Text

Entergys Opposition to NECs Motion to File a Timely New Contention
	ML091420373
Person / Time
Site:	Vermont Yankee File:NorthStar Vermont Yankee icon.png
Issue date:	05/18/2009
From:	Travieso-Diaz M; Entergy Nuclear Operations, Entergy Nuclear Vermont Yankee, Pillsbury, Winthrop, Shaw, Pittman, LLP
To:	; Atomic Safety and Licensing Board Panel
	SECY RAS
References
	50-271-LR, ASLBP 06-849-03-LR, RAS M-421
	Download: ML091420373 (113)
	v • d • e

DOCKETED USNRC May 18, 2009 (1:51pm)

OFFICE OF SECRETARY RULEMAKINGS AND May 18, 2009 ADJUDICATIONS STAFF UNITED STATES OF AMERICA NUCLEAR REGULATORY COMMISSION Before the Atomic Safety and Licensing Board In the Matter of

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Entergy Nuclear Vermont Yankee, LLC

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Docket No. 50-271 -LR and Entergy Nuclear Operations, Inc.

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ASLBP No. 06-849-03-LR

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(Vermont Yankee Nuclear Power Station)

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ENTERGY'S OPPOSITION TO NEC'S MOTION TO FILE A TIMELY NEW CONTENTION I.

INTRODUCTION Applicants Entergy Nuclear Vermont Yankee, LLC and Entergy Nuclear Operations, Inc.

(collectively "Entergy") submit this opposition, pursuant to 10 C.F.R. § 2.309(h)(1) and the provisions in the Board's November 24, 2008 Partial Initial Decision in this proceeding1 and in the Board's March 9, 2009 Order (Clarifying Deadline for Filing New or Amended Contentions)

("March 9, 2009 Order"), to New England Coalition, Inc.'s ("NEC") Motion For Leave to File a Timely New Contention and Motion to Hold in Abeyance Action on this Proposed Contention Until Issuance of NRC Staff Supplemental Safety Evaluation Report, dated April 24, 2009

("NEC Motion").* NEC's proposed contention is inadmissible because it contravenes the Board's admonitions that any new contention must not "rehash or renew any technical challenges that have already been raised and resolved in this proceeding" and "must specifically Partial Initial Decision (Ruling on Contentions 2A, 2B, 3 and 4), LBP-08-25, 68 N.R.C. _

(Nov. 24, 2008) (slip op.) ("LBP-08-25").

2 NEC's filing includes a motion asking that the Board hold in abeyance a ruling on NEC's motion for the admission of its propounded contention until the NRC Staff issues its Safety Evaluation Report Supplement and its Audit Summary relating to the Staffs review of the environmentally assisted fatigue ("EAF") calculations that are the subject of NEC's propounded contention. NEC Motion at 7-8. Entergy and the NRC Staff have filed responses opposing that motion.

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state how the new analyses are not consistent with the legal requirement and the calculations performed for the feedwater nozzle." LBP-08-25 at 67 n.95; see also March 9, 2009 Order at 3.

In addition, NEC's claims are vague, are not supported by the opinion of a technically competent witness, are untimely, and are not shown to be material to the final decision the Board will make with respect to this much litigated issue. Accordingly, NEC's proposed contention must be rejected.

II.

FACTUAL BACKGROUND In August 2007, Entergy issued a set of refined calculations of environmentally assisted fatigue at nine locations. Entergy Exhs. E2-10 through E2-24. These calculations, which came to be known as the "refined fatigue calculations" or the "CUFen Reanalyses," indicated that environmentally adjusted cumulative usage factors at the nine locations would not exceed regulatory limits, and hence component failure due to metal fatigue during the period after renewal of the VY license was not a concern. LBP-08-25 at 14. On September 4, 2007, NEC filed a motion to submit a timely new or amended contention challenging Entergy's CUFen Reanalyses. On November 7, 2007, the Board admitted this contention, designating it "Contention 2A." Memorandum and Order (Ruling on NEC Motions to File and Admit New Contention), LBP-07-15, 66 N.R.C. 261, 267 (2007); LBP-08-25 at 14.

In the course of its review of the CUFen Reanalyses, the NRC Staff raised certain issues with regard to the way the "Green's Function" methodology was used in the CUFen Reanalyses to compute stresses for three reactor components: the feedwater ("FW"), core spray ("CS") and reactor recirculation outlet ("RO") nozzles. The Staff was concerned that the methodology might not lead to conservative estimates of the stress loads at certain locations in those components. To resolve the Staff s concerns, Entergy agreed to perform a confirmatory CUFen 2

analysis without using the Green's Function methodology on one of the three nozzles, the feedwater nozzle, which Entergy and the Staff agreed was bounding. This new environmentally assisted fatigue analysis of the feedwater nozzle was referred to as the "Confirmatory CUFen Analysis." LBP-08-25 at 15.

Entergy provided the Confirmatory CUFen Analysis to NEC in February 2008, and NEC filed a motion to amend Contention 2A to challenge the analysis. NEC asserted that the Confirmatory CUFen Analysis did not validate the results of the CUFen Reanalyses because it only resolved one of many alleged deficiencies in the CUFen Reanalyses and only addressed the feedwater nozzle, which, in NEC's view, is not bounding for the other components. Id. at 15-16.

The Board admitted NEC's proposed contention, which was treated as a subset of Contention 2A and was designated "Contention 2B." Id. at 16; Order (Granting Motion to Amend NEC Contention 2A) (Apr. 24, 2008) at 2.

The Board heard extensive oral testimony on Contentions 2A and 2B at the evidentiary hearing held in Newfane, VT on July 21 and 22, 2008.3 In particular, Messrs. Fitzpatrick and Stevens testified that the Confirmatory CuFen Analysis of the feedwater nozzle performed in early 2008 used the same finite element model, thermal transient definitions, numbers of transient cycles, and water chemistry inputs as were used in the 2007 CUFen Reanalysis for that Prior to the hearing, Entergy had submitted written direct testimony ("Fitzpatrick/Stevens Decl.") and rebuttal testimony ("Fitzpatrick/Stevens Rebuttal Decl.") on these contentions by its witnesses James C. Fitzpatrick and Gary L. Stevens. The direct and rebuttal testimony of Messrs. Fitzpatrick and Stevens, with certain corrections, were admitted into evidence and incorporated into the record of this proceeding. Fitzpatrick/Stevens Decl., Tr. at 763; Fitzpatrick/Stevens Rebuttal Decl., id.; LBP-08-25 at 20-21. Likewise, the NRC Staff had submitted, prior to the hearing, written declarations on these contentions by its witnesses Dr. Kenneth C. Chang ("Chang Decl.")

and Mr. John R. Fair ("Fair Decl.") Those declarations, with certain corrections, were admitted into evidence and incorporated into the record. Chang Dec1., Tr. at 1176; Fair Decl., Tr. at 766-68; LBP-08-25 at 22. Finally, NEC prefiled the direct and rebuttal testimony of its witness Dr. Joram Hopenfeld ("Hopenfeld Dec1." and "Hopenfeld Rebuttal Deci.") on these contentions, and these were admitted into evidence and incorporated into the record.

Hopenfeld Decl. and Hopenfeld Rebuttal Decl., Tr. at 778-79; LBP-08-25 at 23. Messrs. Fitzpatrick, Stevens and Fair and Dr. Hopenfeld presented oral testimony at the hearing. Dr. Chang was unable to testify.

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nozzle. The Confirmatory CuFen Analysis differed in one respect: when the thermal transient stress histories were determined, it computed 6-component stress histories via finite element analysis for each transient, whereas the CUFen Reanalysis used a simplified single stress component difference and subsequently used Green's Functions to obtain the stress time history for all of the transients. Fitzpatrick/Stevens Decl. at A39; Tr. at 916-17, 926-32 (Stevens).4 At the hearing, NEC asserted a number of "errors" in Entergy's 2007 CUFen Reanalysis and the 2008 Confirmatory CUFen Analysis: (1) Entergy used "outdated" statistical equations to calculate the Fen parameters, and should have used instead the results in the 2007 guidance document NUREG/CR-6909; (2) Entergy failed to account for factors that affect the values of the Fen parameters; (3) Entergy did not provide proof that the base metal of the feedwater nozzles is not cracked; (4) Entergy used inappropriate heat transfer equations to calculate the thermal stress for each transient; (5) the number of plant transients estimated to occur during the operating life of VY is not sufficiently conservative; (6) Entergy's calculation of the Fen parameters did not appropriately account for oxygen concentrations and resulting changes in water chemistry; and (7) Entergy failed to perform an error analysis on its calculations. In addition, NEC criticized the 2007 CUFen Reanalysis because it uses a simplified Green's Function methodology. See LBP-08-25 at 32.

The Board found, after receiving extensive testimony on these issues, that NEC's claims were not meritorious and rejected them, except with respect to the use of the simplified Green's Function methodology. LBP-08-25 at 33-46. The Board determined that Entergy's 4 Another difference between the two sets of analyses is that, in the 2008 Confirmatory CuFen Analysis, a maximum environmental correction factor Fen is computed for each paired transient stress state point used for calculating the CUF. The contributions of all stress points are added to produce a composite CUFen. In the 2007 CUFen Reanalysis, on the other hand, a single, maximum Fen was applied to the total CUF resulting from all load pairs, based on the maximum transient temperature for all load pairs. Fitzpatrick/Stevens Decl. at A39. This difference is not relevant to the issues raised in the contention now proposed by NEC.

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Confirmatory CUFen Analysis for the feedwater nozzle "is satisfactory and complies with the regulatory requirements." Id. at 54. However, the Board found that Entergy had failed to show that the Confirmatory CUFen Analysis for the feedwater nozzle proved that the metal fatigue on the CS and RO nozzles during the period of extended operation would "necessarily be below the regulatory requirement of unity." Id. at 55. The Board found that "Entergy must perform the metal fatigue analyses on the core spray and reactor recirculation nozzles (i.e., the CUFens) in compliance with the ASME Code requirements and without using the simplified Green's function methodology in order to satisfy-the ASME Code requirements and 10 C.F.R. §§ 54.21(c)(1) and 54.29(a)." Id.

As a consequence of this finding, the Board concluded that the license renewal for VY can not be authorized until Entergy "either (1) properly recalculates the CS and RR outlet nozzle CUFens such that they demonstrate that these important components will not fail during the PEO (i.e., that the calculations produce a value less than unity), or (2) submits an AMP that demonstrates that aging of these components will be adequately managed during the PEO." Id. at

66. Accordingly, Entergy needs to "(1) recalculate the CUFen analyses for the CS and RR outlet nozzles, in accordance with the ASME Code, NUREG 6583 and 5704, and all other regulatory guidance, (2) resubmit these results to the NRC Staff and serve them on the other parties herein, and (3) either demonstrate that the [time limited aging analyses] are less than unity or submit an adequate [aging management program] for these components." Id. at 67.

The Board ruled that, if Entergy performed confirmatory analyses of the CS and RO nozzles, and "[i]f the CUFen analyses are (1) done in accordance with the above stated guidance and the basic approach used in the Confirmatory CUFen Analysis for the FW nozzle, (2) contain no significantly different scientific or technical judgments, and (3) demonstrate values less than 5

unity, then this adjudicatory proceeding terminates." Id. NEC could file a contention challenging the adequacy of those analyses, but such a contention "must specifically state how the new analyses are not consistent with the legal requirement and the calculations performed for the feedwater nozzle." Id. at 67 n.95.

Entergy proceeded to perform Confirmatory CUFen Analyses of the CS and RO nozzles utilizing the same methodology and approach used in the Confirmatory CUFen Analysis of the feedwater nozzle. On January 8, 2009, Entergy issued and provided to the parties copies of six calculations containing the Confirmatory CUFen Analyses of the CS and RO nozzles.' The methodology applied in these calculations was in accordance with the approach used in the Confirmatory CUFen Analyses for the feedwater nozzle, and contained no significantly different scientific or technical judgments from those used in the feedwater nozzle calculations.6 The results of the calculations showed that the calculated CUFens for the limiting locations of both nozzles are less than unity and are therefore acceptable.7 The NRC Staff conducted a three-day audit of this set of calculations on February 18-20, 2009. In the course of the audit, Entergy and the NRC Staff identified certain minor discrepancies in some of the calculations. While these items did not change the final conclusions of the analyses, Entergy prepared a set of revised calculations that addressed them8 and provided Calculation 0801038.301, Revision 0, "Design Inputs and Methodology for ASME Code Fatigue Usage Analysis of Reactor Core Spray Nozzle;" Calculation No. 0801038.302, Revision 0, "Stress Analysis of Reactor Core Spray Nozzle;" Calculation No. 0801038.303, Revision 0, "Fatigue Analysis of Core Spray Nozzle;" Calculation No. 0801038.304, Revision 0, "Design Inputs and Methodology for ASME Code Fatigue Usage Analysis of Reactor Recirculation Outlet Nozzle;" Calculation No. 0801038.305, Revision 0, "Stress Analysis of Reactor Recirculation Outlet Nozzle;" and Calculation No. 0801038.306, Revision 0, "Fatigue Analysis of Recirculation Outlet Nozzle.".

6 Letter from Matias F. Travieso-Diaz, Esq. to the Board and the parties (January 8, 2009).

7 Id.

8 Calculation No. 0801038.302, Revision 1, "Stress Analysis of Reactor Core Spray Nozzle;" Calculation No.

0801038.303, Revision 1, "Fatigue Analysis of Reactor Core Spray Nozzle;" Calculation No. 0801038.304, Revision 1, "Design Inputs and Methodology for ASME Code Fatigue Usage Analysis of Reactor Recirculation 6

the revised calculations to the parties on March 10, 2009.9 As with the original set, the revised Confirmatory CUFen Analyses of the CS and RO nozzles are done in accordance with the ASME and NRC guidance and used the same approach used in the Confirmatory CUFen Analysis for the feedwater nozzle, contain no significantly different scientific or technical judgments, and demonstrate that the CUFens for the limiting locations of both nozzles are less than unity, and are therefore acceptable.' 0 These are the final CUFen analyses of record for the nozzles.

On April 24, 2009, NEC filed its Motion seeking admission of a contention challenging the Confirmatory CUFen Analyses of the CS and RO nozzles.

II.

NEC'S PROPOSED CONTENTION IS INADMISSIBLE The NEC Motion seeks admission of a contention that broadly challenges as "technically and factually flawed" the "reanalysis of environmentally assisted metal fatigue for Recirculation Outlet (RO) and Core Spray (CS) nozzles" performed by Entergy in accordance with the Board's instructions. NEC Motion at 1, 2. The Motion seeks to differentiate the proposed contention from the issues previously litigated and adjudicated in this proceeding by stating:

NEC does not seek to "rehash technical challenges that have already been resolved in this proceeding." [NEC's witness] Dr. Hopenfeld states specifically how new analyses are not consistent with the legal requirement, the feedwater (FW) nozzle calculations, and the guidance cited in the Board's Partial Initial Decision.

Outlet Nozzle;" Calculation No. 0801038.305, Revision 1, "Stress Analysis of Reactor Recirculation Outlet Nozzle;" and Calculation No. 0801038.306, Revision 1, "Fatigue Analysis of Reactor Recirculation Outlet Nozzle." Calculation 0801038.301, Revision 0, "Design Inputs and Methodology for ASME Code Fatigue Usage Analysis of Reactor Core Spray Nozzle" was not revised.

9 Letter from Matias F. Travieso-Diaz, Esq. to the Board and the parties (March 10, 2009).

]0 Id. Included as Attachments 1-3 to this response are copies of Calculation No. 0801038.304, Revision 1, "Design Inputs and Methodology for ASME Code Fatigue Usage Analysis of Reactor Recirculation Outlet Nozzle;"

Calculation No. 0801038.305, Revision 1, "Stress Analysis of Reactor Recirculation Outlet Nozzle;" and Calculation No. 0801038.306, Revision 1, "Fatigue Analysis of Reactor Recirculation Outlet Nozzle."

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NEC Motion at 6. Notwithstanding this attempt to cast its proposed contention as new and different from previously litigated issues, the contention and the supporting Declaration of Dr.

Joram Hopenfeld (Declaration of Dr. Joram Hopenfeld in Support of New England Coalition's Motion to File a Timely New or Amended Contention on Entergy's Fatigue Reanalysis (Apr. 22, 2009)) ("Hopenfeld Decl.") are concerned "primarily with"'1" two claims that were extensively discussed at the hearing and were resolved in Entergy's favor by the Board: (1) the appropriateness of the heat transfer coefficients used to compute stresses on the reactor nozzles (Tr. 1096-1127); and (2) the dissolved oxygen level in the reactor water during plant transients (Tr. 952-1031). See Hopenfeld Decl. at A5.12 NEC bears the burden to demonstrate that its proposed contention meets three sets of requirements for admissibility:

1 Hopenfeld Decl. at A5.

12 At the end of his Declaration, Dr. Hopenfeld asserts: "Recent discoveries of large cracks in RO nozzles both at the James A Fitzpatrick Nuclear Station (ADAMS Accession Number -ML083300360 "LER 2008-002-00, November 20, 2008) and Oyster Creek Nuclear Generating Station (ADAMS Accession Number ML0090280055

[sic] "Submittal of Analytical Evaluation...") [nozzle indication], January 21, 2009), clearly indicate that Entergy's analysis is not conservative." Hopenfeld Decl. at A24. However, the incipient crack discovered in the RO nozzle at the Fitzpatrick plant was the result of intergranular stress corrosion cracking ("ISGCC"). See James A. FitzPatrick Nuclear Power Plant, LER 2008-002-00, ADAMS Accession No. ML083300360, at 3. As discussed at the hearing in connection with another contention involving the steam dryer, IGSCC is a phenomenon that has nothing to do with environmentally assisted fatigue. See LBP-08-25 at 86 & n.110.

The incipient crack at the Oyster Creek plant occurred at the weld between the RO nozzle and the safe end; the licensee letter to the NRC does not cite fatigue as the potential cause of the crack. See letter RA-09-011 (January 21, 2009) from Exelon Nuclear to NRC, ADAMS Accession No. ML090280055. Therefore, these incidents are irrelevant to the potential vulnerability of the RO and CS nozzles at VY to environmentally assisted fatigue.

In any case, the uncontested evidence in the record demonstrates that the nozzle inspection programs at VY will be effective in detecting incipient cracks in any of the reactor nozzles and will repair them before they risk the integrity of the component. Regular and state-of-the-art ultrasonic ("UT") inspections have revealed no such cracks in the last 20 years. Entergy is obligated to continue those inspections during the period of extended operations in accord with its existing in-service inspection program, and is obliged to take corrective action if a crack is identified. Based on this record, the Board found that Entergy has appropriately addressed the possibility of cracking in the cladding inside the nozzles. LBP-08-25 at 40. Nothing in Dr. Hopenfeld's Declaration suggests otherwise. Finally, there was also uncontested testimony at the hearing that the existence of nozzle cracks is irrelevant to the EAF analysis, since such cracks are presumed identified and eliminated through the plant's ASME Section XI inspection procedures. Tr. 1058-63 (Stevens); LBP-08-25 at 40.

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First, NEC must meet the Board's specific requirements for the admissibility of contentions challenging the confirmatory RO and CS calculations. LBP-08-25 at 67.

Second, NEC must show that its proposed contention meets the general contention admissibility requirements of 10 C.F.R. § 2.309(f)(1).

Third, NEC must demonstrate that the "new" information upon which its proposed contention is based was not previously available, is materially different than information previously available, and was raised in a timely manner based on the availability of the information; or that it meets the admissibility criteria for nontimely filings. See 10 C.F.R. §§ 2.309(f)(2) and (c).

NEC does not - and cannot - meet any of these criteria for admitting its purportedly "new" contention. NEC's proposed contention is inadmissible as set forth further below.

A.

NEC'S NEW CONTENTION CONTRAVENES THE BOARD'S ORDER AND IS THUS INADMISSIBLE NEC's proposed contention is inadmissible because it contravenes the Board's specific instructions on the scope of any contention challenging the confirmatory RO and CS analyses.

NEC's proposed contention must challenge whether the analyses: "are (1) done in accordance with the above stated guidance and the basic approach used in the Confirmatory CUFen Analysis for the FW nozzle, (2) contain no significantly different scientific or technical judgments, and (3) demonstrate values less than unity." LBP-08-25 at 67. NEC's challenge must not "rehash or renew any technical challenges that have already been raised and resolved in this proceeding (e.g., dissolved oxygen, outdated equations, etc.), but rather must specifically state how the new analyses are not consistent with the legal requirement and the calculations performed for the feedwater nozzle." Id. at 67 n.95.

