 650' Ft2155 psig

-80 *80 +---,-------,-.::==;:===;::===j 0 0.2 0.4 0.6 0,8 a 0.2 04 0.6 0.8 Distance from I D Surface (in, Distance from 10 Surface jin)

Figure 21: Post WOL Through-Wall Stress Plots, Paths 5 and 6 (See Figure 19ftlY stress path locations)

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ATTACHMENT 6 Affidavits

AFFIDAVIT COMMONWEALTH OF VIRGINIA ss.

CITY OF LYNCHBURG

1. My name is Gayle F. Elliott. I am Manager, Product Licensing, for AREVA NP Inc. (AREVA NP) and as such I am authorized to execute this Affidavit.

2. I am familiar with the criteria applied by AREVA NP to determine whether certain AREVA NP information is proprietary. I am familiar with the policies established by AREVA NP to ensure the proper application of these criteria.

3. I am familiar with the AREVA NP information contained in the Structural Integrity Associates, Inc. Calculation Package, No. 1000320.310, Revision 0, entitled "Pressurizer Spray Nozzle Weld Overlay Sizing Calculation," and referred to herein as "Document." Information contained in this Document has been classified by AREVA NP as proprietary in accordance with the policies established by AREVA NP for the control and protection of proprietary and confidential information.

4. This Document contains information of a proprietary and confidential nature and is of the type customarily held in confidence by AREVA NP and not made available to the public. Based on my experience, I am aware that other companies regard information of the kind contained in this Document as proprietary and confidential.

5. This Document has been made available to the U.S. Nuclear Regulatory Commission in confidence with the request that the information contained in this Document be withheld from public disclosure. The request for withholding of proprietary information is

made in accordance with 10 CFR 2.390. The information for which withholding from disclosure is requested qualifies under 10 CFR 2.390(a)(4) "Trade secrets and commercial orfinancial information."

6. The following criteria are customarily applied by AREVA NP to determine whether information should be classified as proprietary:

(a) The information reveals details of AREVA NP's research and development plans and programs or their results.

(b) Use of the information by a competitor would permit the competitor to significantly reduce its expenditures, in time or resources, to design, produce, or market a similar product or service.

(c) The information includes test data or analytical techniques conceming a process, methodology, or component, the application of which results in a competitive advantage for AREVA NP.

(d) The information reveals certain distinguishing aspects of a process, methodology, or component, the exclusive use of which provides a competitive advantage for AREVA NP in product optimization or marketability.

(e) The information is vital to a competitive advantage held by AREVA NP, would be helpful to competitors to AREVA NP, and would likely cause substantial harm to the competitive position of AREVA NP.

The information in the Document is considered proprietary for the reasons set forth in paragraphs 6(b) and 6(c) above.

7. In accordance with AREVA NP's policies governing the protection and control of information, proprietary information contained in this Document have been made available, on a limited basis, to others outside AREVA NP only as required and under suitable agreement providing for nondisclosure and limited use of the information.

8. AREVA NP policy requires that proprietary information be kept in a secured file or area and distributed on a need-to-know basis.

9. The foregoing statements are true and correct to the best of my knowledge, information, and belief.

SUBSCRIBED before me this day of YVlJAf ~ ,2011.

Sherry L. McFaden NOTARY PUBLIC, COMMONWEALTH OF VIRGINIA MY COMMISSION EXPIRES: 10/31/14 Reg. # 7079129 SHERRV L MCfADEN Notarv Public Commonwealth of Virginia 7079129 My Commission Expires Oct 31, 2014

"rff"f'flr~1 1n1F,!nrtJrv Associates, Inc.

5215 Hellyer Ave.

Suite 210 San Jose, CA 95138-1025 Phone: 408-978-8200 Fax: 408-978-8964 www slructintG'Om March 23, 2011 AFFIDAVIT I, Marcos Legaspi Hen'era, state as follows:

(1) I am a Vice President of Structural Integrity Associates, Inc. (SO and have been delegated the function of reviewing the infonnation described in paragraph (2) which is sought to be withheld, and have been authorized to apply for its withholding.

(2) The infonnation sought to be withheld is contained in SI Calculation 1000320.3 10.

