ML081820137

Prairie Island, Units 1 and 2 - License Amendment Request for Technical Specifications Changes to Allow Use of Westinghouse 0.422-inch OD 14x14 Vantage+ Fuel

ML081820137
Person / Time
Site:	Prairie Island
Issue date:	06/26/2008
From:	Wadley M D Nuclear Management Co
To:	Document Control Desk, Office of Nuclear Reactor Regulation
References
L-PI-08-047
Download: ML081820137 (293)

Text

Prairie Island Nuclear Generatincl Plant Operated by Nuclear Management Company, LLC L-PI-08-047 10 CFR 50.90 10 CFR 50.46 Document Control Desk U .S. Nuclear Regulatory Commission Washington, DC 20555-0001 Prairie Island Nuclear Generating Plant Units I and 2 Dockets Nos.

50-282 and 50-306 License Nos. DPR-42 and DPR-60 License Amendment Request for Technical Specifications Chanaes to Allow Use of Westinghouse 0.422-inch OD 14x14 VANTAGE+ Fuel Pursuant to 10 CFR 50.90, Nuclear Management Company, LLC (NMC) hereby requests amendments to the Operating Licenses and associated Technical Specifications (TS) for Prairie Island Nuclear Generating Plant (PINGP), Units 1 and 2, in support of the transition from 0.400-inch outer diameter (OD) VANTAGE+ (hereinafter referred to as 400V+) fuel to Westinghouse 0.422-inch OD VANTAGE+ (hereafter referred to as 422V+) fuel. This transition is planned beginning with Cycle 26 (Fall 2009) for Unit 1 and Cycle 26 (Spring 2010) for Unit 2 and will extend over three cycles for each Unit (Cycles 26 through 28).

The 422V+ fuel design is a modification of the physical structure of the Westing house Vantage+ (WCAP-12610-P-A) fuel design already in use at PINGP. The 422V+ fuel design is currently in use at Kewaunee Power Station, Point Beach Nuclear Plant, and R.

E. Ginna Nuclear Power Plant. Enclosure 1 describes the requested TS changes and provides mark-ups of affected TS and TS Bases pages, as well as revised TS pages. Enclosure 1 also provides a detailed discussion of the 422V+ fuel and the analyses, including evaluations of loss-of-coolant-accidents (LOCAs), non-LOCA transients, and pipe breaks inside containment conducted in support of this request. The results of the analyses, including evaluations show that PINGP Units 1 and 2 can operate safely and within their licensing bases with 422V+ fuel, including the transition to 422V+ fuel. Finally, Enclosure 1 provides NMC's analysis with respect to the issue of No Significant Hazards Consideration.

1717 Wakonade Drive East Welch, Minnesota 55089-9642 Telephone:

651.388.1 121 Document Control Desk Page 2 In addition to the TS changes described in Enclosure 1, NMC requests review and approval for this application of the following analyses. (All section numbers refer to Attachment 4 of Enclosure 1

.): Four re-analyzed non-LOCA events (Sections 5.1 .I, 5.1.8, 5.1.9 and 5.1.14). Small Break LOCA (Section 5.2.2).

LOCA Containment Peak Pressure Response (Sections 5.3.1 and 5.3.2). Steam Line Break inside Containment (Section 5.4.1).

As part of the fuel transition to 422V+, the large break loss of coolant accident (LBLOCA) analyses for Units 1 and 2, as previously approved for PINGP, were re-analyzed using the same methodology. These new analyses bound the current 400V+ fuel and the 422V+ fuel or any combination of the two. The results of these re-analyses are presented in Attachment 4 to Enclosure 1 and meet all the acceptance criteria delineated in 10 CFR 50.46. These re-analyses result in a projected Peak Cladding Temperature (PCT) for Unit 1 of 1765°F and a PCT of 1623°F for Unit 2 which both exceed a 50 degree change from the existing analysis results. Therefore, this application for amendment is considered a notification for a significant change in peak clad temperature per the requirements of 10 CFR 50.46. Since this was a re-analysis using NRC- approved methods and meets all acceptance criteria, no re-analysis schedule is necessary or being provided. These new reference values will become effective upon implementation of this license amendment for each Unit. The radiological source term for 422V+ fuel was evaluated for its potential impact on dose consequences relative to the existing 400V+ fuel, using the whole bodylthyroid dose and TEDE models.