NEC and Dr. Hopenfeld agree that the Confirmatory CUFen analyses for the RO and CS nozzles use the same methodology as that employed in the confirmatory analyses for the 9

feedwater nozzle. 13 Hopenfeld Deci. at A6. And, while NEC asserts that the calculations "are not performed in accordance with the ASME Code, NUREG 6583 and 5704, and all other regulatory guidance," NEC Motion at 4, this assertion is not explained or supported anywhere in the Motion or in Dr. Hopenfeld's Declaration. Therefore, the proposed contention violates one of the limitations imposed by the Board, i.e., that any contention that NEC may seek to raise must allege a failure to follow the methodology used in the confirmatory feedwater analyses or a non-compliance with NRC guidance or industry codes.

Further, all of Dr. Hopenfeld's claims were specifically addressed and resolved against NEC in the hearing. Firt, NEC's argument that equations used by Entergy to calculate the heat transfer coefficient longitudinally along the nozzle "are valid only when the flow inside the pipe is fully developed" (Hopenfeld Decl, at A9) is identical to NEC's argument at the hearing. See, e.g., Tr. at 1108-09 (Hopenfeld). Mr. Stevens refuted that claim by explaining that, because of the high flow velocities in the nozzle, the "entrance effects" preventing the existence of fully developed flow are not present. Tr. at 1124-25 (Stevens). Mr. Stevens referred to the very same figure that Dr. Hopenfeld includes as Attachment 2 to his Declaration, which shows that the entrance effects are not significant at high Reynolds numbers applicable to nozzle flows at issue.

13 Dr. Hopenfeld appears to suggest that Entergy even used the same models and inputs for the confirmatory RO nozzle analysis asit did for the feedwater nozzle analysis. Hopenfeld Decl. at A7. That is demonstrably not the case. While - as directed by the Board - Entergy used in the RO confirmatory analysis the same methodology that was upheld for the feedwater nozzle analyses, the geometries, flow conditions, and heat transfer coefficients are modeled differently for each nozzle. This is recognized in the calculation describing the detailed inputs and methodology for the confirmatory analysis of the RO nozzle, Calculation 0801038.304, Revision 1, where nozzle-specific geometry (Section 2.0 of.304, "Finite element analysis will be performed using a previously-developed axisymmetric finite element model (FEM) of the RO nozzle [7]"), transients (Section 3.3 of.304, "[p]reviously developed thermal and pressure transients [11, Tables 2 and 3] are used for this analysis"), and heat transfer coefficients (Section 3.4 of.304, "Heat transfer coefficients are calculated at 3000 F, as in the previous analysis

[4]") were calculated for the RO nozzle, as was done in the CUFen reanalysis for that nozzle. See Calculation VY-16Q-304, Revision 0, "Recirculation Outlet Nozzle Finite Element Model," Entergy Exh. E2-13 and Calculation VY-1 6Q-305, Revision 0, "Recirculation Outlet Nozzle Stress History Development for Nozzle Green Function," Entergy Exh. E2-14.

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Tr. at 1123-25 (Stevens). The flows through the RO nozzle also have high Reynolds numbers, analogous to those for the feedwater nozzle. Thus, Mr. Stevens' explanation at the hearing holds true for the RO nozzle as well."4 Similarly, Dr. Hopenfeld's claim (Hopenfeld Decl. at A14) that it is inappropriate to use a single heat transfer coefficient for natural convection flow because the heat transfer coefficient varies circumferentially around the RO nozzle was also addressed at the hearing. Dr.

Hopenfeld's claim was refuted by Mr. Stevens. See, e.g., Tr. at 1108-09 (Hopenfeld); Tr. at 111143 (Stevens). 15 Finally, with respect to NEC's dissolved oxygen claim, there was testimony at the hearing on what values of dissolved oxygen were used at the RO line and other reactor locations besides the feedwater line. Entergy witnesses testified that they used the EPRI guidance document BWRVIP-120 (also known as the EPRI BWRVIA Model) to determine the dissolved oxygen values at other lines, such as the recirculation line. Entergy Exh. E2-34, Attach. 2 at 1; Fitzpatrick/Stevens Decl. at A56; Tr. at 1003-05 (Stevens); Tr. at 1030 (Fitzpatrick). The Board, citing this testimony, rejected NEC's argument and found that Entergy's use of industry guidance on dissolved oxygen values "was reasonable and appropriate." LBP-08-25 at 38.

NEC's contravention of the Board's instruction not to "rehash or renew any technical challenges that have already been raised and resolved in this proceeding" (LBP-08-25 at 67 n.95) is even 14 See Tables 3 and 4 in Calculation VY-16Q-305, Revision 0, "Recirculation Outlet Stress History Development for Nozzle Green Function" (Entergy Exh. E2-14). Those tables include calculations of forced convection heat transfer coefficients for 100% flow for different regions of the nozzle. The Reynolds numbers for that flow are provided in the table under "Calculated Parameter." For a flow rate of over 28,000 gpm (100% rated flow), the Reynolds numbers range from 2E6 to 27E6, depending on the fluid temperature.

1 The Board found that Dr. Hopenfeld's concern that it was inappropriate to assume that the flow at the feedwater nozzles is fully developed "has not been substantiated and instead has been fairly rebutted by the evidence presented by Mr. Stevens and Mr. Fitzpatrick." LBP-08-25 at 48. In short, despite the Board's warning, NEC is seeking to rehash a technical claim that has already been examined and rejected.

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more striking because the Board specifically cited "dissolved oxygen" as one of the issues that NEC was not to raise. Id.

B.

NEC'S PROPOSED CONTENTION DOES NOT MEET THE STANDARDS FOR ADMISSIBILITY IN 10 C.F.R. § 2.309(f)(1)

In addition to seeking to relitigate issues already resolved - contrary to the Board's express instructions - NEC's proposed contention is also inadmissible because it fails to meet the general admissibility standards in 10 C.F.R. § 2.309(f)(1). In particular, NEC's proposed contention is vague, immaterial, lacks expert and factual support, and fails to demonstrate the existence of a dispute with Entergy on a genuine, material issue of fact involving the VY license renewal application.

1. NEC Fails to Meet Admissibility Standards with Respect to the Confirmatory CS Nozzle Analyses Although NEC claims that its proposed contention is directed at the Confirmatory CUFen Analyses of both the RO and CS nozzles, the discussion in the NEC Motion pertains only to the RO nozzle analysis. In fact, NEC makes only three vague, non-specific references to the CS nozzle in Dr. Hopenfeld's supporting declaration. See Hopenfeld Decl. at A5 ("[m]y concerns are primarily with the lack of conservatism in the heat transfer calculations and the use of non conservative oxygen concentrations in the analysis of the CS and RO nozzles"); id, at A7 ("[t]he RO and the CS nozzles and FW nozzle are located at different sections of the reactor vessel and therefore their local coolant chemistries differ during transients"); id. at A21 ("[flor the CS and RO nozzles, low alloy steel locations, the maximum oxygen concentrations occur at the lowest temperature during the transient").

NEC's proposed contention, as it relates to the CS confirmatory nozzle analysis, must be dismissed outright because it fails to state what factual issues are being controverted with respect 12

to that analysis. See 10 C.F.R. § 2.309(f)(1)(i). Also, NEC does not offer any explanation of the bases for a challenge to the CS nozzle analysis, in contravention of 10 C.F.R. § 2.309(f)(1)(ii).

NEC does not make any showing that the claims against the CS nozzle analysis are material, as required by 10 C.F.R. § 2.3 09(f)(1)(iv). Nor does NEC provide any alleged facts or expert opinion regarding alleged deficiencies in the CS nozzle analysis, contrary to 10 C.F.R. § 2.309(f)(1)(v). Finally, the NEC Motion does not provide "sufficient information' to show that a genuine dispute exists with Entergy on a material issue of law or fact with respect to any specific aspect of the CS nozzle analysis, as required by 10 C.F.R. § 2.309(f)(1)(vi). Therefore, the proposed contention is inadmissible as it relates to the Confirmatory CuFen Analysis of the CS nozzle because it fails to satisfy the requirements of 10 C.F.R. § 2.3 09(f)(1).

2. NEC's Proposed Contention is Inadmissible with Respect to the Confirmatory Analyses for the RO Nozzle

a.

NEC's Proposed Contention Fails to Specify the Issue of Law or Fact to be Raised NEC fails to set forth the text of its proposed contention in the NEC Motion. Instead, NEC alleges that "Entergy has not properly recalculated the Core Spray and Recirculation Outlet nozzle CUFens." NEC Motion at I (emphasis in original). NEC claims that Entergy's confirmatory calculations "are technically and factually flawed and do not conform to ASME, NRC, or National Laboratory guidance, nor do they fully conform to established engineering practice, or the rules of applied physics." Id. at 2. NEC asserts - without specifying the basis for its assertion - that Entergy allegedly "has not, by this flawed reanalysis, demonstrated that the reactor components assessed will not fail due to metal fatigue during the period of extended operation," "complied with the... Partial Initial Decision," or "credibly demonstrated that its new calculations and analyses for the CS and RO nozzles are consistent with the intent of 10 C.F.R. § 54.21." Id. at 4.

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NEC's description of its contention fails to provide the necessary "specific statement of the issue of law or fact to be raised or controverted." 10 C.F.R. § 2.309(f)(1)(i). NEC does not specify in what respect the confirmatory calculations are "technically and factually flawed" nor to what "ASMIE, NRC, or National Laboratory practice" they fail to conform, nor how they fail to meet "established engineering practice, or the rules of applied physics." Such vague contentions are not permissible under the Commission rules of practice. Duke Energy Corp. (Oconee Nuclear Station, Units 1, 2, and 3), CLI-99-11, 49 N.R.C. 328, 334, 338 (1999);

Northeast Nuclear Energy Co. (Millstone Nuclear Power Stations, Units 2 and 3), LBP-01-10, 53 N.R.C. 273, 279, 302 (2001).16

b.

NEC's Proposed Contention does not Demonstrate that the Issues it Raises are Material Earlier in this proceeding, the Board defined the materiality requirements that a proposed contention must meet with respect to 10 C.F.R. § 2.309(f)(1)(iv):

For a contention to be admissible, a petitioner must show "that the issue raised in the contention is material to the findings the NRC must make to support the action that is involved in the proceeding." 10 C.F.R. § 2.309(f)(1)(iv). An issue is only "material" if "the resolution of the dispute would make a difference in the outcome of the licensing proceeding." 54 Fed. Reg. at 33,172. This means that there must be some link between the claimed error or omission regarding the proposed licensing action and the NRC's role in protecting public health and safety or the environment. Dominion Nuclear Connecticut, Inc. (Millstone Nuclear Power Station, Units 2 and 3), LBP-04-15, 60 NRC 81, 89 (2004), aff d, CLI-04-36, 60 NRC 631 (2004).

LBP-06-20, 64 N.R.C. 131, 149 (2006).

16 NEC addresses the requirements of 10 C.F.R. § 2.309(f)(1)(i) as follows: "However, NEC contends, as explained in some detail in the attached Declaration of Dr. Joram Hopenfeld (Attached as Exhibit A) that Entergy's submitted recalculations do, 'involve complex scientific and technical judgments and discretion, and are not merely ministerial,' but are not performed in [sic], 'in accordance with the ASME Code, NUREG 6583 and 5704, and all other regulatory guidance.'

Thus, Entergy has not, by this flawed reanalysis, demonstrated that the reactor components assessed will not fail due to metal fatigue during the period of extended operation. Nor complied with the requirements set forth in Board's Partial Initial Decision. Nor has Entergy credibly demonstrated that its new calculations and analyses for the CS and RO nozzles are consistent with the intent of 10 C.F.R. § 54.21." NEC Motion at 4. This explanation, however, is as vague as the rest of the Motion.

14

(1)

NEC's Proposed Changes to the RO Nozzle Confirmatory Analysis are Immaterial NEC's proposed contention fails to satisfy 10 C.F.R. § 2.309(f)(l)(iv) with respect to the RO nozzle confirmatory analyses because neither NEC nor Dr. Hopenfeld provide any demonstration that the issues NEC raises are material to the findings that the Board must make. 17 The Confirmatory CUFen Analysis of the RO nozzle indicates that the 60-year CUFen for the safe end of the nozzle is 0.0360, and the 60-year CUFen for the nozzle blend radius is 0.111.

Calculation No. 0801038.306, Revision 1, "Fatigue Analysis of Reactor Recirculation Outlet Nozzle," Attachment 3 hereto, Section 6.0 at 8. Both values are less than the ASME Code allowable value of 1.0, and are therefore acceptable. LBP-08-25 at 29. Indeed, the calculated CUFen at the limiting location - the nozzle blend radius - would have to increase by almost an order of magnitude before the ASME Code allowable was exceeded. Neither NEC nor Dr.

Hopenfeld provide any sound basis to justify that an order of magnitude increase in CUFen at the nozzle's limiting location will result if a new fatigue analysis of the RO nozzle is performed along the lines that Dr. Hopenfeld suggests. 18 Because NEC has failed to demonstrate that a Confirmatory CUFen Analysis of the RO nozzle with the changes NEC suggests would result in values greater than the ASME Code allowable value, NEC has failed to demonstrate that the issues it raises are material.

17 All of the reasons discussed in this subsection why NEC's proposed contention lacks materiality under 10 C.F.R.

§ 2.309(f)(1)(iv) apply also to NEC's failure to demonstrate compliance with 10 C.F.R. § 2.309(f)(1)(vi), which requires that an admissible contention include sufficient information to show that a genuine dispute exists with the applicant/licensee on a material issue of law or fact.

18 Dr. Hopenfeld states that using a dissolved oxygen value of 0.4 ppm to compute the environmental adjustment factor for the nozzle would result in an order of magnitude increase in the CUFen for the nozzle. Hopenfeld DecI.

at A22. As further discussed below, the testimony at the hearing demonstrated that Dr. Hopenfeld's assertion was unsound. The Board rejected Dr. Hopenfeld's claim that a 0.4 ppm value should be used.

15

(2)

NEC's Heat Transfer Coefficient Arguments are Immaterial Dr. Hopenfeld's assertions that "fatigue life is very sensitive to even a very small change in the heat transfer coefficient" and "use of incorrect heat transfer coefficients would result in invalid fatigue life predictions," Hopenfeld Decl. at A8, do not demonstrate any material issue.

First, it has already been established that "entrance effects" are insignificant for high Reynolds number conditions; therefore the heat transfer coefficients do not change. Second, as further discussed below, Dr. Hopenfeld has disclaimed expertise in the area of numerical stress analysis.

Tr. at 831 (Hopenfeld). Third, Dr. Hopenfeld does not indicate what changes would need to be made to the Confirmatory CUFen Analysis of the RO nozzle to address his concerns or provide any indication that, if those changes were made, they would result in having the CUFen for the RO nozzle exceed unit at any nozzle location. Finally, Dr. Hopenfeld provides no citation or other information in support of this bald statement. This is not surprising because the statement is not true.

During the NRC's February 2009 audit of the initial version of the confirmatory analyses for the RO and CS nozzles,' 9 the NRC Staff asked whether the CUFen results for those nozzles were dependent on the value of heat transfer coefficient used in the analyses. Entergy performed a sensitivity analysis, which demonstrated that the effect of changes in the heat transfer coefficient on CUFen estimates is minimal. 2 ) Therefore, the heat transfer -analysis deficiencies J9 Revision 0, later superseded by the March 2009 revised calculations in Revision 1.

20 Attachment 4 hereto is a Declaration of Gary L. Stevens that describes the sensitivity analysis performed by Entergy in connection with the NRC audit that demonstrated that changes in the value of heat transfer coefficient have only a minor impact on the CUFen.

16

raised by Dr. Hopenfeld, even if they were correct, would have little or no impact on the value of CUFen for the RO nozzle and would raise no material issues of fact that needed adjudication.21 Likewise, Dr. Hopenfeld's claim that, for natural convection flow, the heat transfer coefficient varies circumferentially around the pipe is immaterial. As further discussed below, the calculations show that convection flow (as opposed to the forced flow that occurs during some transients) does not contribute significantly to the component stress.

Neither NEC nor Dr. Hopenfeld have provided any indication that the alleged deficiencies in the heat transfer equations for the RO nozzle make a material difference in the value of CUFen for the RO nozzle. Dr. Hopenfeld vaguely asserts:

I believe that Entergy should be required to demonstrate that the incorrect heat transfer equations that they used actually result in conservative CUFens. This can be done by repeating the calculations with heat transfer equations which are valid for the nozzle geometries and which take into account the local variation in the heat transfer instead of using average values. I believe that such calculations will show that the present results are not conservative.

Hopenfeld Decl. at Al 7 (emphases in original). Dr. Hopenfeld does not quantify by how much he claims the CUFen computations for the RO nozzle are not conservative, nor does he claim that he has performed (or is able to perform) such a computation. In the absence of such a quantification, his claim that Entergy's calculation is non-conservative does not rise above mere speculation. Dr. Hopenfeld furthermore does not identify what changes would need to be made 21 Table 6 on p. 15 of the confirmatory CUFen calculation for the RO nozzle (Calculation No. 0801038.306, Revision 1, "Fatigue Analysis of Reactor Recirculation Outlet Nozzle," Attachment 3 hereto) shows that for the safe end of the RO nozzle the main contributor to fatigue is the first load pair combination (loads 47 and 48),

which occurs during transient 9 (see Table 1 on p. 10 of the same calculation). As shown in Table 1 on p. 10 of Calculation No. 0801038.304, Revision 1, "Design Inputs and Methodology for ASME Code Fatigue Usage Analysis of Reactor Recirculation Outlet Nozzle" (Attachment I hereto), Transient 9 occurs during forced flow conditions. Thus, for the safe end of the nozzle, convection flow is only a minor contributor to fatigue, and the circumferential variation on heat transfer coefficient alleged by Dr. Hopenfeld is of no consequence. Also, the main contributor to fatigue in the nozzle comer region occurs during pressure, not temperature, transients; so Dr.

Hopenfeld's allegation is immaterial for that nozzle as well. See Calculation No. 0801038.305, Revision 1, "Stress Analysis of Reactor Recirculation Outlet Nozzle" (Attachment 2 hereto), Section 4.3 at p. 6.

17

to the Confirmatory CUFen Analysis of the RO nozzle to address his concerns or provide any indication that, if those changes were made, they would result in having the CUFen for the RO nozzle exceed unity at any nozzle location. NEC has provided no evidence that the deficiencies that Dr. Hopenfeld alleges raise a material issue, as required by 10 C.F.R. § 2.309(f)(1)(iv).

(3)

NEC's Dissolved Oxygen Claim is Immaterial Dr. Hopenfeld's assertions (Hopenfeld Decl. at A22) that Entergy should have used a 0.4 ppm concentration for dissolved oxygen at the RO nozzle, as "dictated by ANL," and that had it done so, the CUFen "would have been increased by an order of magnitude" lack a factual foundation for the premise that a 0.4 ppm dissolved oxygen concentration should have been used anywhere in the RO nozzle area. The Board heard extensive testimony at the hearing on the significance of the 0.4 ppm figure in the ANL report, NUREG/CR-6909, which it summarized as follows: "With regard to Dr. Hopenfeld's argument that Entergy should have used the DO values of 0.4 ppm for carbon and low-alloy steels and 0.05 ppm for austenitic stainless steels specified in NUREG/CR-6909, Mr. Fitzpatrick pointed out that the NUREG statement was not prescriptive. Tr. at 997-98 (Fitzpatrick). Mr. Fair clarified that NUREG/CR-6909 calls for the use of the DO values of 0.4 ppm and 0.05 ppm only as default values." Tr. at 998 (Fair); LBP-08-25 at 37. Dr. Hopenfeld's dissolved oxygen claim lacks materiality, a fact that he conceded at the hearing. When asked about his concerns regarding the dissolved oxygen issue, he testified:

"First of all, I would like to comment that this is not a major concern." Tr. at 959 (Hopenfeld).

6.0 REFERENCES

19 APPENDIX A: ANSYS INPUT FILE: RONVY.INP.................................................

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List of Tables Table 1: Vessel and Nozzle/Safe End Transients...............................................................

10 Table 2: H eat Transfer Coefficients.......................................................................................

11 Table 3: Temperature-Dependent Material Properties...........................................................

11 Table 4: Recirculation Outlet Nozzle Attached Piping Loads and Dimensions [9, 11]......... 14 Table 5: Membrane Plus Bending Stresses Due to Piping Loads.....................................

14 Table 6: 0% Flow Regions 1 and 3 Heat Transfer Coefficients........................................

15 Table 7: 0% Flow Region 5 Heat Transfer Coefficient...................................................

16 List of Figures Figure 1: Nozzle and Vessel Wall Thermal Boundaries...................................................