Rev. 0, "Pressurizer Spray Nozzle Weld Overlay Sizing Calculation," This calculation is to be treated as SI proprietary inf(mnation, because it contains significant information that is deemed proprietary and confidential to AREV A NP. AREVA NP design input information was provided to SI in strictest confidence so that we could generate the aforementioned calculation on behalf of Sl's client, Exelon Nuclear Company, LLC (Exelon),

Paragraph 3 of this Aflidavit provides the basis for the proprietary detemlination.

(3) 81 is making this application for withholding of proprietary infonnation on the basis that such information was provided to 81 under the protection of a Proprietary/Confidentiality

,md Nondisclosure Agreement between Sl and AREVA Nfl. In a separate Affidavit requesting withholding of such proprietary infonnation prepared by AREV A NP, AREVA NP relies upon the exemption of disclosure set forth in NRC Regulation 1() CFR 2.390(a)(4) pertaining to "trade secrets and commercial or financial information obtained from a person and privileged or confidential" (Exemption 4). As delineated in AREV A NP's Atlidavit, the material for which exemption from disclosure is herein sought is considered proprietary fl)r the ft)llowing reasons (taken directly from ftems 6(b) and 6(c) of AREVA NP's Affidavit):

a) Use of the inltmnation by a competitor would permit the competitor to significantly reduce its expenditures, in time or resources, to design, produce, or market a similar product or service; and

SI Affidavit for Calculation 1000320.310, Rev. 0 March 23, 2011 Page 2 01'2 b) The information includes test data or analytical techniques concerning a process, methodology, or component, the application of which results in a competitive advantage for AREVA NP.

Public disclosure of the infonnation sought to be withheld is likely to cause substantial hann to AREVA NP with which SI has established a Proprietary/Confidentiality and Nondisclosure Agreement.

I declare under penalty of petjury that the above information and request are true, correct, and complete to the best of my knowledge, inf()rmation, and belief:

Executed at San Jose, Califomia on this n rd day of March, 2011.

Nuclear Plant Services State of Califomia Subscribed and swom to (or affirmed) bet()re me on this day of _-'-~~~--:--_ _' 20_\_1_,

Year by proved to me on the basis of satisfactory evidence to be the person who appeared before me (.) (~

(and proved to me on the basis of satisf~tctory evidence to be the person who appeared before me.)

Place Nolary Sell I illllllor Slamp Above 7fYIJ'",ftiflll Ul1f.mn.nl Associates, Inc.

AFFIDAVIT COMMONWEALTH OF VIRGINIA )

) ss.

CITY OF LYNCHBURG )

1. My name is Gayle F. Elliott. I am Manager, Product Licensing, for AREVA NP Inc. (AREVA NP) and as such I am authorized to execute this Affidavit.

2. I am familiar with the criteria applied by AREVA NP to determine whether certain AREVA NP information is proprietary. I am familiar with the policies established by AREVA NP to ensure the proper application of these criteria.

3. I am familiar with the AREVA NP information contained in the Structural Integrity Associates, Inc. Calculation Package, No. 1000320.314, Revision 0, entitled "Residual Stress Analysis of Pressurizer Spray Nozzle with Weld Overlay Repair," and referred to herein as "Document." Information contained in this Document has been classified by AREVA NP as proprietary in accordance with the policies established by AREVA NP for the control and protection of proprietary and confidential information.

4. This Document contains information of a proprietary and confidential nature and is of the type customarily held in confidence by AREVA NP and not made available to the public. Based on my experience, I am aware that other companies regard information of the kind contained in this Document as proprietary and confidential.

5. This Document has been made available to the U.S. Nuclear Regulatory Commission in confidence with the request that the information contained in this Document be withheld from public disclosure. The request for withholding of proprietary information is

made in accordance with 10 CFR 2.390. The information for which withholding from disclosure is requested qualifies under 10 CFR 2.390(a)(4) "Trade secrets and commercial or financial information."

6. The following criteria are customarily applied by AREVA NP to determine whether information should be classified as proprietary:

(a) The information reveals details of AREVA NP's research and development plans and programs or their results.

(b) Use of the information by a competitor would permit the competitor to significantly reduce its expenditures, in time or resources, to design, produce, or market a similar product or service.

(c) The information includes test data or analytical techniques concerning a process, methodology, or component, the application of which results in a competitive advantage for AREVA NP.

(d) The information reveals certain distinguishing aspects of a process, methodology, or component, the exclusive use of which provides a competitive advantage for AREVA NP in product optimization or marketability.