12 Figure 2: Coordinate System for Forces and Moments.....................................................

17 Figure 3: RO Nozzle and Safe End Geometry [20]..........................................................

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1.0 OBJECTIVE The objective of this calculation package is to establish the design inputs and methodology to be used for an ASME Code,Section III fatigue usage calculation of the reactor pressure vessel (RPV) recirculation outlet (RO) nozzle at Vermont Yankee Nuclear Power Station (VYNPS)1.

This calculation, along with subsequent calculations for stress and fatigue, are being performed to assess the impact of using finite element analysis using all six components of stress in lieu of the Green's Function approach used in SI project VY-16Q [4, 7, and 11]. Therefore, to the extent possible, inputs from that project will be maintained and used.

1.1 Changes Made in Revision 1 of this Calculation Description of changes made in Revision 1 of this calculation:

a. Transient 9 described in Table 1 was changed to more precisely match the Green's Function analysis.

b. All remaining changes marked throughout this calculation are editorial changes made to the text of the calculation package.

2.0 METHODOLOGY A detailed fatigue usage calculation of the RO nozzle will be performed using the methodology of Subarticle NB-3200 of Section III of the ASME Code [1]. The 1998 Edition including the 2000 Addenda of the ASME Code [10] is also used for material properties. Only the fatigue calculation portion of the ASME Code methodology will be used and the analysis will be a fatigue assessment only, not a complete ASME Code analysis.

Finite element analysis will be performed using a previously-developed axisymmetric finite element model (FEM) of the RO nozzle [7]. Thermal transient analysis will be performed using the FEM for each defined transient. Concurrent with the thermal transients are pressure and piping interface loads; for these loads, unit load analyses (finite element analysis for pressure, and manual calculations for piping loads) will be performed. The stresses from these analyses will be scaled appropriately based on the magnitude of the pressure and piping loads during each thermal transient, and combined with stresses from the thermal transients. Other stress concentration factors (SCFs) will be applied as appropriate.

All six components of the stress tensor will be used for stress calculations. The stress components for the non-axisymmetric loads (shear and moment piping loads) can have opposite signs depending upon which side of the nozzle is being examined. Therefore, when combining stress components from these loads with stress components from thermal transients and other loads, the signs of the stress components will be adjusted to maximize the magnitude of the stress component ranges. The fatigue analysis will be performed at locations that were determined in a previous calculation [4]. Stresses will be linearized at these locations.

1 The methodology described and applied herein and in the two additional recirculation outlet nozzle fatigue calculations is in accordance with the approach used in the SIA calculations for the feedwater nozzle [16, 17, 18] and contains no significantly different scientific or technical judgments used in those calculations.

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The linearized primary plus secondary membrane plus bending stress will be used to determine the value of Ke to be used in the simplified elastic-plastic analysis in accordance with ASME Code NB-3200 methodology. Environmental fatigue multipliers will be applied in accordance with NUREG/CR-6583 [2]

for the low alloy steel forging and NUREG/CR-5704 [15] for the stainless steel safe end.

3.0 ASSUMPTIONS / DESIGN INPUTS 3.1 Assumptions

1. Extended power uprate (EPU) effects are considered as being applied to the entire 60-year period of operation. The higher pressures, flows, and temperatures at uprate conditions are used in determining and applying heat transfer coefficients [4, Section 3.2] [11, Section 4.1].

2. The Boltup transient does not affect the RO nozzle because there is no pressure or temperature change, and the nozzle is sufficiently removed from the vicinity of the flange such that stresses due to head stud tensioning are insignificant at the nozzle location [8]. The Boltup transient is therefore excluded from the transients analyzed.

3. For the blend radius and safe end transient definitions, steady state condition time steps were assumed to be 5,000 seconds for Transients 3, 5, 6, 8, 9, and 40,000 seconds for Transients 1, 2, 4, 7, 10.

4. The effect of non-uniform geometries is judged to be insignificant for flow inside the safe end, because of the smooth transition and small geometry changes, as shown in Figure 3. The nominal inner diameter for all heat transfer regions was used to calculate heat transfer coefficients.

5. Density, p, and Poisson's ratio, v, used in the FEM are assumed typical values of p = 0.283 lb/in 3 and v = 0.3, respectively.

6. For purposes of linearizing stress at the nozzle blend radius, the cladding is ignored.

7. Stress components due to piping loads are scaled assuming no stress occurs at an ambient temperature of 70'F and the full values are reached at reactor design temperature, 575°F, as was done in the previous analysis [11, Section 3.4].

8. Consistent with Reference [4], 12% of the available temperature difference (AT) between the fluid and surface was assumed for all natural convection thermal heat transfer coefficients.

9. The instant temperature change for transients is assumed as a 1-second time step.

3.2 ASME Code Edition The analysis will be performed in a manner consistent with the fatigue usage rules in NB-3200 of Section III of the ASME Code; the 1998 Edition with Addenda through 2000 [1] will be used, for consistency with the previous analysis [11].

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3.3 Transients Previously developed thermal and pressure transients [11, Tables 2 and 3] are used for this analysis. The transients to be evaluated are shown in Table 1. For each transient, the time, nozzle fluid temperature, RPV pressure, percent reactor recirculation flow rate, and number of cycles are included. In some cases, flow rates and nozzle temperature values from the nozzle thermal cycle diagram [8, Attachment 1, p. 4] are used to reduce excess conservatism. Note that the only difference between the vessel and the safe end/nozzle transients is the temperature difference between the two regions for Transient 9.

At the inside surface of the RPV, the Region B or B 1 bulk fluid temperature from the reactor thermal cycle diagram [8, Attachment 1, p. 2] shall be applied.

3.4 Heat Transfer Coefficients Heat transfer coefficients are calculated at 300* F, as in the previous analysis [4]. The heat transfer coefficients for the 100% flow and 50% flow cases were calculated from Reference [5] as follows:

f )

.(2 0.2 hDf = h300 ( fDfII2

~25) 1Df)

Where:

hDf= the heat transfer coefficient at a Diameter and flow rate h300 = the heat transfer coefficient from Reference [5] at 300°F,f= 25 ft/sec, and D = 26" = 4,789 BTU/hr-ft2-°F fDf= the flow velocity corresponding to hDf (ft/sec)

DDf ý the diameter corresponding to hDf (in)

The heat transfer coefficients for 0% flow were calculated in spreadsheet HTCOEF.xls for natural convection and are shown in Tables 6 and 7.

As shown in Figure 1, the following heat transfer coefficients were applied:

Region 1 The heat transfer coefficient, h, for 100% flow is 4789

'17.364 2

08 (62.8) = 3583 BTU/hr-fý-°F at 300'F, where 17.364 ft/sec is converted from 28,294 GPM and 25.8 in ID [20].

The heat transfer coefficient, h, for 50% flow is 4789 682

( 26 2= 2058 BTU/hr-ft2-Y 25 )

~25.8)

OF at 300'F, where 8.682 ft/sec is converted from 14,147 GPM and 25.8 in ID [20].

(2.084/°

( 26 ° The heat transfer coefficient, h, for 12% flow is 4789 2

4 25 2

I2 = 657 BTU/hr-ft2-OF

( 25)k 25.8) at 300'F, where 2.084 ft/sec is converted from 3,395 GPM and 25.8 in ID [20].

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The heat transfer coefficient, h, for 0% flow is 112 BTU/hr-ft2 -°F at 300'F. (Table 6, for natural convection)

Region 2 The heat transfer coefficient for Region 2 is linearly transitioned from the value of the heat transfer coefficient used in Region 1 to the value used for Region 3.

Region 3 (the point between Region 2 and Region 4)

The inside diameter of Region 3, as measured on the ANSYS model, is 35.49 inches.

The heat transfer coefficient, h, for 100% flow is 4789 ( 9.17 6 0'.(

= 2018 BTU/hr-ft2-OF at 300'F, where 9.176 ft/sec is converted from 28,294 GPM and 35.49 in. ID.

The heat transfer coefficient, h, for 50% flow is 4789 4.588

(=

1159 BTU/hr-ft2-

\\ 25 )"35.49)

=15'T/rf2 OF at 3000F, where 4.588 ft/sec is converted from 14,147 GPM and 35.49 in. ID.

The heat transfer coefficient, h, for 12% flow is 4789 1.101 08 26°02 = 370 BTU/hr-ft2 2oF

( 25 )

" 35-.

=304T/h-t)° at 300'F, where 1.101 ft/sec is converted from 3,395 GPM and 35.49 in. ID.

The heat transfer coefficient, h, for 0% flow is 112 BTU/hr-ftZ-°F at 300'F. using the same HTC as Region 1 (Table 6, for natural convection)

Region 4 The heat transfer coefficient for Region 4 (Nozzle Blend Radius) is linearly transitioned from the value of the heat transfer coefficient used in Region 3 to the value used in Region 5.

Region 5 A value of 0.5 x Region 1 HTC from Reference [5, page I-T9-4, 6] is used to simulate the interior of the RPV shell for all conditions.

The heat transfer coefficient, h, for 100% flow is 0.5 x 3583.3 = 1,792 BTU/hr-ft2 -OF at 3000F.

The heat transfer coefficient, h, for 50% flow is 0.5 x 205 8.1= 1029 BTU/hr-ft2-OF at 3000F.

The heat transfer coefficient, h, for 12% flow is 0.5 x 657.2= 329 BTU/hr-ft2 -OF at 3000F.

The heat transfer coefficient, h, for 0% flow is 101 BTU/hr-ft2e-F at 300'F. (Table 7, for natural convection) by using 40 in. hydraulic diameter [5].

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Region 6 The heat transfer coefficient, h, is 0.4 BTU/hr-ft2 -OF [5].

A summary of the heat transfer coefficients (HTC) to be used is shown in Table 2.

3.5 Finite Element Model The ANSYS program [6] will be used to perform the finite element analysis. A previously developed axisymmetric model will be used [7, file RONVY.INP], except that temperature-dependent material properties will be used. Table 3 shows the applicable material properties [10].

Stresses will be extracted and linearized at two locations, both on the inside surface of the model, one at the safe end, and one at the blend radius, as was done previously [4].

3.6 Nozzle Blend Radius Pressure Stress The axisymmetric model has the effect of modeling the cylindrical RPV as spherical. The following paragraphs describe the details of the modeling used to account for the differences in this approximation and the actual geometry of two intersecting cylinders.

The radius of the vessel in the finite element model was multiplied by a factor of 2 to account for the fact that the vessel portion of the axisymmetric model is a sphere, but the true geometry is a cylinder. The equation for the membrane hoop stress for a sphere is:

(pressure) x (radius) 2 x thickness The equation for the membrane hoop stress in a cylinder is:

(pressure) x (radius) 0"=

thickness The factor of two was verified in Reference [4], where actual stress results were compared to the results of this analytical form.

The pressure stress components for the safe end and blend radius paths will be extracted using ANSYS [6].

3.7 Piping Inteeface Loads Per Reference [9, 11], the RO nozzle piping loads, which conservatively use the design loads for the seismic, thermal and deadweight load combination, are stated in Table 4 along with relevant dimensions.

The coordinate system used for these are shown in Figure 2 and is consistent with Reference [9]. The finite element model coordinate system is shown in Figure 1.

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3.8 SCFs, Safe End At the safe end inside surface, guidance is taken from the piping analysis rules in Subarticle NB-3600 of Section III of the ASME Code [1]. The stresses caused by the piping will be hand calculated and require a stress concentration factor, if appropriate. The stress concentration factor for the safe end location is 1.53 [5, page I-S9-4E, Table 5]. This value is conservatively used for both the C2 and K2 values required by the ASME code [1, NB-3600]. The piping loads are relatively minor in comparison to the other loads this nozzle experiences so the conservative C2 and K2 values will have a small impact on the analysis.

These factors are conservatively applied to all six components of the stress tensor.

3.9 Environmental Fatigue Multipliers The environmental fatigue multipliers for the safe end will be calculated in accordance with NUREG/CR-5704 methodology [15], and the environmental fatigue multipliers for the nozzle blend radius will be calculated in accordance with NUREG/CR-6583 methodology [2].

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Table 1: Vessel and Nozzle/Safe End Transients Transient Time Temp Time Step Pressure Flow Rate Transient Time Temp Time Step Pressure Flow Rate Number US W*u p

GFM Number LsS

.(s) ipsiq)

(GPMI

1. Normal Startup with 0

100 0

14147.0

6. Reactor Overpressure 0

526 1010 28294 Heatup at 100Flhr 16164 549 16164 1010 (50%)

1 Cycle (1,2) 2 526 2

1375 (100%)

300 Cycles (2) 56164 549 40000 1010 32 526 30 940

2. Turbine Roll and 0

549 1010 28294 1832 526 1800 940 Increase to Rated Power 1

542 1

1010 (100%)

2252 549 420 1010 300 Cycles (1, 2) 601 542 600 1010 2312 549 60 1010 602 526 1

1010 2313 542 1

1010 40602 526 40000 1010 1

2913 542 600 1010

3. Loss of Feedwater 0

526 1010 28294 2914 526 1

1010 Heaters 1800 542 1800 1010 (100%)

7914 526 5000 1010 Turbine Trip 25% Power 2100 542 300 1010

7. SRV Blowdown 0

526 1010 28294 10 Cycles (2) 2460 526 360 1010 1 Cycle (2) 600 375 600 170 (100%)

3060 526 600 1010 11580 70 10980 50 3960 542 900 1010 51580 70 40000 50 4260 "542 300 1010

8. SCRAM Other 0

526 1010 28294 6060 526 1800 1010 228 Cycles (1, 2) 15 526 15 940 (100%)

11060 526 5000 1010 1815 526 1800 940

4. Loss of Feedwater 0

526 1010 0

2235 549 420 1010 Pumps 3

526 3

1190 (0%)

2295 549 60 1010 10 Cycles (1, 2) 13 526 10 1135 2296 542 1

1010 233 300 220 1135 2356 542 60 1010 2213 500 1980 1135 2357 526 1

1010 2393 300 180 885 7357 526 5000 1010 6773 500 4380 1135

9. Improper Startup 0

526 1010 3395 7193 300 420 675 14147 1 Cycle (1,2) 1

13011, 1

1010 (12%)

7493 300 300 675 (50%)

27 130 i" 26 1010 11093 400 3600 240 28 526 1

1010 16457 549 5364 1010 5028 526 5000 1010 16517 549 60 1010'

10. Shutdown 0

549 1010 14147 16518 542 1

1010 28294 300 Cycles (2) 6264 375 6264 170 (50%)

17118 542 600 1010 (100%)

6864 330 600 88 17119 526 1

1010 16224 70 9360 50 57119 526 40000 1010 56224 70 40000 50

5. Turbine Generator Trip 60 Cycles (1, 2) 0 10 15 30 1830 2250 2310 2311 2911 2912 7912 526 526 526 526 526 549 549 542 542.

526 526 10 5

15 1800 420 60 1

600 1

5000 1010 1135 1135 940 940 1010 1010 1010 1010 1010 1010 28294

11. Design Hydrostatic (100%)

Test 120 Cycles (2)

I 100 0

1100 50 1981 (7%)

12. Hydrostatic Test 100 1

50 1981 I Cycle (2) 1 5 6 3 (7%)

1. The instant temperature change is assumed as 1-second time step.

2. The number of cycles is for 60 years [8].

3. 130°F is the Region 1 temperature for Transient 9, whereas the blend radius is at 268°F and the vessel is at 2687F, as was modeled previously [11].

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Table 2: Heat Transfer Coefficients Flow Rate Thermal Region 100%

50%

12%

0% (Natural Convection)

Region 1 3583 2058 657 112 Region 2 Linear transition from Region 1 and Region 3 values Region 3 2018 1159 370 112 Region 4 Linear transition from Region 3 and Region 5 values Region 5 1792 1029 329 101 Region 6 0.4 for all flow rates Note: All Heat transfer coefficients are in units of BTU/hr-ft2-°F and are evaluated at 300'F.

Table 3: Temperature-Dependent Material Properties Material No.

Young's Tempera-

Modulus, Description tue F EX16

ture, (F

E x 10)

(psi)

Mean Coefficient of Thermal Expansion, a x 10.6 (in/in-0F)

Conductivity, k

(BTU/hr-ft-0 F)

(see Note 1)

Diffusivity, d

(ft2/hr)

Specific Heat, ep (BTU/Ibm-*F)

(see Note 4) 4 SA533 Grade B, 70 29.2 7.0 23.5 0.458 0.105

[Vessel Wall]

200 28.5 7.3 23.6 0.425 0.114 (Mn-'1/2Mo-1/2Ni) 300 28.0 7.4 23.4 0.401 0.119 400 27.4 7.6 23.1 0.378 0.125 500 27.0 7.7 22.7 0.356 0.130 600 26.4 7.8 22.2 0.336 0.135 2

SA-508 Class 2 70 27.8 6.4 23.5 0.458 0.105

[Nozzle Forging]

200 27.1 6.7 23.6 0.425 0.114 300 26.7 6.9 23.4 0.401 0.119 400 26.1 7.1 23.1 0.378 0.125 500 25.7 7.3 22.7 0.356 0.130 (See Note 2) 600 25.2 7.4 22.2 0.336 0.135 1,3 SA 240 Type 70 28.3 8.5 8.6 0.151 0.116 304, SS Clad, 200 27.6 8.9 9.3 0.156 0.122 SA182 Type 300 27.0 9.2 9.8 0.160 0.125 F316 400 26.5 9.5 10.4 0.165 0.129

[Clad, Safe End]

500 25.8 9.7 10.9 0.170 0.131 (see Note 3) 600 25.3 9.8 11.3 0.174 0.133 Notes:

1.

2.

3.

4.

Convert to BTU/sec-in-°F for input to ANSYS.

Properties of A508 Class II are used (3/4Ni-1/2Mo-1/3Cr-V).

Properties of 18Cr - 8Ni austenitic stainless steel are used.

Calculated as [k/(pd)]/12 3.

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Region 5 Region 6 Region 4 S

Region 2 Region 3 Region 1 x

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4.0 CALCULATIONS 4.1 Piping Interface Loads From general structural mechanics [14], the membrane plus bending stresses at the inside surface of a thick-walled cylinder are:

(z5 = axial stress due to axial force = Ft/A (Y2 = axial stress due to bending moment = My(ID/2)/I (z5 = yz 1 + uz2 Tr0 = shear stress due to torsion = Mz(ID/2)/J crz = shear stress due to shear force = 2Fxy/A, where Fx, Fy, F,, Mx, My, and M, are forces and moments at the pipe-to-safe end weld MxL = moment about x axis translated by length z = -L = Mx - Fy L MyL = moment about y axis translated by length z = -L = My + Fx L Mxy = resultant bending moment = (MxL 2 + MyL2)0 5 FXy = resultant shear force = (Fx2 + Fy2)0 5 ID, OD = inside and outside diameters A = area of cross section = (it4)(OD2 - ID 2)

I = moment of inertia = (it/64)(0D 4 - ID 4)

J = polar moment of inertia = (n/32)(OD4 - ID 4)

The.shear stresses are expressed in a local coordinate system with r radial (X in ANSYS coordinates), 0 circumferential (Z in ANSYS coordinates), and Z axial (Y in ANSYS coordinates). Tables 4 and 5 show the calculation of stresses; ID, OD, and L are taken from the previous piping load stress calculations [11, Section 3.4]. Forces and moments are taken from Reference 11, Table 1. Note that the IDs shown in Table 4 for the safe end and nozzle blend radius (25.938" and 37.368", respectively) represent the two most limiting locations for the nozzle (See Figure 3), and therefore do not represent the ID values where the HTCs were calculated.

File No.: 0801038.304 Page 13 of 20 Revision: 1 F0306-01i

V Structural Integrity Associates, Inc.

Table 4: Recirculation Outlet Nozzle Attached Piping Loads and Dimensions [9, 11]

Safe End Nozzle Blend Radius Fx, kip 20.0 20.0 Fy, kip 20.0 20.0 Fz, kip 30.0 30.0 M,, kip-in 2004.0 2004.0 My, kip-in 3000.0 3000.0 Mz, kip-in 2004.0 2004.0 L, in 4.25 42.77 OD, in 28.38 55.88 ID, in 25.938 37.368 Table 5: Membrane Plus Bending Stresses Due to Piping Loads Mýj, kip-in MyL, kip-in Mxy, kip-in FV, kip-in A, in2 I, in4 J, in4 CYzI, ksi a,,, ksi cz, ksi trO, ksi

'rz, ksi Safe End 1919.00 3085.00 3633.15 28.28 104.18 9624.85 19249.69 0.288 4.895 5.183 1.350 0.543 Blend Radius 1148.60 3855.40 4022.86 28.28 1355.76 382912.48 765824.95 0.022 0.196 0.218 0.049 0.042 File No.: 0801038.304 Revision: 1 Page 14 of 20 F0306-0E

Structural Integrity Associates, Inc.

Table 6: 0% Flow Regions 1 and 3 Heat Transfer Coefficients Pipe Inside Diameter, 2

=

0, inches=

2.150 ft

.. r 0.655 m

Outer Pipe, Inside radius, r.=

12.9 inches =

1.075 ft 0.328 rn Inner Pipe Outside Diameter, D I

inches =

0.000 ft 0.000 m

Inner Pipe, Outside radius, r =

0 inches =

0.000 ft 0.000 m

Fluid Velocity, V =

0.000 flsec =

gpm=

0 ilthr CharacteristlcLengthL=D=

2.150 ft=

0.855 m

T,

-T

.. AT= assumedt be 12% at fluidtemperature=

8.40, 12.00D 24.001 3,.000 48.00D 60.00 72.00 F

=

4.67 6.67 13.33 20.00 26.67 3133 40.00 T

Value at Fluid: Temperature. T [i2]

Units Conversion 70 100 200 300 400 500 600 T

Water Property Factor [19]

21.1t 37.78 93.33 148.89 204.A4 260.00 315.56 T

k 1.7307 0.5997 0.6300 0.6784 0.618M 0.6611 0.8040 0.5071 W/m-'C Cmermsl~~ý-ý 9enuc?2ty 0.46 0.60 032

.00 0320 0.3490 D.2930 Bta/tir-ft-'F 4.1869 4.185 4.179 4.229 4.313 4.522 4.982 6.322 Ulkg/-C (Specific Heat) 1.000 0.993 1.010 1.030 1.080 1.190 1.510 Btu/lbr-'F p

16.018 997.1 994.7 962.7 917.8 658.6 784.9 679.2 kgvly (Oensity) 62.3 62.1 60.1 57.3 53.6 49.0 42.4 Ibnft i f -

*ii

K '*;Y*:i 4} :

........ m } *;

b

1.

1.800-04 3203466041000 0.-..

1.8-0---- 3.15-0 em/m-;C (Volurmetric Rate ofExpansion) 1.05E-04 1.80E-D4 3.70E-04 5.609-04 7.80E-94 1.10E-03 1.75E-03 ftlf9--F g

0.3048 9.906 9.806 9.806 9.806 9.806 9.806 9.806 ms*

(Gravitation.al Constant) 32.17 32.17 32,17 32,17 32.17 32.17 32.17 ifa=

............... *4..................-..

f......* * ' ' "*o :-i

.~ *6*...