(e) The information is vital to a competitive advantage held by AREVA NP, would be helpful to competitors to AREVA NP, and would likely cause substantial harm to the competitive position of AREVA NP.

The information in the Document is considered proprietary for the reasons set forth in paragraphs 6(b) and 6(c) above.

7. In accordance with AREVA NP's policies governing the protection and control of information, proprietary information contained in this Document have been made available, on a limited basis, to others outside AREVA NP only as required and under suitable agreement providing for nondisclosure and limited use of the information.

8. AREVA NP policy requires that proprietary information be kept in a secured file or area and distributed on a need-to-know basis.

9. The foregoing statements are true and correct to the best of my knowledge, information, and belief.

SUBSCRI BED before me this day of ~lhv ,2011.

Sherry L. McFaden NOTARY PUBLIC, COMMONWEALTH OF VIRGINIA MY COMMISSION EXPIRES: 10/31/14 Reg. # 7079129 SHERRV l. MCFAOEN Notary Public Commonwealth of Virginia 7079129 My Commission Expires Oct 31, 2014

5215 Hellyer Ave Suite 210 San Jose, CA 95138*1025 Phone: 408*978-8200 Fax 408,978-8964 www.structintcom March 23, 2011 AFFIDAVIT I, Marcos Legaspi Herrera, state as follows:

(]) I am a Vice President of Structural Integrity Associates, Inc. (SI) and have been delegated the function of reviewing the information described in paragraph (2) which is sought to be withheld, and have been authorized to apply for its withholding.

(2) The information sought to be withheld is contained in SI Calculation ]000320.314, Rev. 0, "Residual Stress Analysis of Pressurizer Spray Nozzle with Weld Overlay Repair." This calculation is to be treated as SI proprietary infonnation, because it contains significant information that is deemed proprietary and confidential to AREVA NP. AREV A NP design input infonnation was provided to SI in strictest confidence so that we could generate the aforementioned calculation on behalf of SI's client, Exelon Nuclear Company, LLC (Exelon).

Paragraph 3 of this Atlidavit provides the basis f()r the proprietary determination.

(3) SI is making this application for withholding of proprietary information on the basis that such information was provided to SI under the protection of a Proprietary/Confidentiality and Nondisclosure Agreement between SI and AREVA NP. In a separate Atlidavit requesting withholding of such proprietary infonnation prepared by AREVA NP, AREV A NP relies upon the exemption of disclosure set forth in NRC Regulation I() CFR 2.390(a)(4) pertaining to *'trade secrets and commercial or financial infonnation obtained from a person and privileged or confidential" (Exemption 4). As delineated in AREVA NP's Affidavit, the material t()r which exemption from disclosure is herein sought is considered proprietary f{)r the following reasons (taken directly Ii'om Items 6(b) and 6(c) of AREVA NP's Affidavit):

a) Use of the int()[mation by a competitor would permit the competitor to significantly reduce its expenditures, in time or resources, to design. produce, or market a similar product or service; and

SI Affidavit for Calculation 1000320.314, Rev. 0 March 23, 2011 Page 2 of2 b) The infonnation includes test data or analytical techniques concerning a process, methodology, or component, the application of which results in a competitive advantage for AREVA NP.

Public disclosure of the information sought to be withheld is likely to cause substantial harm to AREVA NP with which SI ha.c; established a Proprietary/Confidentiality and Nondisclosure Agreement.

I declare under penalty of peIjury that the above infonnation and request are true, correct, and complete to the best of my knowledge, infonnation, and belief.

rd Executed at San Jose, California on this 23 day of March. 2011.

J'

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M cos LegaspI Herrera, P.E.

Vice President Nuclear Plant Services State of California Subscribed and sworn to (or affinned) before me County on this day of aA{il dl ,20_U_,

Month Year by proved to me on the basis of satisfactory evidence to be the person who appeared before me (. ) ~1 (and 1~ * *'*. .* * * * * * * **

• c * . * **f C. METZGER Commission 1# 1866327

~ Notary Public

• California I Z \. Santa Clara County ~ proved to me on the basis of satisfactory evidence t ... ;..M! SOT"1 tx~rts 2e2J.}~l:J e to be the person who before me.)

Place Notary Seal and/or Stamp Abo\(:

tnu~lnf.::I1 lfl'IPf1r111l Associates, Inc.