JU

~~~~1.4581 9.968-04 6.9*E*0...07ý-04 1:93-04

.1'.3-3-9E',D4 1.,0'40-0--D4 8.....S'6*2"E-0"5 O

kzg/rn-s (Oynemic Visonaity) 6.690-04 4.58E-04 2.00E-14 1.30E-04 9.30E-05 7.00E-05 5

5.7-K-05 lbrrdft-Pr 6.980 4.510 1.910 1.220 0.950 0.859 1.070 (Prandtl Number)

Calculated Parameter Formula 70 100 200 300 400 500 600 Reynolds Number, Re P.VO/M 0

0 0

Grashof Number, Gr g5TL'tWOp) 2441754517 1.2897E+10 1417E011 1.252E+12 3.9766F+12 1.1134E÷13 2.16049E+13 GrashofNmumber, Grm g:PAT(r 0-r,(jlpr 3,05E+08 1.59E009 3.02E+10 1.578+11 4.978+11 1.29E012.

2.70E+12 P.ayleigh Number, Ra GrPr 17043446531 5.72659+10 4.616E÷11 1.528E,12 3.7777E+12 8,8535+12 2.31172E+13 Rayleigh Number, Re.

GrPr 2.13E÷09 7.16f009 5.7-E+10 1.91E.11 4.720+11 1.11E+12 2,89F+12 Frmm (19):

Insids Surface ANatural Convection Heat Transfer Coefficient:..

Case:

Enclosed cylinder C =.n

=

(

pa

e.ij Hý,

C(GrPr)tlrL 181.&5 258.65 468.34 637,89 773.57 675.17 933.22 VnT_-

C 32.03 45.55 62.60 1 ý'12.14 136.24 154.13 164.35 Btu/hr-fF-~FI File No.: 0801038.304 Revision: 1 Page 15 of 20 F0306-01'

Structural Integrity Associates, Inc.

Table 7: 0% Flow Region 5 Heat Transfer Coefficient Pipe Inside Diameter, D =

inches=

3.333 ft

=

1.016 m

Outer Pipe. Inside radius, r. =

20 inches =

1.687 ft 0,508 m

Inner Pipe Outside Diameter, D=

Dinctes=

0.000

='0.000 m

Inner Pipe, Outside radius, r:

0 in ches =

0.000 ft 0.000 m

Fluid Velocity V =

0.000 Itsec =

gp(n=

0 14lbfhr CharacteristicLength, L= D =

3,333 f=

1.016 m

Tý - T.

LT= assurned to be 12%/a of fluid temperature=

8.40 12.00 24.00 36.00r 48.00 60.00 4.67 6.67 13.33 20.00 26.67 33.33 72.00 TF 40.00

.C Value at Fluid Temperature, T f121 Units Conversion 70 100 200 300 400 500 800 F

Water Property Factor [19]

24.11 37,78 93.33 14B.89 204.44 260400 3M6.56 k

1.7307 0.05997 0.6300 0.6784 0.6836 0.6611 0.6040 0.5071 Wmln-'C (Thermal Conductivity) 0.3-65 D.2-4-0 0.1,020 0,3950, 0.3820 0,349D 0.29311 Btuthr-f.t-'F er..c..%..::. )"...........................................o.... s.

...........o.3.e..

...........o:* ?.o........ *.3..s.............*............. :.............

.o r..-..F.

C, 4.189 4.185 4.179 4.229 4.313 4.522 4.982 6.322 Jl/fg-'C (Specific Heat) 1.000 0.098 1.010 1.030 1.080 1.100 1.510 Bttr'ibn--F

...s

ýcp

.*.e....)............... I...........................

...: q......

9.......... -...

-o 4.....i ----------

9..... * **

.. L...........t., '...

p 16.018 097.1 99.7 0,"2.7 017.0

-50.6 704.9 879.2 IgZrr (Density) 62.3.

62,1 60.1 57.3 53.6 49.0 42.4 bhrtvw A:

1.8 1.89E-04 3.24E-04 6.6E-04 1.01E-03 1.40E-03 1.96E-03 3.10E-03 m*r.m-C (Volumutric Rate ofl2xppsnian) 1.05E-04 1.80-04 0.702-04 5.600-04.

7.80E-04

1. 1 DEý(3 1,70.0-03 ft lf*-'F

..!.?:.

.o........

U............ *....-

7.:............ ! !. -.

o....

9 0.3048.

9.806 9.806 9.808 9.806 9.80*

06 0.06 0.00 nffs (Gravitational Constant) 32.17 32.17 32.17 32.17 32.17 32.17 32.17 fue n...........*

- *...........:.......................o.

......:*F.............

M1.4881 9.960-04 6.82E-04 3.117E-04 1.932-G4 1.39E-04 1.04E-04 8.02E0-kglm-s Pr 6.080 4.510 1.910 1.220 0.050 0.859 1.070 (Prandtl Nlumber)

Calculated Parameler Formula 70 100 200 300 400 500 600 Reynold's Number, Re pVDIA 0

0 0

Greshiftumbesr, Gr g0ATLl21);

909811606 4.7319E010 9.006E+11 4.667E012 1.4819f+13 3.854F-13 8.05143E+13 Grashaf Number, Gru gAT(r*r?1u*tdp)'

1.14E+019 5.91E+00 1.13E011 5.83E+11 1.852412 4.020+12 1.01E+13 Rsyleigh Number, Re GrPr 63515281008 2.1341E÷11 1.72E012 5.693,E+12 1.4076E+13 3.31E013 8.61503E+13 Rayleigh Number, Re Gr,,Pr 7.94E009 2.67+1t0 2.1SE011 7.12.E+1 1.76E+12 4.14E012 1.08E+13.

F. om [19]:

inside Surface Natural Convection Heat Transfer Coefficienrt:

Case:

Enclosed cylinder C=

a.'

n Mt-C(GrPrLIdL 162.97 231.79 420.60 571.68 603.25 784.30 368.32 W/ni-,C 25.70 40.82 74.07

1 00 6F 122.09 130.13 147.20 Btuthrrftn-F File No.: 0801038.304 Revision: 1 Page 16 of 20 F0306-01

Structural Integrity Associates, Inc.

UP tMz Z

Figure 2: Coordinate System for Forces and Moments SST CLAD 103.0,0 R' MAN LOW ALLOY STEEL NOZZLE SA 508 CL It 2.50 R NOTE:

REFERENCE 14 Figure 3: RO Nozzle and Safe End Geometry [20]

File No.: 0801038.304 Revision: 1 Page 17 of 20 F0306-01

jStructural Integrity Associates, Inc.

5.0 RESULTS OF ANALYSIS This calculation package specifies the ASME Code Edition, finite element model, thermal and pressure transients (Table 1), and HTCs (Table 2) to be used in a fatigue usage calculation of the RO nozzle at Vermont Yankee. Thermal transient and pressure stress components will be calculated using ANSYS [6]

and will be combined with piping loads in subsequent calculations.

Linearized stress components will be used for the fatigue usage calculation. For the nozzle blend radius location, the stresses used in the evaluation will be for the base metal only; that is, the cladding material will be unselected prior to stress extraction consistent with ASME Code rules and Reference [13].

The fatigue usage calculation will consider all six stress components, and will be performed using the rules of Subarticle NB-3200 of Section III of the ASME Code [1]. Calculated fatigue usage factors will be multiplied by the appropriate environmental fatigue multipliers computed for each location.

The results of this calculation are to be used in SIA calculations: No. 081038.305, Stress Analysis of Reactor Recirculation Outlet Nozzle and No. 081038.306, Fatigue Analysis of Recirculation Outlet Nozzle File No.: 0801038.304 Revision: 1 Page 18 of 20 F0306-01.

Structural Integrity Associates, Inc.

6.0 REFERENCES

1. American Society of Mechanical Engineers, Boiler and Pressure Vessel Code,Section III, Subsection NB, 1998 Edition, with Addenda through year 2000.

2. NUREG/CR-6583 (ANL-97/18), "Effects of LWR Coolant Environments on Fatigue Design Curves of Carbon and Low-Alloy Steels," March 1998.

3. J. P. Holman, 'Heat Transfer," 5th Edition, McGraw Hill Inc., 1981.

4. Structural Integrity Associates Calculation No. VY-16Q-305, Revision 0, "Recirculation Outlet Stress History Development for Nozzle Green Function."

5. CB&I, RPV Stress Report Sections S9 "Stress Analysis Recirculation Outlet Nozzle Vermont Yankee Reactor Vessel." and T9 "Thermal Analysis Recirculation Outlet Nozzle Vermont Yankee Reactor Vessel." CB&I Contract 9-6201, SI File No. VY-16Q-204.

6. ANSYS, Release 8.1 (w/Service Pack 1), ANSYS, Inc., June 2004.

7. Structural Integrity Associates Calculation No. VY-16Q-304, Revision 0, "Recirculation Outlet Nozzle Finite Element Model."

8. Entergy Design Input Record (DIR), Rev. 1, EC No. 1773, Rev. 0, "Environmental Fatigue Analysis for Vermont Yankee Nuclear Power Station," 7/26/07, SI File No. VY-16Q-209.

9. GE Drawing No. 919D294, Revision 11, Sheet 7, "Reactor Vessel, Spec. Control," SI File No. VY-05Q-241.

10. American Society of Mechanical Engineers, Boiler and Pressure Vessel Code,Section II, Part D, 1998 Edition, with Addenda through year 2000.

11. Structural Integrity Associates Calculation No. VY-16Q-306, Revision 0, "Fatigue Analysis of Recirculation Outlet Nozzle."

12. N. P. Cheremisinoff, "Heat Transfer Pocket Handbook," Gulf Publishing Co., 1984.

13. NUREG/CR-6260 (INEL-95/0045), "Application of NUREG/CR-5999 Interim Fatigue Curves to Selected Nuclear Power Plant Components," March 1995.

14. Warren C. Young, "Roark's Formulas for Stress & Strain," Sixth Edition, McGraw - Hill Book Company, 1989.

15. NUREG/CR-5704, "Effects of LWR Coolant Environments on Fatigue Design Curves of Austenitic Stainless Steels," April, 1999.

File No.: 0801038.304 Page 19 of 20 Revision: 1 F0306-01

Structural Integrity Associates, Inc.

16. SI Calculation No. VY-19Q-301, Revision 0, "Design Inputs and Methodology for ASME Code Confirmatory Fatigue Usage Analysis of Reactor Feedwater Nozzle."

17. SI Calculation No. VY-19Q-302, Revision 0, "ASME Code Confirmatory Fatigue Evaluation of Reactor Feedwater Nozzle."

18. SI Calculation No. VY-19Q-303, Revision 0, "Feedwater Nozzle Environmental Fatigue Evaluation."

19. J. P. Holman, "Heat Transfer," 4th Edition, McGraw Hill Inc., 1976.

20. GE Stress Report 23A4316, Revision 0, "Stress Report-Reactor Vessel Recirculation Outlet Safe End," SI File No. VY-16-204.

File No.: 0801038.304 Revision: 1 Page 20 of 20 F0306-01:

Structural Integrity Associates, Inc.

APPENDIX A:

ANSYS Input File: RONVY.INP File No.: 0801038.304 Revision: 1 Page A-1 of A-23 F0306-01*

Structural Integrity Associates, Inc.

ANSYS Input File: RONVY.INP finish

/clear, start

/prep7

/title, Recirc Outlet Nozzle Finite Element Model

/com,

PLANE42, 2-D Solid et, 1, PLANE42,,, 1

!Axisymmetric

/com,

/com, Material Properties

/com,

MPTEMP,

, 70,200,300,400,500,600 tmp = 3600*12 hr-ft to sec-in

/COM, Material #1 Safe-End and Portion of Piping (SA-182 F316)

(18Cr-8Ni)

MPDATA,EX

,1,, 28.3e6, 27.6e6, 27.0e6, 26.5e6, 25.8e6, 25.3e6 MPDATA,ALPX,1,

, 8.5e-6, 8.9e-6, 9.2e-6, 9.5e-6, 9.7e-6, 9.8e-6

MPDATA, KXX,1,

, 8.6/tmp, 9.3/tmp, 9.8/tmp, 10.4/tmp, 10.9/tmp, ll.3/tmp

MPDATA, C,l,

, 0.116, 0.122, 0.125, 0.129, 0.131, 0.133 mp, nuxy, 1, 0.3 mp,dens,1,0.283

/COM, Material #2 (Nozzle Forging) SA-508 Class 2 (3/4Ni-l/2Mo-l/3Cr-V)

MPDATA,EX

,2,

, 27.8e6, 27.1e6, 26.7e6, 26.1e6, 25.7e6, 25.2e6 MPDATA,ALPX,2,

, 6.4e-6, 6.7e-6, 6.9e-6, 7.le-6, 7.3e-6, 7.4e-6

MPDATA, KXX,2,

, 23.5/tmp, 23.6/tmp, 23.4/tmp, 23.1/tmp, 22.7/tmp, 22.2/tmp

MPDATA, C,2,

, 0.105, 0.114, 0.119, 0.125, 0.130, 0.135 mp, nuxy, 2,0.3 mp,dens, 2,0.283

/COM, Material #3 (Cladding) SA-240 Type 304 (18Cr-8Ni)

MPDATA,EX

,3,

, 28.3e6, 27.6e6, 27.0e6, 26.5e6, 25.8e6, 25.3e6 MPDATA,ALPX,3,

, 8.5e-6, 8.9e-6, 9.2e-6, 9.5e-6, 9.7e-6, 9.8e-6

MPDATA, KXX,3,

, 8.6/tmp, 9.3/tmp, 9.8/tmp, 10.4/tmp, 10.9/tmp, ll.3/tmp

MPDATA, C,3,

, 0.116, 0.122, 0.125, 0.129, 0.131, 0.133 mp, nuxy, 3, 0. 3 File No.: 0801038.304 Page A-2 of A-23 Revision: 1 F0306-01

Structural Integrity Associates, Inc.

mp, dens, 3,0.283

/COM, Material #4 (Vessel) SA-533, GR.

B (Mn-i/2Mo-i/2Ni)

MPDATA,EX

,4,

, 29.2e6, 28.5e6, 28.0e6, 27.4e6, 27.0e6, 26.4e6 MPDATA,ALPX,4,

, 7.0e-6, 7.3e-6, 7.4e-6, 7.6e-6, 7.7e-6, 7.8e-6

MPDATA, KXX,4,

, 23.5/tmp, 23.6/tmp, 23.4/tmp, 23.1/tmp, 22.7/tmp, 22.2/tmp

MPDATA, C,4,

, 0.105, 0.114, 0.119, 0.125, 0.130, 0.135 mp, nuxy, 4, 0.3 mp, dens, 4, 0. 283

AFUN, DEG

/com, Geometric Parameters *

set,vira, (103+3/16)

!Actual Vessel Inner Radius to base metal used for model

set,vir, 2.0*vira

!2.0 time of Vessel Inner Radius to base metal used for model

set,tvw,5+5/8-3/16

!Vessel Wall Thickness

set, ril, 25.75/2

set, rol, 28.375/2

set,L1,5

set, ro2,28.375/2

set, L2,4.25

set, ro3,28.875/2

set, ro4,48.75/2

set, L3, 1.5

set,L4,5.25

set, L5, 7+1/16

set, L6, 12+13/16

setL7,9+7/8

set,L8,9+3/8

set, L9, 31+15/16

set, L0, L9-12-13/16-tvw

set, ra, 7

set,rb,1

set, rc, 5.25

set, rd, 2.5

set, tv, 3/16

set,dimA,vir-(tv*2.0)+L9+11+Ll

!Vessel Centerline to End of Safe End used for model

set,L21,1

set,L22,4.25

set,ri2l, (25+15/16)/2

/com, Geometry File No.: 0801038.304 Page A-3 of A-23 Revision: 1 F0306-01

0 Structural. Integrity Associates, Inc.

local, 13,0,, dimA,,,,

csys, 13

/com, Begin at end of Safe-End Carbon Section k,

9, ril,

-l*(dimA) k, 2,

ril+tv, -1*(dimA) k, 3,

rol, -l*(dimA) k, 4, ril,

-i* (dimA-Ll) k, 5,

ril+tv,

-l*(dimA-Ll) k, 6,

rol, -i*(dimA-Ll) k, 7,

ril,

-T*(dimA-Ly-L2) k, 8,

ril+tv, -1*(dimA-L5-L2) k, 9,

ro2, -r*(dimA-Li-L2) k, 10, ril,

-I*.(dimA-LI-L2-L3) k, 2 iv,r

-*(dimA-Lt-L2-L3) k, 12, ro3, -r*(dimA-Lb-L2-L3) k, 13,

ril,

-!*(dimA-LI-L2-L3-L4) k, 14, ril+tv,

-1*(dimA-LI-L2-L3-L4) k, 15, ro3, -1*(dimA-Lt-L2-L3-L4) k, 16,

ril,

-1*(dimA-Lt-L2-L3-L4-L5) k, 17, ril+tv,

-I*(dimA-Ln-L2-L3-L4-L5) k, 18,

ro3,

-1*(dimA-Lt-L2-L3-L4-L5) k,19,

ro4,

-0*(dimA-LI-L2-L3-L4-LS-L7)! Temporary Point 1,19,18 1,18,15 if illt, 1, 2, ra k,22, ro4+(L8+6)*tan(15),

-I*(dimA-LI-L2-L3-L4-L5-L7-(L8+6))

1,19,22 LFILLT, 1, 4, rb k,

25,

ril,

-I*(dimA-LI-L2-L3-L4-L6) k, 26, ril+tv, -I*(dimA-LI-L2-L3-L4-L6) k, 27, ril+(Ll0+tvw+tv+4)*tan(15),

-1*(vir-tv-4) k, 28, ril+tv+(Ll0+tvw+tv+4)*tan(15),

-l*(vir-tv-4) k,29, (vir+tvw+tv)*sin(45),

-1*(vir+tvw+tv)*cos(45) k,30, 0,

-l*(vir+tvw+tv)

!Temporary Point k,31, 0,

0 ! Temporary Point larc, 29, 30,31, vir+tvw+tv k, 32, (vir+tv)*sin(45),

-l* (vir+tv) *cos (45)

FileNo.: 0801038.304 Page A-4 of A-23 Revision: 1 F0306-01:

V Structural Integrity Associates, Inc.

k,33, 0,

-1*(vir+tv)

Temporary Point larc, 32,33, 31, vir+tv k, 34, k, 35, larc, 34

LSTR, LSTR,

LSTR, LSTR, vir*sin(45), -1*vir*cos(45) 0,

-1*vir

,35, 31,vir 4,

5, 6,

9, 12, 5,

4, 7,

8, 11, 10, 13, 14, 16, 17, 26, 25, 4,

1, 2,

3, 5,

7, 8,

12, 11, 13, 14,

! Temporary Point 5

6 9

12 15 8

7 10 11 14 13 16 17 25 26 28 27 1

2 3

6 2

8 9

11 10 14 15 FLST,2,2,4,ORDE,2 FITEM, 2,4 FITEM, 2, 6 LPTN, P51X FLST,2,2,4,ORDE, 2 FITEM, 2,8 FITEM, 2,25 LPTN, P51X FLST,2,2,4,ORDE,2 File No.: 0801038.304 Revision: 1 Page A-5 of A-23 F0306-01.

V Structural Integrity Associates, inc.

FITEM, 2,27 FITEM, 2,24 LPTN, P51X FLST,2, 6,4,ORDE, 6 FITEM, 2, 6 FITEM, 2,25 FITEM, 2,37 FITEM, 2,40 FITEM, 2,42 FITEM, 2,44 LDELE,P51X,

.1 1*

LFILLT,4,41,rd, LFILLT, 43, 8,rd,,

1*

LFILLT,39,38,rc, FLST, 2,3, 4,ORDE, 3 FITEM, 2,1 FITEM, 2,3 FITEM, 2,5 LCOMB,P51X,

,0

LSTR, 16, 17

LSTR, 17, 21

LSTR, 25, 26

LSTR, 26, 24

LSTR, 22, 30

LSTR, 30, 35

LSTR, 27, 28

LSTR, 28, 33

LSTR, 29, 32

LSTR, 32, 34 k,39, 0,

-1*(vir+tvw+tv)

!Create Areas FLST, 2,4,4 FITEM, 2,27 FITEM, 2,30 FITEM, 2,26 FITEM, 2,9 AL, P51X FLST,2,4,4 FITEM, 2,28 FileNo.: 0801038.304 Page A-6 of A-23 Revision: 1 F0306-01:

V Structural Integrity Associates, Inc.

FITEM, 2,29 FITEM, 2, 30 FITEM, 2,30 AL, P51X FLST, 2, 4, 4 FITEM, 2, 11 FITEM, 2,32 FITEM, 2,10 FITEM, 2, 14 AL, P51X FLST, 2, 4,4 FITEM, 2,15 FITEM, 2, 14 FITEM, 2, 9 FITEM, 2, 31 AL, P51X FLST, 2, 4,4 FITEM, 2,32 FITEM, 2,33 FITEM, 2, 12 FITEM, 2,17 AL, P51X FLST, 2,4,4 FITEM, 2,16 FITEM, 2, 17 FITEM, 2, 31 FITEM, 2,34 AL, P51X FLST, 2,4,4 FITEM, 2,36 FITEM, 2, 13 FITEM, 2,33 FITEM, 2, 18 AL, P51X FLST, 2,4,4 FITEM, 2, 19 FITEM, 2, 18 FITEM, 2,35 FITEM, 2,34 AL, P51X FLST, 2,4,4 FITEM, 2,2 FITEM, 2,5 FITEM, 2,36 FITEM, 2,21 AL, P51X FLST, 2, 4,4 File No.: 0801038.304 Page A-7 of A-23 Revision: 1 F0306-0K

Structural integrity Associates, Inc.

FITEM, 2,20 FITEM, 2, 21 FITEM, 2, 3 FITEM, 2,35 AL, P51X FLST, 2, 4, 4 FITEM, 2, 1 FITEM, 2,37 FITEM, 2,23 FITEM, 2,5 AL, P51X FLST, 2,4,4 FITEM,2,22 FITEM, 2,23 FITEM, 2,25 FITEM, 2, 3 AL, P51X FLST, 2,4,4 FITEM, 2,38 FITEM, 2, 42 FITEM, 2,37 FITEM, 2,8 AL, P51X FLST, 2,4,4 FITEM, 2,4 FITEM, 2, 8 FITEM, 2,25 FITEM, 2,40 AL, P51X FLST, 2,4,4 FITEM, 2,24 FITEM, 2,45 FITEM, 2, 7 FITEM, 2,42 AL, P51X FLST, 2,4,4 FITEM, 2, 6 FITEM, 2, 7 FITEM, 2,44 FITEM, 2,40 AL, P51X FLST, 2,4,4 FITEM, 2, 41 FITEM,2, 43 FITEM,2,47 FITEM, 2,44 AL, P51X File No.: 0801038.304 Page A-8 of A-23 Revision: 1 F0306-O1

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FLST, 2, 4, 4 FITEM, 2,39 FITEM, 2,46 FITEM, 2,45 FITEM, 2,43 AL, P51X

! define materials FLST, 5,8,5,ORDE, 2 FITEM, 5,1 FITEM, 5, -8 CM, Y,AREA

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ASEL,

,P51X File No.: 0801038.304 Page A-9 of A-23 Revision: I F0306-01

Structural Integrity Associates, Inc.

CM, _Y1,AREA CMSEL, S, _Y CMSEL,S, _YI

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File No.: 0801038.304 Page A-10 of A-23 Revision: I F0306-01:

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FLST, 5,3, 4, ORDE, 3 FITEM, 5,39 FITEM, 5, 41 FITEM, 5,43 CM, _Y, LINE

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FLST, 5, 6, 4, ORDE, 6 FITEM, 5,2 FITEM, 5,20 File No.: 0801038.304 Page A-I Iof A-23 Revision: 1 F0306-01:

VStructural Integrity Associates, Inc.

FITEM, 5,-21 FITEM, 5,26 FITEM, 5,29 FITEM, 5, -30 CM, _Y,LINE

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!Modify the safe end ID FLST, 2, 6, 5, ORDE, 2 FITEM, 2,1 FITEM, 2, -6 ACLEAR, PSIX FLST, 2,6, 5, ORDE, 2 File No.: 0801038.304 Page A-12 of A-23 Revision: 1 F0306-01:

Structural Integrity Associates, Inc.

FITEM, 2, 1 FITEM, 2, -6 ADELE, P51X FLST, 2,9, 4,ORDE, 7 FITEM, 2,9 FITEM, 2,14 FITEM, 2, -17 FITEM, 2,26 FITEM, 2, -27 FITEM, 2,30 FITEM, 2, -31 LDELE,P51X, 1

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LDELE, 16,

,I FLST, 2,4,3 FITEM, 2,11 FITEM, 2,23 FITEM, 2,41 FileNo.: 0801038.304 Page A-13 of A-23 Revision: 1 F0306-01

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FITEM, 2, 12 A, P51X FLST, 2,4,3 FITEM, 2,23 FITEM, 2, 8 FITEM, 2, 9 FITEM, 2, 41 A, P51X FLST, 2,4,3 FITEM, 2,8 FITEM, 2, 7 FITEM, 2, 6 FITEM, 2, 9 A, P51X FLST, 2,4,3 FITEM, 2,7 FITEM, 2,5 FITEM, 2,3 FITEM, 2, 6 A, P51X FLST, 2,4,3 FITEM, 2, 10 FITEM, 2,20 FITEM, 2,23 FITEM, 2, 11 A, P51X FLST, 2,4,3 FITEM, 2,20 FITEM, 2, 4 FITEM, 2, 8 FITEM, 2,23 A, P51X FLST, 2,4,3 FITEM, 2, 4 FITEM, 2, 2 FITEM, 2,7 FITEM, 2, 8 A, P51X FLST, 2,4,3 FITEM, 2,2 FITEM, 2, 1 FITEM, 2, 5 FITEM, 2, 7 A, P51X FLST, 5,8,5,ORDE, 4 FITEM, 5,1 FITEM, 5, -6 File No.: 0801038.304 Page A-14 of A-23 Revision: 1 F0306-01I

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FITEM, 5, 19 FITEM, 5,-20 CM, _Y, AREA

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,P51X CM, _Y1,AREA CMSEL,S, _Y 1*

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LSEL,

,,P51X CM,__YI,LINE File No.: 0801038.304 Page A-15 of A-23 Revision: 1 F0306-01

Structural Integrity Associates, Inc.

CMSEL,, _Y 1*

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FLST, 2,2,5, ORDE, 2 File No.: 0801038.304 Page A-16 of A-23 Revision: 1 F0306-01*

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FITEM, 2,17 FITEM, 2, -18 ACLEAR, P51X csys,0 k, 51,62/2,0,0 k, 52,62/2,60,0

LSTR, 51, 52 FLST, 2,2, 5, ORDE, 2 FITEM, 2,17 FITEM, 2, -18 ADELE, P51X

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.1 FLST, 2,4,4 FITEM, 2,54 FITEM,2, 62 FITEM,2,55 FITEM, 2,44 AL, P51X FLST, 2,4,4 FITEM, 2,55 FITEM, 2,63 FITEM, 2,58 FITEM, 2,45 AL, P51X FLST,2,4,4 FITEM, 2, 63 FITEM, 2,56 FITEM, 2,57 FITEM, 2., 46 AL, P51X FLST, 2,4,4 FITEM, 2,47 FITEM, 2,59 FITEM, 2, 57 FITEM, 2,62 AL, P51X File No.: 0801038.304 Page A-17 of A-23 Revision: I F0306-01'

Structural Integrity Associates, Inc.

CM, _Y, AREA

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,f f

File No.: 0801038.304 Page A-18 of A-23 Revision: 1 F0306-01'

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1*

FLST, 5,3,4,ORDE, 3 FITEM, 5,56 FITEM, 5, -57 FITEM, 5,59 CM, _Y,LINE

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CMDELE, Y

CMDELE, _Y1 CMDELE,_Y2 File No.: 0801038.304 Page A-19 of A-23 Revision: 1 F0306-01Y

Structural Integrity Associates, Inc.

!Simulating Butter FLST, 2,2,5, ORDE, 2 FITEM, 2,9 FITEM, 2, -10 ACLEAR, P51X FLST, 2,2,5, ORDE, 2 FITEM, 2,9 FITEM, 2, -10 ADELE, P51X KGEN,2,15,

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FITEM, 2,68 FITEM, 2, 3 AL, P51X FLST, 2,4,4 FITEM, 2,39 FITEM, 2, 5 FITEM, 2, 2 FITEM, 2,53 AL, P51X FLST, 2,4,4 FITEM, 2,20 FITEM, 2, 60 FITEM, 2,53 FITEM, 2,41 AL, P51X FLST, 2,4,4 FITEM, 2,72 FITEM, 2,68 FITEM, 2,69 FITEM, 2,41 AL, P51X FLST, 2,4,4 FITEM, 2,21 FITEM, 2,60 FITEM, 2,36 FITEM, 2,43 AL, P51X FLST, 2,4,4 FITEM,2,66 FITEM, 2,69 FITEM, 2,35 FITEM, 2,43 AL, PSIX CM, _Y, AREA

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CMDELE, Y1 FLST, 5,3,5,ORDE,3 FITEM, 5,9 File No.: 0801038.304 Page A-21 of A-23 Revision: 1 F0306-O1

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FITEM, 5,23 FITEM, 5,-24 CM, _Y, AREA

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,P51X CM, _Y1,AREA CMSEL,S, _Y 1*

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LSEL,

,P51X File No.: 0801038.304 Page A-22 of A-23 Revision: 1 F0306-01

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CM, _Y1,LINE CMSEL,, _Y

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save finish FileNo.: 0801038.304 Page A-23 of A-23 Revision: 1 F0306-01I Structural Integrity Associates, Inc.

File No.: 0801038.305 CALCULATION PACKAGE Project No.: 0801038 Quality Program Z Nuclear FD Commercial PROJECT NAME:

VY Confirmatory Analysis for the CS and RO Nozzles CONTRACT NO.:

10163217 Amendment 5 CLIENT:

PLANT:

Entergy Nuclear Operations, Inc Vermont Yankee Nuclear Power Station CALCULATION TITLE:

Stress Analysis of Reactor Recirculation Outlet Nozzle Document Affected Project Manager Preparer(s) &

Revision Pages Revision Description

. Approval Checker(s)

Signature & Date Signatures & Date 0

1 - 16 Initial issue.

Preparer:

Gary L. Stevens Computer Files 01/07/09 Tyler Novotny 01/07/09 Checker:

R. D. Dixon 01/07/09 1

1-9,11,15 Revised per summary

/

Preparer:

contained in Section 1.1.

Zm-Changes are marked with revision bars" in right-Gary L. Stevens hand margin.

03/09/09 Tyler D. Novotny 03/09/09 Checker:

Tim D. Gilman 03/09/09 Page 1 of 16 F0306-01RO

Structural Integrity Associates, Inc.

Table of Contents 1.0 O B JE C T IV E.................................................................................................................................

3 1.1 Changes M ade in Revision 1 of this Calculation...........................................................

3 2.0 M ETHODOLOGY.......................................................................................................................

3 3.0 ASSUM PTIONS / DESIGN INPUTS....................................................................................

4 4.0 CALCULATIONS........................................................................................................................

4 4.1 Finite Element Unit Pressure Stress Analysis..............................................................

4 4.2 Thermal Transient Stress Analysis.................................................................................

4 4.3 Determining Critical Stress Paths....................................................................................

5 4.4 Stress Calculation.......................................................................................................

6 4.5 P ip in g L oads........................................................................................................................

7 5.0 RESULTS OF ANALYSIS.....................................................................................................

8 6.0 RE FERENCES.............................................................................................................................

9 List of Tables Table 1: Pressure Stress Intensity Results (1,000 psi)...................................................................

7 Table 2: Stresses Under Unit Pressure Load, psi...........................................................................

10 Table 3: Example Thermal Stress Result Output, psi....................................................................

11 List of Figures Figure 1. RO Nozzle Internal Pressure Distribution......................................................................

12 Figure 2. RO Nozzle Pressure Cap Load & Boundary Condition..................................................

13 Figure 3. RO Nozzle Vessel W all Boundary Condition...............................................................

14 Figure 4. Safe End Critical Thermal Stress Intensity Location............................ I.......................

15 Figure 5. Nozzle Blend Radius Limiting Pressure Stress Intensity Location................................

16 Figure 6. Limiting Stress Paths.....................................................................................................

16 FileNo.: 0801038.305 Page 2 of 16 Revision: 1 F0306-O1RO

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1.0 OBJECTIVE The objective of this calculation package is to obtain stress distributions for the reactor pressure vessel (RPV) recirculation outlet (RO) nozzle at the Vermont Yankee Nuclear Power Station.

ANSYS [ 1 ] thermal transient and pressure stress analyses are performed, along with calculation of stresses due to attached piping loads. The stress results will be used for a subsequent ASME Code,Section III NB-3200 [2] fatigue usage calculation.

1.1 Changes Made in Revision 1 of this Calculation Description of changes made in Revision 1 of this calculation:

a. Transient 9 described in Section 4.3 was changed to more precisely match the Green's Function analysis. This also required modification of the input files VYRONTRAN9-T.INP and VY RON TRAN9-S.INP.

b. The input files VY_RONTRAN2-TINP and VYRON_TRAN2-S.INP were modified to include a finer time step around 601 seconds.

c. A Kt value of 1.53 that was conservatively applied to piping loads at blend radius was changed to Kt = 1.0 to match the Green's Function analysis.

d. Table 3 was revised because the input file VY_RONTRAN4-T.INP was updated to correct a conservative misapplication of a temperature ramp rate.

e. Figure 4 was revised because Transient 9, which produced Figure 4, was modified.

f. All remaining changes marked throughout this calculation are editorial changes made to the text of the calculation package.

2.0 METHODOLOGY The methodology to be used for this evaluation was established in a previous calculation package

[3]. A previously developed finite element model (FEM) [3] of the RO nozzle is used to perform thermal and pressure stress analyses using ANSYS [1]. A thermal transient analysis is performed for each defined transient. Concurrent with the thermal transients are pressure and piping interface loads. For these loads, unit load analyses (based on finite element analysis for pressure and manual calculations for attached piping loads) are performed. All six components of the stress tensor are determined in the stress calculations.

The fatigue usage calculation and enviromnental fatigue usage analysis will be performed in a separate calculation package. That subsequent calculation will utilize the thermal and pressure stresses determined in this calculation, along with stresses due to attached piping loads provided in Tables 4 and 5 of Reference [3]. The stresses due to pressure and the attached piping loads will be scaled based on the temperature and pressure magnitudes during each individual transient, and the location being analyzed. The appropriate nozzle blend radius effects factor will also be applied to the total stresses for the nozzle blend radius location.

FileNo.: 0801038.305 Page 3 of 16 Revision: 1 F0306-01RO

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3.0 ASSUMPTIONS / DESIGN INPUTS Assumptions and design inputs were previously established in Section 3.0 of the Reference [3]

calculation. Assumption 3.1.3 of Reference [3] was verified in this calculation package by plotting the stress components of each transient in ANSYS. If the stress components plot did not contain a step change at the end of the transient, the steady state portion, the steady state time step assumed was detenrined to be adequate.

4.0 CALCULATIONS 4.1 Finite Element Unit Pressure Stress Analysis A uniform pressure of 1,000 psi was applied to the FEM along the inside surface of the RO nozzle and the RPV wall (Figure 1). A pressure load of 1,000 psi was used because it is easily scaled up or down to account for different pressures that occur during transients. In addition, a membrane stress "cap load" was applied to the modeled end of the piping attached to the RO nozzle safe end. This membrane stress was calculated as follows:

PDi2 Vcap -

2 D2 DO - Di where:

P = Pressure = 1,000 psi unit load Di= Inner Diameter at end of model = 25.9375 in Do= Outer Diameter at end of model = 28.375 in Therefore, the membrane stress is 5,082 psi. The calculated value is given a negative sign in order for it to exert tension on the piping end of the model. The FEM geometry input file is taken from the calculation that specifies the design and methodology inputs [3, input file RONVY1NP]. The ANSYS input file VY RONP.INP contains the pressure loading. Figure 1 shows the applied 1,000 psi internal pressure distribution. At the vessel wall, a symmetric boundary condition is applied. At the piping end of the model, axial displacement is coupled to simulate the effect of the attached piping that is not modeled. Figure 2 and Figure 3 show the boundary conditions.

4.2 Thermal Transient Stress Analysis The FEM geometry input file is taken from the calculation that specifies the design and methodology inputs [3, file RON_VY.INP], and is used as input to the files in which the thermal transient and pressure stress analyses are performed.

For the thermal transient ANSYS analyses, previously defined thermal transients [3, Table 1] are evaluated, applying heat transfer coefficients [3, Table 2], as appropriate, based on the flow rates for each individual transient.

Each thermal transient is evaluated in ANSYS to determine the resulting temperature distributions.

The thermal results are used as input for the stress analysis for each transient. The boundary FileNo.: 0801038.305 Page 4 of 16 Revision: 1 F0306-01RO

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conditions used for the pressure load case were also applied to the thermal stress cases. Figure 2 and Figure 3 show the application of these boundary conditions.

All ANSYS input files for the thermal analyses, as listed below, are saved in the project computer files:

RON_VYINP: Geometry and material properties VYRONTRAN1-T1NP, VY RON TRANJ-S.INP: Transient 1, thermal and stress analyses VY RONTRAN2-T.NP, VYRONTRAN2-S.INP: Transient 2, thermal and stress analyses VYRONTRAN3-TINP, VYRONTRAN3-S.INP: Transient 3, thermal and stress analyses VY RONTRAN4-T.INP, VY RON TRAN4-S.INP: Transient 4, thermal and stress analyses VYRON TRAN5-TJNP, VYRONTRANS-S.INP: Transient 5, thermal and stress analyses VYRONTRAN6-TINP, VYRONTRAN6-S.INP: Transient 6, thermal and stress analyses VYRONTRAN7-T.INP, VY RON TRAN7-S.INP: Transient 7, thermal and stress analyses VY RON TRAN8-TINP, VYRONTRAN8-S.INP: Transient 8, thermal and stress analyses VYRONTRAN9-TINP, VYRONTRAN9-S.INP: Transient 9, thermal and stress analyses VY_RONTRANJO-T.INP, VY RON TRANJO-S.INP: Transient 10, thermal and stress analyses VYRONTRANJJ-T.INP, VYRONTRANJJ-S.INP: Transient 11, thermal and stress, analyses VYRONTRAN12-T.INP, VYRONTRAN12-S.INP: Transient 12, thermal and stress analyses 4.3 Determining Critical Stress Paths The thermal transient that is to be used in determining the critical stress path at the safe end was determined by the most severe temperature difference over the shortest amount of time. This transient, Transient 9, is intended to represent the worst case thermal transient. This occurs during the Improper Startup cycle per Reference [3, Table 1]. The thermal transient conditions are:

12% flow rate heat transfer coefficients.

" Thermal shock from 526°F to 130'F along the inside surface of the nozzle safe end and piping and a blend radius and lower vessel thermal shock from 5260F to 2680F.

" Constant temperatures from previous step for 26 seconds

" Thermal shock from 130°F to 526°F along the inside surface of the nozzle safe end and piping and a blend radius and lower vessel thermal shock from 268°F to 526°F.

" Steady state temperature conditions following thermal shocks.

" Constant temperature of 120'F on the outside surface of the model.

The ANSYS input files for the analysis, as listed below, are saved in the project computer files:

RON_VYINP. Geometry and material properties VYRONTRAN9-TINP, VY RON TRAN9-S.INP: Thermal and stress analysis for the worst case transient for the safe end An interactive review of the worst case thermal stress results (which are controlling for the safe end) showed the critical location in the model to be at Node 6395. The location of Node 6395 is shown in File No.: 0801038.305 Page 5 of 16 Revision: 1 F0306-O1RO

Structural Integrity Associates, Inc.

Figure 4. This location was selected since it possessed the highest stress intensity during the worst case thermal transient. This is the same location evaluated in Reference [4].

A critical stress location in the nozzle blend radius will also be analyzed. This location is chosen based upon the highest pressure stress (which is controlling in the nozzle blend radius) in the base metal. An interactive review of the pressure stress intensity results showed the critical location in the nozzle blend radius to be at Node 3829 (Figure 5). This is the same location evaluated in Reference [4].

Figure 6 shows the two critical stress paths that will be used to extract the linearized stresses at the safe end and nozzle blend radius.

4.4 Stress Calculation Linearized stresses from Node 6395 (safe end inside surface) and Node 3829 (nozzle blend radius inside surface of base metal) are used for the fatigue usage analysis, as shown in Figure 6. For the nozzle blend radius location, the stresses used are for the base metal only; since the cladding is of the integrally bonded type and is less than 10% of the total thickness of the section the material is unselected prior to stress extraction, per NB-3 122.3 [2].

The pressure stress intensities for the safe end and blend radius paths were extracted using the ANSYS file VYRONP.1NP. This produced one file, ROPRESSURE.lin, that contains results of the critical stress paths.

Table 1 shows the final pressure stress intensity results for the safe end and blend radius. The results at the blend radius are slightly different from those reported in Table 2 of Reference [4] as a result of the revised material properties (i.e., temperature dependent material properties were used in the current evaluation vs. constant material properties in Reference [4]).

Results were also extracted from the vessel portion of the model to verify the accuracy of the results obtained from the ANSYS model, and to check the results due to the use of the 2.0 multiplier on the vessel radius. These results are contained in the file RO PRESSURE. lin. The radius of the finite element model (FEM) was multiplied by a factor of 2.0 [4] to account for the fact that the vessel portion of the axisymmetric model is a sphere, but the true geometry is the intersection of two cylinders.

The equation for the membrane hoop stress in a thin wall sphere is:

0r =(pressure) x (radius)

S2 x thickness Considering an actual vessel base metal radius, R, of 105.906 inches increased by a factor of 2.0, a vessel base metal thickness, t, of 5.4375 inches, and an applied pressure, P, of 1,000 psi, the calculated stress for a thin wall sphere is PR/(2t) = 19,477 psi. This compares very well with the remote vessel wall membrane hoop stress from the ANSYS result file, ROPRESSURE.lin, of File No.: 0801038.305 Page 6 of 16 Revision: 1 F0306-O1RO

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18,070 psi. Thus, considering the peak total pressure stress of 31,270 psi, the stress concentrating effect of the nozzle blend radius is 31,270/19,477 = 1.61. In other words, the peak nozzle blend radius stress is 1.61 times higher than nominal vessel wall stress for the axisymmetric model.

The equation for the membrane hoop stress in a thin wall cylinder is:

(pressure) x (radius)

\\

thickness Based on the previous dimensions, the calculated stress for a cylinder without the 2.0 factor is 19,477 psi. Increasing this by a factor of 1.61 yields an expected peak nozzle blend radius stress of 31,358 psi, which would be expected from a cylindrical geometry that is representative of the nozzle configuration. Therefore, the result from the ANSYS file for the peak nozzle blend radius stress (31,270 psi) is close to the peak nozzle blend radius stress for a cylindrical geometry because of the use of the 2.0 multiplier. This is consistent with SI's experience where a factor of two increase in radius is typical for representing the 3-D effect in an axisymmetric model.

4.5 Piping Loads The piping loads were taken from Table 4 of Reference [3]. To determine the piping load stresses, the distances from the applied piping loads to the limiting stress locations were first determined. The limiting stress path locations from Section 4.3 are in the same locations assumed in Table 4 of Reference [3]; this means that no reconciliation of the lengths in Table 4 of Reference [3] is needed.

Reference [3, Section 4.1] methodology was used to calculate the piping load stresses. The piping loads and piping load stresses are found in Table 4 and Table 5 of Reference [3].

Table 1: Pressure Stress Intensity Results (1,000 psi)

Membrane plus Total Stress Location Bending Stress Intensity Intensity (psi)

(psi)

Safe End 11,350 11,490 (Path 1 Inside)

Blend Radius (Path 2 Inside) 30,540 31,270 File No.: 0801038.305 Revision: 1 Page 7 of 16 F0306-O1RO

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5.0 RESULTS OF ANALYSIS A thermal transient analysis for each defined transient, as well as unit pressure stress and piping interface load analyses were performed for the RO nozzle at Vermont Yankee. All six components of the stress tensor were extracted from the ANSYS model at the two limiting path locations, which are the same two locations previously evaluated [4]. Table 2 provides the unit (1,000 psig) pressure stress analysis results. The unit pressure load results are used to choose the location to analyze at the nozzle blend radius and will be scaled up or down based on applied pressures in the fatigue analysis.

Table 5 of Reference [3] provides the piping stresses at the two critical locations. Table 3 shows an example of thermal stress results. The remaining thermal stress results are contained in the ANSYS output files, listed below, which are saved in the project computer files:

ROPRESSURE. lin: Unit pressure stress analysis results VYRONTRAN1-S.lin: Transient 1, thermal stress analysis results VY RON TRAN2-S.lin: Transient 2, thermal stress analysis results VYRONTRAN3-S.lin: Transient 3, thermal stress analysis results VY RON TRAN4-S.lin: Transient 4, thermal stress analysis results VYRONTRAN5-S.lin: Transient 5, thermal stress analysis results VYRONTRAN6-S.lin: Transient 6, thermal stress analysis results VY RON TRAN7-S.lin: Transient 7, thermal stress analysis results VYRONTRAN8-S.lin: Transient 8, thermal stress analysis results VYRONTRAN9-S.lin: Transient 9, thermal stress analysis results VY RON TRANIO-S. lin: Transient 10, thermal stress analysis results VYRONTRANlI-S.lin: Transient 11, thermal stress analysis results VYRONTRAN12-S.lin: Transient 12, thermal stress analysis results A fatigue calculation using the methodology of Subarticle NB-3200 of Section III of the ASME Code [2] and an environmental fatigue usage analysis will be performed in a separate calculation package using the stress results from this calculation.

The results of this calculation are to be used in SI Calculation No. 081038.306, "Fatigue Analysis of Recirculation Outlet Nozzle."

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6.0 REFERENCES

1. ANSYS, Release 8.1 (w/Service Pack 1), ANSYS, Inc., June 2004.

2. ASME Boiler and Pressure Vessel Code,Section III, Subsection NB, 1998 Edition with 2000 Addenda.

3. SI Calculation No. 0801038.304, Revision 1, "Design Inputs and Methodology for ASME Code Confirmatory Fatigue Usage Analysis of Reactor Recirculation Outlet Nozzle."

4. SI Calculation No. VY-16Q-305, Revision 0, "Recirculation Outlet Stress History Development for Nozzle Green Function."

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Table 2: Stresses Under Unit Pressure Load, psi I

I Membrane plus Bending Total Node S"

SY Sz Sxy Syz Sxz Sx Sy Sz Sxy Syz Sxz SE 6395

-955.2 4420 10390 15.26 0

0

-955.2 4912 10530

-222.6 0

0 BR 3829

-718.7

-951.7 25000 4708 0

0

-718.7 206.2 30150 733.2 0

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Table 3: Example Thermal Stress Result Output, psi Time Membrane Plus Bending Total (s)

Sx Sy Sz Sxy Syz Sxz Sx Sy Sz Sxy Syz Sxz 0

-33

-3379 196 351 0

0

-33

-3539 139 209 0

0 3

-33

-3367 207 351 0

0

-33

-3518 160 209 0

0 13

-33

-3340 231 350 0

0

-33

-3493 180 208 0

0 233 180 11400 12840 210 0

0 180 16290 17350

-536 0

0 2213

-74

-5983

-2660 293 0

0

-74

-7056

-3558 322 0

0 2393 149 8475 9884 164 0

0 149 12580 13670

-416 0

0 6773

-51

-4443

-1020 320 0

0

-51

-5018

-1463 256 0

0 7193 231 12680 13780 145 0

0 231 17340 18140

-588 0

0 6395 7493 10

-142 2054 221 0

0 10 164 2398 45 0

0 11093

-40

-3276

-654 256 0

0

-40

-3669

-954 192 0

0 16457

-47

-4080

-479 352 0

0

-47

-4491

-773 244 0

0 16517

-41

-3813

-231 351 0

0

-41

-4095

-404 230 0

0 16518

-28

-3689

-110 350 0

0

-28

-3383 297 199 0

0 17118

-33

-3241 307 349 0

0

-33

-3393 255 204 0

0 17119 3

-2918 623 348 0

0 3

-1521 2098 125 0

0 57120

-33

-3283 279 350 0

0

-33

-3439 223 206 0

0 0

3078 2100 4262 554 0

0 3078 4281 5859 577 0

0 3

3078 2100 4262 554 0

0 3078 4280 5856 577 0

0 13 3078 2099 4263 554 0

0 3078 4278 5853 576 0

0 233 823 6811

-8426

-847 0

0 823 12480 38540 5953 0

0 2213 3002

-447 2916 683 0

0 3002 1782

-3944

-735 0

0 2393 799 3298

-10540

-506 0

0 799 9988 25870 4515 0

0 6773 2953

-85 3049 980 0

0 2953 2409

-2931

-397 0

0 7193 1539 6354

-2971 49 0

0 1539 9542 24620 4575 0

0 7493 1642 7294 6946 137 0

0 1642 6282 20660 2675 0

0 11093 2290 364 2825 500 0

0 2290 2225 882

-131 0

0 16457 3195 285 3758 754 0

0 3195 3045 526

-230 0

0 16517 3191 304 3705 753 0

0 3191 3131 687

-181 0

0 16518 3182 300 3699 752 0

0 3182 3120 680

-180 0

0 17118 3157 1120 3848 706 0

0 3157 3802 3273 233 0

0 17119 3127 1109 3832 704 0

0 3127 3771 3247 235 0

0 57120 3077 2085 4216 543 0

0 3077 4274 5877 573 0

0 Note: Not all time steps are listed in this table.

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Figure 1. RO Nozzle Internal Pressure Distribution File No.: 0801038.305 Revision: 1 Page 12 of 16 F0306-O1RO

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AN Figure 2. RO Nozzle Pressure Cap Load & Boundary Condition 0-1.,V

-1.VI PIMU.

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ELEMENTS U

ANŽ~

Recirc Outlet Nozzle Finite Element Model Figure 3. RO Nozzle Vessel Wall Boundary Condition File No.: 0801038.305 Revision: 1 Page 14 of 16 F0306-01RO

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NODAL SOLUTION STE.P=686:

SUB =1 SINT

(,Av)

DMX -. 85555 3 SMN =175.054 smx:=89456 GI~b~l X -

12.968 SInT 2 945E.7 Hlp AN>

175.1.015A Vermont Yankee

/

.20015

.39.8,5:5, 159695 17 95 *36 10095 2.9 9 35.5:

4ý9 7 7 5 Recirc Otuhlet.Nozzle, Transient 9 Stress 69616 89:456 Figure 4. Safe End Critical Thermal Stress Intensity Location File No.: 0801038.305 Revision: 1 Page 15 of 16 F0306-01RO

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STEP--I SUB =1 TIME--

SIMW (AVG)

Mxx --+/-Z72 IL38 7835 2.4531 4486 11183 17879 Recirc outlet. NOZZ1 Fini~te EleMent Hc~del AN x:

$~Z 3/4&Y2~ *$~P~W~4 21227 27924 31272 24575 Figure 5. Nozzle Blend Radius Limiting Pressure Stress Intensity Location NOdc $701 Node (6391 NIode329 Figure 6. Limiting Stress Paths File No.: 0801038.305 Revision: 1 Page 16 of 16 F0306-01RO Structural Integrity Associates, Inc.

File No.: 0801038.306 CALCULATION PACKAGE Project No.: 0801038 Quality Program Z Nuclear [] Commercial PROJECT NAME:

VY Confirmatory Analyses for CS and RO Nozzles CONTRACT NO.:

10163217 Amendment 5 CLIENT:

PLANT:

Entergy Nuclear Operations, Inc.

Vermont Yankee Nuclear Power Station CALCULATION TITLE:

Fatigue Analysis of Reactor Recirculation Outlet Nozzle Project Manager Preparer(s) &

Revision Description Approval Checker(s)

Signature & Date Signatures & Date 1 - 18 Computer Files Initial issue.

Gary L. Stevens 01/07/09 Tyler Novotny 01/07/09 Jennifer E. Smith 01/07/09 i

1 1-4, 6-12, 14-19 Computer Files Revised per summary contained in Section 1.1.

Changes are marked with "revision bars" in right-hand margin.

Gary L Stevens 03/09/09 Preparer:

Tyler D. Novotny 03/09/09 Checker:

William F. Weitze 03/09/09 Page 1 of 19 F0306-O1RO

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Table of Contents 1.0 OBJECTIVE.................................................................................................................................

3 1.1 Changes M ade in Revision 1 of this Calculation...........................................................

3 2.0 M ETHODOLOGY......................................................................................................................

3 3.0 DESIGN INPUTS.........................................................................................................................

4 3.1 Stress Calculation......................................................................................................

4 3.2 Fatigue Usage Analysis, General....................................................................................

4 3.3 Event Cycles, VESLFAT................................................................................................

5 3.4 M aterial Properties, VESLFAT......................................................................................

5 3.5 Stress Indices......................................................................................................................

6 4.0 CALCULATIONS........................................................................................................................

6 5.0 RESULTS OF ANALYSIS...................................................................................................

7

6.0 CONCLUSION

S AND DISCUSSIONS.................................................................................

7

7.0 REFERENCES

8 List of Tables Table 1: Safe End Load Sets as Input to VESLFAT......................................................................

9 Table 2: Nozzle Blend Radius Load Sets as Input to VESLFAT.......................................................

11 Table 3: Temperature-Dependent M aterial Properties for VESLFAT (3)........................................... 12 Table 4: Carbon/Low Alloy Steel and Stainless Steel Fatigue Curves..........................................

13 Table 5: Pressure and Attached Piping Unit Load Case Stress Components.................................

14 Table 6: Fatigue Usage Calculation for the Safe End....................................................................

15 Table 7: Fatigue Usage Calculation for the Nozzle Blend Radius...............................................

16 Table 8: EAF Fatigue Usage Calculation for the Nozzle Blend Radius Location........................ 17 Table 9: Linearized Stress Files Compiled for VY-RO-StressResults.xls......................................

19 File No.: 0801038.306 Page 2 of 19 Revision: 1 F0306-O1RO

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1.0 OBJECTIVE The objective of this calculation package is to perform an ASME Code,Section III fatigue usage evaluation and a plant-specific evaluation of reactor water environmental effects for the reactor pressure vessel (RPV) recirculation outlet (RO) nozzle at the Vermont Yankee Nuclear Power Station.

1.1 Changes Made in Revision 1 of this Calculation Description of changes made in Revision 1 of this calculation:

a.

Editorial changes were made to Table 1 to more precisely describe the transient load sets.

b.

All but one of the changes made to Table 2 were editorial to more precisely describe the portions of the transients. The one non-editorial change was to move a time split in Transient 9 to better catch a stress peak or stress valley.

c.

Table 3 and the corresponding VESLFAT input file were revised to reflect actual material properties for the safe end. Revision 0 of this calculation tabulated SA-182 F304 (18Cr -8Ni) properties, but actually used properties for an Alloy 600 material.

d.

Table 5 was changed to eliminate the application ofKt = 1.53 to the nozzle comer piping loads.

e.

Tables 6, 7, and 8 were revised to reflect the new fatigue usage and environmental assisted fatigue summaries as a result of the changes associated with Bullets b and c above.

f.

Table 8 was revised for editorial changes.

g.

The results of various sensitivity studies on fatigue usage were added to Section 5.0.

h.

Revision of CUF values in Sections 5.0 and 6.0 to reflect revised analyses.

i.

All remaining changes marked throughout this calculation are editorial changes made to the text of the calculation package.

2.0 METHODOLOGY The methodology to be used for this evaluation was established in a previous calculation package

[2]. Based on that methodology, thermal stresses, pressure stresses, and attached piping load stresses were developed in the Reference [1] calculation for use in this fatigue calculation. The thermal stresses are added to pressure stresses and attached piping load stresses1. Both the pressure and piping load stresses are scaled based on the magnitudes of the pressure and nozzle fluid temperature during each transient. All six components of the stress tensor from the stress results are used in the fatigue calculation.

Stress components due to piping loads are scaled assuming no stress occurs at an ambient temperature of 70'F and the full values are reached at a reactor design temperature of 575°F [2, Assumption 3.1.7]. In addition, design seismic and deadweight loads are also included and scaled in combination with the thermal loads for each transient. This combination, coupled with assigning the stress due to these loads the same sign as the thermal stress, is considered to be a very conservative treatment of the loads overall in that deadweight and design seismic loads are considered and scaled for every transient.

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The fatigue calculation is performed for both the limiting safe end and nozzle blend radius locations, as determined in the Reference [1] calculation, and uses the methodology of Subarticle NB-3200 of Section III of the ASME Code [3]. An environmental fatigue usage analysis is also performed in this calculation applying the methodology and associated environmental fatigue multipliers described in Reference [6].

3.0 DESIGN INPUTS 3.1 Stress Calculation Linearized stress components at Node 6395 (limiting safe end path at inside surface) and Node 3829 (limiting nozzle blend radius path at inside surface) are used for the fatigue usage calculation, as shown in Figure 6 of Reference [1 ]. For the nozzle blend radius location, the stresses used in the evaluation are for the base metal only; that is, the cladding material is unselected prior to stress extraction. The stress components from the thermal stress analyses are combined with stress components due to pressure and piping loads. The linearized thermal stress components for each transient are taken from the relevant output files in the Reference [1] calculation (a sample of which was provided in Table 3 of Reference [1]). The unit pressure stress component results are taken from Table 2 of Reference [1]. Piping load stress components are taken from Table 5 of the Reference [2] calculation.

3.2 Fatigue Usage Analysis, General Structural Integrity's VESLFAT program [4] is used to perform the fatigue usage calculation in accordance with the fatigue usage portion of ASME Code,Section III, Subarticle NB-3200 [3].

VESLFAT performs the analysis required by NB-3222.4(e) [3] for Service Levels A and B conditions defined by the user. The VESLFAT program computes the primary-plus-secondary and total stress ranges for all events and performs a correction for elastic-plastic analysis, if necessary.

The program computes the stress intensity range based on the stress component ranges for all event pairs [3, NB-3216.2]. The program evaluates the stress ranges for primary-plus-secondary and primary-plus-secondary-plus-peak stresses based on all six components of stress (3 normal and 3 shear stresses). If the primary-plus-secondary stress intensity range is greater than 3 Sm, the total stress range must be increased by the simplified elastic-plastic strain correction factor, Ke, as described in NB-3228.5 [3]. The design stress intensity, Sm, is specified as a function of temperature. The input maximum temperature for both states of a load set pair is used to establish the Sm value used in the fatigue calculations from the user-defined input values.

When more than one stress set is defined for either of the event pair loadings, the stress differences are determined for all of the potential stress pairs, and the pair producing the largest alternating total stress intensity (Salt), including any effects of K,, is used. The principal stresses for the stress ranges are determined by solving for the roots of the following cubic equation2:

S 3 _.(Cyx + a

+ *z)S2 + ((7x (T +

-y az +

7z cyx-2Txz 2 _ "y 2 )S 2 Note that a., ay, a., etc. are used synonymously with S., S,, Sz, etc., in this calculation.

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-((Y" 3'y T, "+- 2 Txy Txz Zy - a, TXy2 _ Ty "xz2 _ Tx "yz2 o

The stress intensities for the event pairs are reordered in decreasing order of SaIt, including a correction for the ratio of modulus of elasticity (E) from the fatigue curve divided by E from the material evaluated at the maximum event temperature. This allows a fatigue table to be created to eliminate the number of cycles available for each of the transient events. This fatigue table is based on a worst-case progressive pairing of events in order of the most severe alternating stress to the least severe, allowing determination of a bounding fatigue usage per NB-3222.4(e) [3]. For each load set pair in the fatigue table, the allowable number of cycles is determined based on Salt.

3.3 Event Cycles, VESLFAT For the Vermont Yankee RO nozzle analysis, transients that consist of combined stress peaks or valleys are split so that each successive peak or valley is treated separately. Therefore, there are 61 load sets based on the combined stress changes for the safe end, and 46 load sets based on the combined stress changes for the nozzle blend radius location. The reason the number of load sets are not equal for each path is because the time history stress results of those paths differ. Tables 1 and 2 show the load sets applicable to plant operation, with cycle counts per Table 1 of Reference [2].

These are used as input to VESLFAT for the safe end and nozzle blend radius locations, respectively. The cycle counts of Reference [2, 7] consider 60 years of operation. The data from Table 1 is entered into the VESLFAT input files VY-RO-VFAT-1L CYC (safe end) and the data from Table 2 is entered into the file VY-RO-VFAT-21. CYC (nozzle blend radius).

3.4 Material Properties, VESLFAT Material properties are entered in VESLFAT input files VY-RO-VFAT-1I.FDT (safe end) and VY-RO-VFAT-2I.FDT (nozzle blend radius). Table 3 lists the temperature-dependent material properties used in the analysis [5]. Table 4 lists the fatigue curve for the nozzle blend radius and safe end materials

[3, Appendix I, Table 1-9.1 and Figure 1-9.1 (UTS *80.0 ksi) for the nozzle blend radius, and Tables 1-9.1 and 1-9.2.2 (Curve C) and Figures 1-9.2.1 and 1-9.2.2 for the safe end location]. Curve C is selected for the safe end location because it is the most conservative curve among the three extended curves for austenitic steel. VESLFAT automatically scales the stresses by the ratio of E on the fatigue curve to E in the analysis, for the purposes of determining allowable numbers of cycles, as required by the ASME Code.

Other material properties are input as follows:

m = 1.7, n = 0.3, parameters used to calculate Ke for the safe end location [3, Table NB-3228.5(b)-l]

m = 2.0, n = 0.2, parameters used to calculate Ke for the nozzle blend radius location [3, Table NB-3228.5(b)-i]

E from fatigue curve = 28,300 ksi [3, Appendix I, Figure 1-9.2] for the safe end location.

E from fatigue curve = 30,000 ksi [3, Appendix I, Figure 1-9.1] for the nozzle blend radius location.

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3.5 Stress Indices The limiting stress path for the RO nozzle safe end is defined in Reference [I]. The stresses caused by the piping were hand calculated and do require a stress concentration factor, if appropriate. The stress concentration factor for the safe end location is 1.53 [2, Section 3.8]. This value is conservatively used for both the C2 and K2 values required by the ASME Code [3, NB-3600]. The piping loads are relatively minor in comparison to the other loads this nozzle experiences so the conservative C2 and K2 values will have a small impact on the analysis. Table 5 shows the piping loads after applying the C2 and K2 values as appropriate.

4.0 CALCULATIONS Table 5 contains the stress components at the locations of interest for the 1,000 psi unit pressure stress case [1, Table 2]. Table 5 also contains the stress components for the attached piping load unit stress case [2, Table 5], which correspond to a reactor design temperature of 575°F [2, Section 3.1.7].

The attached piping load stress components were applied assuming the same signs as the thermal stress, which yields the largest stress component ranges.

The calculations of all of the VESLFAT stress inputs are automated in Excel workbooks VY-RO-VFAT-]i.xls (safe end) and VY-RO-VFAT-2i.xls (nozzle blend radius). These files are organized with sheets labeled as follows:

Overview: Contains general information.

Other Stresses: Contains pressure and attached piping load stresses. As shown in Table 5, the pressure stresses use the membrane-plus-bending and total stress from the finite element analysis [1].

Rearranger: There are 12 Rearranger sheets, one for each thermal transient as analyzed by ANSYS. In these sheets, thermal stresses are copied from Excel workbook VY-RO-StressResults.xls, and rearranged to conform to VESLFAT input format (including switching the shear stress components S,. and Sy, as required by VESLFAT). VY-RO-StressResults.xls contains the results of the ANSYS stress linearization for each transient. The files contained within this workbook are shown in Table 9. Time-varying scale factors for the attached piping loads (based on path metal temperature) and pressure are determined, and used to scale the unit load case stresses, which are then added to the thermal stresses. Since the attached piping loads can act in any direction, the stresses due to the attached piping loads are assigned the same sign as the thermal stresses to maximize the component stresses.

Algebraic summation of all six stress components is performed for pressure, piping loads, and thermal stresses at each transient time step. The VESLFAT stress input also includes time-varying metal temperature, as obtained from the ANSYS output, which is used to determine temperature-dependent properties from the values in Table 3.

VESLFAT: Contains the VESLFAT stress input, as obtained from the Rearranger sheets.

Load set numbers are entered on this sheet, as defined in Table 1 and Table 2. These sheets are saved to VESLFAT input files VY-RO-VFAT-li.STR (safe end) and VY-RO-VFAT-2i.STR (nozzle blend radius).

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5.0 RESULTS OF ANALYSIS Table 6 and Table 7 provide the detailed calculated 60-year fatigue usage, as obtained from VESLFAT output files VY-RO-VFA T-1J.FAT (safe end) and VY-RO-VFAT-2LFA T (nozzle blend radius). All VESLFAT input and output files are saved in the project computer files associated with this calculation.

From Table 6, the safe end cumulative usage factor (CUF) is 0.00308 for 60 years. From Table 7, the nozzle blend radius CUF is 0.0 175 for 60 years.

From Table 1 of Reference [6], it was determined that hydrogen water chemistry (HWC) is available for 47% of the total 60-year operating period, and normal water chemistry (NWC) is present for the remaining 53% of the total 60-year operating period. From Table 1 of Reference [6], the dissolved oxygen values for the recirculation line (which is applicable to the RO nozzle) are 48 ppb for HWC conditions and 122 ppb for NWC conditions.

For the stainless steel piping, the environmental fatigue factors for post-HWC and pre-HWC are 15.35 and 8.36 from Table 2 of Reference [6]. The overall environmental multiplier is found by (15.35 x 47% + 8.36 x 53%), which equals 11.645, conservatively rounded up to 11.7. Therefore, the overall environmental multiplier is 11.7, which results in an EAF adjusted CUF of 11.7 x 0.00308 =

0.0360 for 60 years, which is acceptable (i.e., less than the allowable value of 1.0).

Based on the detailed CUF calculation shown in Table 7, a detailed EAF adjusted CUF evaluation on a load-pair basis is provided for the nozzle blend radius location in Table 8. The EAF usage from Table 8 is 0.111 for 60 years, which is less than the allowable value of 1.0 and is therefore acceptable. The effective overall Fen is 0.111/0.0175 = 6.32.

As a part of fatigue analysis calculations, it was noted that using Fy = -20 kips in the piping loads caused a slightly higher total stress intensity. However, the change was determined to have an insignificant effect on fatigue usage results. In addition, the effect of modeling the distinct material properties of both Type F304 and Type F316 in the ANSYS analysis (as opposed to using 18Cr-8Ni properties) was determined to have an insignificant effect on fatigue usage results. Finally, the effect of applying a minimum temperature of 130'F for thermal boundary Region 2 (see Figure 1 of Reference [2]) was determined to have an insignificant effect on fatigue usage results. These investigations and associated results are contained in the project files.

6.0 CONCLUSION

S AND DISCUSSIONS Detailed fatigue calculations for the Vermont Yankee RO nozzle were performed based on the results of stress analyses previously performed [1]. The thermal stresses were combined with stresses due to pressure and attached piping loads, both of which were scaled based on the magnitudes of the pressure and metal temperature during each thermal transient. All six components of the stress tensor were used for the fatigue calculations. The fatigue calculations were performed at previously-determined limiting locations in the safe end and nozzle blend radius, and used the methodology of Subarticle NB-3200 of Section III of the ASME Code [3].

File No.: 0801038.306 Page 7 of 19 Revision: 1 F0306-O1RO

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The 60-year CUF for the safe end location was determined to be 0.00308 and the CUF for the nozzle blend radius location was determined to be 0.0175. Both values are less than the ASME Code allowable value of 1.0, and are therefore acceptable.

Detailed EAF assessments were also performed for the two RO nozzle locations. The 60-year EAF CUF for the safe end location was determined to be 0.0360. The 60-year EAF CUF for the nozzle blend radius location was determined to be 0.111 using temperature-dependent Fen multipliers for each load pair. Both values are less than the ASME Code allowable value of 1.0, and are therefore acceptable.

7.0 REFERENCES

1. Structural Integrity Associates Calculation No. 0801038.305, Revision 1, "Stress Analysis of Reactor Recirculation Outlet Nozzle."

2. Structural Integrity Associates Calculation No. 0801038.304, Revision 1, "Design Inputs and Methodology for ASME Code Fatigue Usage Analysis of Reactor Recirculation Outlet Nozzle."

3. ASME Boiler and Pressure Vessel Code,Section III, 1998 Edition with 2000 Addenda.

4. VESLFAT, Version 1.42, 02/06/07, Structural Integrity Associates.

5. ASME Boiler and Pressure Vessel Code,Section II, Part D-Properties, 1998 Edition with 2000 Addenda.

6. SI Calculation No. VY-16Q-303, Revision 0, "Environmental Fatigue Evaluation of Reactor Recirculation Inlet Nozzle and Vessel Shell/Bottom Head."

7. Entergy Design Input Record (DIR) EC No. 1773, DIR. Revision 1, "Environmental Fatigue Analysis for Vermont Yankee Nuclear Power Station," 7/26/07, SI File No. VY-16Q-209.

8. Deleted (not used in this calculation).

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Table 1: Safe End Load Sets as Input to VESLFAT YES Lo~

LFAT Transient Start ad Set Time, see 1

1Trn1 0

2 2Trnl 1616.4 3

1Tin2 0

4 2Trn2 0.4 5

3Tm2 301 6

4Trn2 601.4 7

1Trn3 0

8 2Trn3 250 9

3Trn3 2050 10 4Tm3_

2960 11 5Tm3_

5560 12 lTrn4 0

13 2Tm4_

2 14 3Tin4 7

15 4Tm4 46 16 5Trn4_

992 17 6Tn4_

2294 18 7Trn4 3050 19 8Tm4_

6899 20 9Trn4_

7745 21 1OTrn4_

8645 22 11Trn4_

11057 23 12Tm4_

16166 24 13Tm4_

16818 25 14Trh4_

17118 26 lTrn5 0

27 2Tm5_

1.5 28 3Trn5 24 29 4Trn5 2310 30 5Trn5 2611 31 6Trn5 2911.4 32 ITm6 0

33 2Tm6 0.6 34 3Trn6 20 35 4Tin6 2312 36 5Trn6_

2613 37 6Tm6_

2913.6 Temp Change Pressure Change Up Up Down Down Down Down Up Up Down Up & Down Down None None Down Down & Up Up & Down Down & Up Up & Down Down & Up Up Up Up Up & Down None Down None None Up Down None Down None None Up Down None Down Up Up None None None None None None None None None Up Up & Down Down None Down Down & Up Up & Down Down Down Down Up Up None None Up Up & Down Down & Up None None None Up Up & Down Down & Up None None None Cycles 300 300 300 300 300 300 10 10 10 10 10 10 10 10 10 10 10 10 10 10 10 10 10 10 10 60 60 60 60 60 60 1

1 1

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Table 1 (continued): Safe End Load Sets as Input to VESLFAT VESLFAT Load Set 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60 61 Start Transient ties Time, see 1Trn7_

0 2Tm7_

37.5 3Trn7 600 4Tin7 4443 1Trn8_

0 2Tn8_

3 3Trn8 2295 4Tin8_

3927 1Trn9 0

2Tm9_

0.12 3Trn9_

27.92 4Trn9_

290.15 1Trn10 0

2Trnl0_

730.8 3Trnl0 6314 4TmlO 6844 5TrnlO 9555 6Tmr0_

14937 1Tmrl 1 0

2Trnl 1 0

3Tmll 0

1Trnl2 0

2Trnl2 0

3Trnl2_

0 Temp Change Down Down Down Down None Up Down None Down Down & Up Up None Down Down Down Down Down Down None None None None None None Pressure Change Cycles Down Down Down Down Down Down & Up None None None None None None Down Down Down Down Down Down None Up Down None Up Down 228 228 228 228 300 300 300 300 300 300 120 120 120 1

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Table 2: Nozzle Blend Radius Load Sets as Input to VESLFAT VESLFAT Load Set 1

2 3

4 5

6 7

8 9

10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 Start Transient

Time, sec ITml 0

2Trn1 808.2 1Tin2_

0 2Trn2 0.4 3Trn2 401 lTrn3_

0 2Trn3_

250 3Trn3_

2325 4Trn3_

3510 5Trn3 5060 1Trn4_

0 2Tm4_

2 3Tm4_

7 4Trn4_

46 5Tin4 1091 6Tm4_

2348 7Tm4_

3269 8Tm4_

6983 9Trn4_

7745 1OTm4_

13839 11Trn4_

16918 12Trn4_

18986 1Tm5_

0 2Trn5_

24 3Trn5_

2611 1Trn6_

0 2Tm6_

0.6 3Trn6_

20 4Tm6_

2663 1Tm7_

0 2Trn7_

37.5 3Trn7_

2247 1Trn8_

0 Temp Change Pressure Change Cycles Up Up Down Down Down Up Up & Down Down & Up Up & Down Down None None Down Down & Up Up & Down Down & Up Up & Down Down & Up Up Up & Down Down None None Up & Down Down None None Up & Down Down Down Down Down None Up Up None None None None None None None None Up Up & Down Down None Down Down & Up Up & Down Down Down & Up Up None None Up & Down Down & Up None Up Up & Down Down & Up None Down Down Down Down 300 300 300 300 300 10 10 10 10 10 10 10 10 10 10 10 10 10 10 10 10 10 60 60 60 1

1 1

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Table 2 (continued): Nozzle Blend Radius Load Sets as Input to VESLFAT VESLFAT Load Set Transient Start

Time, sec Temp Change Pressure Change Cycles 34 35 36 37 38 39 40 41 42 43 44 45 46 2Trn8 3Trn8 lTrn9 2Trn9 3Trn9_

1Tml0_

2Trnl0_

1Tml 1_

2Trn 1_

3Tml I-lTrnl2_

2Tml22 3Tm12_

3 2025 0

9 58 0

313.2 0

0 0

0 Up & Down Down Down Up None Down Down None None None None None None Down & Up None None None None Down Down None Up Down None Up Down 228 228 300 300 120 120 120 1

1 Table 3: Temperature-Dependent Material Properties for VESLFAT (3)

Material T, OF E x 106, psi S., ksi SY, ksi SA-508 Class 2 (nozzle blend radius(2))

70 200 300 400 500 600 70 200 300 400 500 600 27.8 27.1 26.7 26.1 25.7 25.2 28.3 27.6 27.0 26.5 25.8 25.3 26.7 26.7 26.7 26.7 26.7 26.7 20 20 20 19.3 18.0 17.0 50.0 47.0 45.5 44.2 43.2 42.1 30 25.9 23.4 21.4 20.0 18.9 SA-182 F316 (Safe End (1))

Notes:

1.

For the safe end material, SA-182 F316 (16Cr-12Ni - 2Mo) austenitic stainless steel properties are used.

2.

For the nozzle blend radius material, SA508 Class 2 material properties are used (3/4Ni-1/2Mo-1/3Cr-V), per Reference [2].

3.

All values are taken from Reference [5].

4.

SA-508 Class 2 in the Code of Construction is the same as SA-508 Gr. 2 Class 2 in the 1998 ASME Code [5].

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Table 4: Carbon/Low Alloy Steel and Stainless Steel Fatigue Curves S,, ksi Sa, ksi Number of Cycles Carbon/Low Alloy )

Austenitic 10 580 708 20 410 512 50 275 345 100 205 261 200 155 201 500 105 148 1000 83 119 2000 64 97 5000 48 76 10000 38 64 20000 31 55.5 50000 23 46.3 100000 20 40.8 200000 16.5 35.9 500000 13.5 31 1000000 12.5 28.2 2.E+06 N/A 22.8(2) 5.E+06 N/A 18.4(2) 1.E+07 N/A 16.4(2) 2.E+07 N/A 15.2(2) 5.E+07 N/A 14.3(2) 1.E+08 N/A 14.1(2) 1.E+09 N/A 13.9(2) 1.E+ 10 N/A 13.7(2) 1.E+11 N/A 13.6(2)

Note:

1. Using UTS < 80 ksi curve.

2.

Using Curve C for austenitic steel.

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Table 5: Pressure and Attached Piping Unit Load Case Stress Components Node Membrane plus Bending (1)

Total (')

s~ s s()s(5)

Sx Sy S

Sx Sy(z)(5)

Sa (2)xy S

xz YZ

(

Sxz (5) yz Pressure (3) 6395

-955.2 4420 10390 15.26 0

0

-955.2 4912 10530

-222.6 0

0 3829

-718.7

-951.7 25000 4708 0

0

-718.7 206.2 30150 733.2 0

0 Piping (4) 6395 0

7930 0

831 2066 0

0 12133 0

1271 3160 0

3829 0

218 0

42 49 0

0 218 0

42 49 0

Notes:

1.

All stress values are in units of psi.

2.

The safe end location is represented by Node 6395, and the nozzle blend radius location is represented by Node 3829.

3.

The stresses for both nodes represent the stress due to an applied pressure of 1,000 psig.

4.

Piping stresses for both locations represent the stress due to full attached piping loads at an RIPV temperature of 575'F.

5.

Sy. and S. components have been rearranged from the ANSYS output in order to be in correct order for VESLFAT.

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Table 6: Fatigue Usage Calculation for the Safe End Load

1 47 15 15 19 17 28 18 28 34 43 6

6 6

2 2

33 13 50 50 3

Desc.

Load Desc.

1

2

2 2Trn9_

48 3Trn9_

4Trn4_

49 4Trn9_

4Trn4_

28 3Trn5_

8Trn4_

28 3Trn5_

6Trn4 28 3Trn5_

3Trn5_

39 2Trn7_

7Trn4_

28 3Trn5_

3Trn5_

44 3Trn8_

3Trn6_

44 3Trn8_

2Trn8_

44 3Trn8_

4Trn2_

43 2Trn8_

4Trn2_

35 4Trn6_

4Trn2_

29 4Trn5_

4Trn2_'

22 11Trn4_

4Trn2 23 12Trn4 2Trnl_

6 4Trn2_

2Trnl_

31 6Trn5_

2Trnl_

37 6Trn6_

2Trnl 25 14Trn4 2Trnl_

40 3Trn7_

2Trnl_

16 5Trn4_

2Trn6_

52 3Trnl0_

2Trn4_

52 3Trn 10_

lTrnlO_

52 3TmlO0 lTrnlO_

53 4Trn 10_

1Trn2_

53 4Trnl0_

n (cycles)

Sn (psi) 1 1

9 10 10 1

10 20 1

207 21 1

60 10 10 198 60 1

10 1

10 289 11 289 79715 30275 29755 26926 25213 20321 19961 4606 4606 4606 4028 3519 3484 11783 3202 3193 3319 3319 1702 18894 5069 12380 10470 9634 18796 18795 Ke 2.62 1

1 1

Salt Si)

Nallow U

(psi) 169777 331.52 0.00302 23722 1757500 0.00000 23610 1784800 0.00001 21352 2647400 0.00000 20492 3155800 0.00000 16926 8269400 0.00000 16731 8866300 0.00000 16450 9819700 0.00000

16450, 9819700 0.00000 16450 9819700 0.00002 16176 11335000 0.00000 15752 14441000 0.00000 15637 15446000 0.00000 15613 15666000 0.00000 15588 15895000 0.00000 15583 15936000 0.00001 15531 16430000 0.00000 15531 16430000 0.00000 15055 23098000 0.00000 14987 24732000 0.00000 14487 41157000 0.00000 14460 42317000 0.00000 13875 1.336E+09 0.00000 13841 1.968E+09 0.00000 13770 4.465E+09 0.00000 13769 4.491E+09 0.00000 Total 0.00308 Note:

usage =

All other load pairs have an alternating stress, Salt, that is below the endurance limit of the fatigue curve. Therefore, they do not contribute to fatigue usage.

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Table 7: Fatigue Usage Calculation for the Nozzle Blend Radius Load Desc.

Load Desc.

n Salt

1

1

2

2 (cycles)

(psi)

Nallow U

1 1

2 1

1 1

2 2

4 4

4 10 35 35 9

7 7

7 3

3 31 1Trn1_

14 4Trn4_

10 1Trnl_

37 2Trn9_

1 1Trnl1 16 6Trn4_

10 1Trnl_

27 2Trn6_

1 2Trnl1 45 2Trn 12 1

1Trn1_

15 5Trn4_

10 1Trnl 18 8Trn4 10 1Trnl_

36 lTrn9_

1 1Trnl_

13 3Trn4_

10 1Trn1_

38 3Trn9_

1 1Trnl_

12 2Trn4_

10 1Trnl_

23 1Trn5_

60 1Trn1_

17 7Trn4_

10 1Trn1_

5 3Trn2_

166 2Trnl 5

3Trn2_

134 2Trnl_

28 3Trn6_

1 2Trnl_

11 1Trn4_

10 2Trnl_

26 1Trn6_

1 2Trnl_

25 3Trn5_

60 2Trnl_

29 4Trn6_

1 2Trnl1 8

3Trn3 10 2Trnl_

4 2Trn2_

82 2Trn2_

41 1Trnl 1_

120 2Trn2_

32 3Trn7_

1 2Trn2_

40 2Trnl0 97 5Trn3_

40 2Trnl0_

10 3Trn8_

40 2Trn 10 193 3Trn8_

43 3Trnl 11 35 4Trn3_

43 3Trnl 1_

10 2Trn3_

46 3Trn 12_

1 2Trn3 44 1Trn12_

1 2Trn3 43 3Trn 11 8

1Trn2_

43 3Trnl 1 67 1Trn2_

19 9Trn4_

10 2Trn7_

42 2Trn 11 1

21902 1.00 43085 6889 0.0015 21390 1.00 32177 17617 0.0001 15100 1.00 31137 19701 0.0005 42381 1.00 27020 30496 0.0000 45773 1.00 26852 31084 0.0000 18457 1.00 26707 31604 0.0003 13066 1.00 26562 32139 0.0003 28617 1.00 24546 40947 0.0000 34179 1.00 24042 43643 0.0002 25904 1.00 23939 44218 0.0000 36762 1.00 23612 46129 0.0002 35051 1.00 22617 54348 0.0011 22210 1.00 22533 55358 0.0002 29847 1.00 22312 58126 0.0029 29301 1.00 22309 58168 0.0023 33856 1.00 22227 59234 0.0000 33460 1.00 21959 62919 0.0002 32908 1.00 21661 67330 0.0000 29068 1.00 21226 74454 0.0008 29068 1.00 21226 74454 0.0000 29847 1.00 21214 74661 0.0001 30245 1.00 21092 76819 0.0011 32229 1.00 20851 81328 0.0015 30983 1.00 20125 96967 0.0000 30982 1.00 20124 96981 0.0010 31344 1.00 20033 99198 0.0001 29931 1.00 19888 102050 0.0019 29651 1.00 19696 105678 0.0003 30915 1.00 19357 112494 0.0001 30523 1.00 19349 112655 0.0000 30523 1.00 19349 112655 0.0000 30523 1.00 19349 112655 0.0001 31236 1.00 19331 113042 0.0006 23810 1.00 16958 181219 0.0001 27376 1.00 11515 infinite 0.0000 Total 0.0175 Usage =

Note:

All other load pairs have an alternating stress, Salt, that is below the endurance limit of the fatigue curve. Therefore, they do not contribute to fatigue usage.

File No.: 0801038.306 Revision: 1 Page 16 of 19 F0306-01RO

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Table 8: EAF Fatigue Usage Calculation for the Nozzle Blend Radius Location VY RO Nozzle Corner Environmental Fatigue Calculation CUF Calculation from file VY-RO-VFAT-2i fat:

Index Load#fIDescription #1h, icycles) t5 Load2n#Description:n, (cycles) ý5) t (cycles) (5 S, (psi)

K.

Sý (psil i

U I

1 1

1Trn_

300 14 i

4Trn4 10 10 21902 1.00 43085 68&9 G.0015 2

1 1Trn_

290 37 2Trnm_

1 1

21390 1.00 32177 17617 0.0001 3

1 1Tin1 289 16 6Trn4 10 10 15100 1.00 31137 19701 0.0005 4

1 i

Trn1 279 27 2Trn6 1

1 42381 1.00 27020 30496 0.0000 5

2 2Trnl 1 300 45

[

2Trn12 1

1 4577.3 1,00 L6852 31084 0.0000 6

1 iTrnl 278 15 5Trn4 10 10 18457 1.00 26707 31604 0.0003 7

1 1Trnt_

m6 18 BTrn4 10 1.0 130-6 1.00 26562 32139 0.0003 8

1 1TrnlI 258 36 1lTrn.

[

1 1

28617 1.00 24546 40947 0.0000 9

1 1Trnl i

257 13 i

3Trn4 i

10 10 34179 1.00 24042 43643 D.0002 10 1

1 ITrnl_

247 38 3Trn_

1 1

25904 1.00 23939 44218 0.0000 11 1

ITrnl 246 12 2Trn4 10 10 38762 1.00 23612 468129 0.0002 12 1

lTrnlI 236 23 lTrnS 60 6(]

35051 1.00 22617 54348 0.0011 13 1

1Tml _

176 17 7Trn4 10 t10 22210 1.00 22533 55356 0.0002 14 1

ITrnlI 1C-6 5

3Trn2_

300-166 29847 1.00 22312 58126 0.0029 15 2

2TrnlI 299 5

3Trn2 134 134 29301 1.00 22309 58168

.0023 16 2

Trnl_

165 28 3Trn6 1

1 33856 1.00 22227 59234 0.0000 IT 2

trr i

164 11 lTrn4_

10 10 33460 1.00 21959 652919 0.0002 18 2

2Trn 1 154 26 TrnS 1

1 32908 1.00 21661 67330 0.0000 19 2

2Trnl 153 25 3Trn5 60 60 29068 1.00 21226 744,54 0.0008 20 2

i 2Trnl_

93 29 4Trn6 1

1 29068 1.00 21226 74454 0.0000 21 2 z 2Trnl I

2 5

3Krn3._

10 10 29847 1.00 21214 74661 0.0001 22 2

i2Trnt 2

4 2Trn2_

300 82 30245 1.00 2109G2 76819 0.0011 23 4

2rn2_

216 41 i

1Trnl_

120 120 32229 1.00 20851 81328 0.0015 24 4

i 27rn2 98 32 3Trn7 1

1 30983 1.00 20125 96067 0.0000 5

4 i

2Trn2_

97 40 2TrnlO 30 97 30982 1.00 20124 96081 0.0010 26 10 5Trn3 10 40 2TrnlO 203 10 31344 1.00 20033 99198 G.0001 27 35 3TrnB 228 40 2TrnlO 193 193 29931 1.00 198,8 102050 0.0019 28 35 i

3Trn_

35 43 3Trn1l 120 35 251 1.00 19H96 1 G%78 0.0003 29 9

4Trn3 I0 43 3Trnt1_

B5 10 30915 1.00 19357 112494 0.0001 30 7

2Trn3 10 46 3Trn12 1

1 30523 1.00 19349 112655 0.0000 31 7

27rn3 9

44 1TrnI2 I1 1

30523 1.00 19349 112655

.0O000 32 7

1 2Trn3 8

43 3TrnI11 75 8

30523 1.00 19349 112655 0.0001 33 3

1Trn2 300 43 3Trnt1_

67 67 31236 1.00 19331 113042 0.0006

34.

3 1Trn2 233 19 9Trn4 10 10 23810 1.00 16958 181219 0.0001 35

,31 2Trn7 1

42 2Trn11 120 1

27376 1.00 11515 infinite 0.0000 Total, U =

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Table 8 (continued): EAF Fatigue Usage Calculation for the Nozzle Blend Radius Location FAT -a'culanions:

MWIX M5D WttC DO Lao d H,"V..#.

c.

T,,4b e/RW~,~veff,.

49 122 ppb

% HWC =

47%

53%

= % NIWC Transient Maximum Temperatures:

From v'Y-R0-VFA7-2iALL':

Index Load#4IDesc.#l Load#21 Desc,#2 Line#

TI"(4) sa1(4) T2(4) s2(4) Sn (psi) T(**(.1) 1 1

iTrl_

14 4Trn4.

175 6 3

14 19 21902 339 2

1 iThol 37 2Trn9.

6065 1

3 37 62 21390 437 3

1 1Tin1_

16 6Thr4 1968 1

3 16 7

15100 329 4

1 ITrml_

27 2TrnR_

3734 1

3 27 8

42381 526 5

2 2Trnl_

45 2Trn12 201558 2

1 45 1

45773 120 6

1 ITrnl_

15 5Trn4 1927 1

3 1i 49 18457 394 7

1 ITral_

18 STrn4 2236 1

3 18 10 13066 335 a8 1

Tirn1_

36 1Trn9.

5157 1

3 36 41 20617 495 S

1 1Trnl_

13 3Trr.4 151 1

3 13 15 34179 516 10 1

1Trnl 31 3Trn9 6657 1

3 3a 1

25904 490 11 1

1Trnl 12 2Tro4.

159 1

3 12 3

36762 5,29 12 1

ITrn1 23 ITrs_

3115 1"

3 23 27 35051 526 13 1

1Tnl_

17 7Trn4.

2152 1

3 17 56 22210 426 14 1

ITni_

5 3Trn2 952 1

3 5

80 29847 530 15 2

2Trnl_

5 3Trn2_

8718 2

1 5

79 29301 530 16 2

2Trnl_

28 3Trn._

99727 2

1 28 1

33856 526 17 2

2Trnl_

11 lTrn_

42455 2

1 11 4

33450 526 18 2

2Tnl_

26 lTrn9_

98465 2

1 26 3

32908 526 19 2

27rnl_

25 3TrnS_

89557 2

1 25 22 29069 529 20 2

2TrMl 29 4TrnA.

105593 2

1 29 21 29068 5E29 21 2

2Trnl_

6 3Trn _

35741 2

1 6

5 26847 528 22 2

2,rnl_

4 2Tr 2m.

7777 2

1 4

7 30245 543 23 4

2 Trn 41 1Trn lI 233450 4

7 41 1

32229 543 24 4

2Trn,2_

32 3Trm7_

223647 4

7 32 126 30083 543 25 4

2Tro2 40 2TrnlO 232587 4

7 40 209 30982 543 26 10 5Trnq 40 2Trn 1_

1138571 10 21 40 209 31344 527 27 35 3Trn6.

40 2TrnlO 2891140 35 51 40 299 29031 528 28 35 3Trn,-

43 3Trn 11 2910647 35 51 43 1

29651 528 29 9

4Tfrn 3.

43 3Trnll_

1069326 9

28 43 1

30915 538 30 7

2Trn.3_ 46 3Trn12_

96**274 7

42 46 1

30523 536 31 7

7Trn3_

44 1Trn12_

9-68190 7 42 44 1

30523 536 32 7

2Trn3 43 3Trnol_

968148 7

42 43 1

30523 536 33 3

4Trn2_

a3 STrn1I_

206,818 3

1 43 1

31239 549 34 3

lTrn2..

19 STrn4_

203153 3

1 19

04 23510 549 35 31 1

n7 42 2Trnl 1 2625522 31 90 42 1

27376 339 Uenv 131 0.004 437 225 1 2-45 0.900 329 165 2.45 2.92 0.001 526 274 2.45 10-4 9.O09 120 46 245 2.45 0.009 394 201 2.45 446 0.001 335 169 2.45 3.04 9.001 465 257 2.45 859 9.000 516 2F9 2.45 9.68 0.001 490 254 2.45 6.31 0.000 526 274 2.45 1-3.40 0.001 526 274 2.45 10.49 0.007 426 216 2.45 5.40 0.001 530 277 2.45 10.7S

0.

020 530 277 2.45 10.76 0.016 526 274 2.45 10.49 0.000 526 274 2.45 10.49 0.001 526 274 2.45 10.49 0.000 529 276 2.45 10.69 0.405 529 276 2.45 1 0.89 0.000 528 276 2,45 10.63 O.901 543 264 2.45 11.71 0.908 543 284 2.45 11.71 0.011 543 284 2.45 11.71 0.00O 543 284 2.45 11.71 0.007 527 275 2.45 10.56 0.00i 528 276 2.45 1 3-63 0.013 528 276 2.45 10.63 0.002 536 260 2.45 11.19 0.0G'1 536 250 2.45 11.9 0.009 536 280 2.45 11.19 0.00O 53 250 2.45 11.12 0.001 549 287 2.45 12.18 0.005 549 287 2.45 12.18 0.030 339 171 2.45 3.12 0.000 Notes:

1. T0.:iethe maximum temperature of the two paired load states, sod represents the motal (nodal) temperature at the location being analyzed. This.

which is included as 7' in the'Trao sient M.aximum Temperatoras"' table above.

determined from the VESLFA output.

2. F.,

values computed using the low alloy steel equationo from Section 3.0 of Reference

161, with S' cons-ervathrelyaot to a maximum value of10.015, an~d the transformed strain rate conoservetively set to a minimum value of hn (0.001)

=-6.908 for ali load pairs.

-3. U.ý = [U x HVVC F_ x. % H!,C[ + [U x. NIWC F, x %S NINC1.

4.1 eI nd 72 represent the load number for LoGad 4#1 and Load #2, respectively, ind zl and s2 represent the --

tate number for each of thoze oapds!

Total. U 1 1.111 Overall Fen=

6.32

51. For each load pair, n, is.the number of available cycles for Load #1, n_ is the nomter of available cycles for Load 42. andunLe the available number of cyclen for the load pair (i.e.. the minimum of1n, and 0n).

File No.: 0801038.306 Revision: 1 Page 18 of 19 F0306-01RO

Structural Integrity Associates, Inc.

Table 9: Linearized Stress Files Compiled for VY-RO-StressResults.xls Filename Description VY RONTRAN1-S.csv VYRONTRAN2-S.csv VYRON_ TRAN3-S.csv VYRONTRAN4-S.csv VY RONTRAN5-S.csv VY RON TRAN6-S.csv VY RON TRAN7-S.csv VYRONTRAN8-S.csv VYRONTRAN9-S.csv VYRONTRAN1O-S.csv VY_ RON_ TRAN11-S.csv VYRONTRAN12-S.csv Transient 1 linearized stress Transient 2 linearized stress Transient 3 linearized stress Transient 4 linearized stress Transient 5 linearized stress Transient 6 linearized stress Transient 7 linearized stress Transient 8 linearized stress Transient 9 linearized stress Transient 10 linearized stress Transient 11 linearized stress Transient 12 linearized stress Note: All files are from the Reference [1] supporting computer files.

File No.: 0801038.306 Revision: I Page 19 of 19 F0306-0IRO ATTACHMENT 4 May 15, 2009 UNITED STATES OF AMERICA NUCLEAR REGULATORY COMMISSION Before the Atomic Safety and Licensing Board In the Matter of

))

Entergy Nuclear Vermont Yankee, LLC

)

Docket No. 50-271-LR and Entergy Nuclear Operations, Inc.

)

ASLBP No. 06-849-03-LR

)

(Vermont Yankee Nuclear Power Station)

)

DECLARATION OF GARY L. STEVENS Gary L. Stevens states as follows under penalty of perjury:

1. My name is Gary Lance Stevens. I am a Senior Associate at Structural Integrity Associates, Inc. ("SIA"). I have previously testified on behalf of Applicants Entergy Nuclear Vermont Yankee, LLC and Entergy Nuclear Operations, Inc.

("Entergy") in this proceeding.

2. I led the SIA team that prepared in January 2009 a set of Confirmatory Environmentally Assisted Fatigue Analyses ("CUFen Analyses") of the Core Spray

("CS") and Recirculation Outlet ("RO") reactor nozzles at the Vermont Yankee Nuclear Power Station ("VY") utilizing the same methodology and approach SIA had used in the Confirmatory. CUFen Analysis of another reactor nozzle, the feedwater nozzle. These analyses were reflected in the following calculations: Calculation 0801038.301, Revision 0, "Design Inputs and Methodology for ASME Code Fatigue Usage Analysis of Reactor Core Spray Nozzle;" Calculation No. 0801038.302, Revision 0, "Stress Analysis of Reactor Core Spray Nozzle;" Calculation No. 0801038.303, Revision 0, "Fatigue Analysis of Core Spray Nozzle;" Calculation No. 0801038.304, Revision 0, "Design Inputs and Methodology for ASME Code Fatigue Usage Analysis of Reactor

Recirculation Outlet Nozzle;" Calculation No. 0801038.305, Revision 0, "Stress Analysis of Reactor Recirculation Outlet Nozzle;" and Calculation No. 0801038.306, Revision 0, "Fatigue Analysis of Recirculation Outlet Nozzle" (collectively referred to as "the initial CS and RO confirmatory calculations"). I also led the SIA team that prepared, in March 2009, a set of revised Confirmatory CUFen calculations for the CS and RO nozzles, i.e.,

Calculation No. 0801038.302, Revision 1, "Stress Analysis of Reactor Core Spray Nozzle;" Calculation No. 0801038.303, Revision 1, "Fatigue Analysis of Reactor Core Spray Nozzle;" Calculation No. 0801038.304, Revision 1, "Design Inputs and Methodology for ASME Code Fatigue Usage Analysis of Reactor Recirculation Outlet Nozzle;" Calculation No. 0801038.305, Revision 1, "Stress Analysis of Reactor Recirculation Outlet Nozzle;" and Calculation No. 0801038.306, Revision 1, "Fatigue Analysis of Reactor Recirculation Outlet Nozzle."

3. Entergyprovided copies of the initial CS and RO confirmatory calculations to the Staff of the U.S. Nuclear Regulatory Commission ("NRC"), who performed a technical audit of the calculations. During the course of the audit, the NRC Staff asked whether the CUFen results for those nozzles were dependent on the value of heat transfer coefficient used in the analyses. SIA staff, under my supervision, performed a sensitivity analysis of the effect of variations in the nozzle comer heat transfer coefficients on the CUFen for the CS nozzle. The analysis was completed in March 2009.

4. A copy of a table prepared by SIA that tabulates the CUFen for the CS nozzle as a function of heat transfer coefficient is enclosed as Exhibit A. The table demonstrates that the effect of changes in the heat transfer coefficient on CUFen estimates is minimal.

While the sensitivity analysis focused on the CS nozzle, its results are applicable to the RO nozzle as well, since the forced convection heat transfer coefficients for the CS nozzle are higher (and therefore bound) the heat transfer coefficients calculated for the RO nozzle. Refer to Table 3 of Exh. E2-14 (VY-16Q-305, Rev. 0) at p. 12 (Hforced =

2713 Btu/hr-ftZ-OF at 300'F for Region 3 (near nozzle comer)) vs. Table 19 of 0801038.301, Rev. 0 at p. 40 (Hforced = 3921 Btu/hr-ft2 _°F for Region 9 (near nozzle comer)).

2

5. I declare under penalty of perjury that the foregoing is true and correct.

ary L. Stevens Executed on May 15, 2009 3

EXHIBIT A 0801038-307 CS Nozzle 1

2 CUF Difference Safe End Material 1(Inconel) 0.0001740 0.0001740 0.000000 Safe End Material 2 (S.S.)

0.0007422 0.0007422 0.000000 Blend Radius 0.0171387 0.0155733

-0.0015654

1. The results of 0801038.303.R0

2. For this iteration 0801038.302.RO ANSYS input files for all transients were modified by changing the Region 9 heat transfer coefficient to 500Btu/hr-fi?-°F for all flow cases. This change then forces Region 9, 10 and 11 to have a heat transfer coefficient of 500Btu/hr-ftf2 -F for all flow cases.

Region!!

Region tO Li J

Region 9 Region 8 713----

Region 5 Region Region 3

.on 2 Region 12 Region I

ML091420373

Text

6.0 REFERENCES

6.0 REFERENCES

6.0 REFERENCES

6.0 CONCLUSION

7.0 REFERENCES

6.0 CONCLUSION

7.0 REFERENCES

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