The evaluation determined that the potential dose impact of the 422V+ fuel is essentially the same as that for the 400V+ fuel. For information purposes, it should be noted that an update to the existing radiological accident analyses has been prepared and is being submitted under separate cover for NRC approval. The re-analyses utilize a source term that reflects a bounding core inventory for either 400V+ or 422V+ fuel. NMC requests approval of this LAR by June 30, 2009 with 90 days to implement the associated changes on Unit 1 to support the Unit 1 Cycle 26 refueling outage.

Document Control Desk Page 3 Summaw of Commitments This letter contains no new commitments and no revisions to existing commitments.

If you have any questions or require additional information regarding this request, please contact Mr. Lenny Sueper at (612) 330-6917, Leonard.Sueper@nmcco.com. I declare under penalty of perjury that the foregoing is true and correct.

Executed on, JfdN 8 6 20@ Michael D. Wadley Site Vice president, Prairie Island Nuclear Generating Plant, Units 1 and 2 Nuclear Management Company, LLC Enclosure 1 : Evaluation of the Change cc: NRC Regional Administrator NRC Project Manager NRC Resident Inspector(s) State of Minnesota Page 1 of 9 Evaluation of the Change License Amendment Request to the Prairie Island Generating Plant Technical Specifications for Use of West inghouse 14X14 422 VANTAGE+ Fuel 1. Background and Summary

2. Detailed Description of Requested Amendment

2.1 Change

to Fuel Design

2.2 Requested

Technical Specificat ion Changes and Justifications 2.3 Changes to TS Bases 2.4 Supporting Analyses and Evaluations 2.5 Radiological Source Term

3. Technical Evaluation

3.1 Change

to Fuel Design

3.2 Technical

Specification Changes

3.3 Supporting

Analyses and Evaluations 4. Regulatory Evaluation 4.1 Applicable Regulatory Requirements/Criteria

4.2 Precedent

4.3 Significant Hazards Consideration Analysis

4.4 Conclusions

5. Environmental Considerations

6. References Attachments:

1. Technical Specification Page Markups 2. Bases Page Markups 3. Retyped Technical Specification Pages

4. Prairie Island Units 1 and 2 422V+ Reload Transition Licensing Report Page 2 of 9 1. Background and Summary This evaluation supports a request to amend Operating Licenses DPR-42 and DPR-60 for Prairie Island Nuclear Generating Plant (PINGP). This license amendment request (LAR) seeks to amend Operating Licenses DPR-42 and DPR-60 and associated Technical Specifications (TS) for PINGP Units 1 and 2, respectively, in cluding review and approval for the use of 422V+ fuel, related and conforming changes to the TS and supporting analyses and methods, as described below. 2. Detailed Description of Requested Amendment The Prairie Island Nuclear Generating Plant (PINGP) plans to refuel and operate with upgraded Westinghouse fuel , commencing with Cycle 26 for both Units 1 and 2. The upgraded fuel is 0.422-inch outer diameter (OD), 14x14 VANTAGE+ fuel, hereafter referred to as 422V+. This fuel is similar to the 422V+ fuel assemblies in operat ion at the Point Beach Nuclear Power Plant, Kewaunee Power Stat ion, and R. E. Ginna Nu clear Power Plant. The current fuel in operation at PINGP is 0.400-inch OD, 400V+. A comparison of the mechanical properties of 400V+

and 422V+ fuel is given in Section 2 of Attachment 4. 2.1 Change to Fuel Design PINGP requests review and approval for the use of 422V+ fuel in Units 1 and 2, including the use of a mix ed core beginning with partial core replacement in Refueling Cycle 26 fo r each Unit and transitioning to all 422V+ fuel over the succeeding two cycles for each Unit. As shown in

Section 4 of Attachment 4, the 422V+

fuel provides additional departure from nucleate boiling ratio (DNBR) ma rgin compared to the 400V+ fuel. 2.2 Requested Technical Specification Changes and Justifications A description of the proposed TS changes a nd their justifications is provided below. TS 2.1.1.1, "Reactor Core Safety Limit" Remove the fuel-specific wording "for OFA fuel" from the departure from nucleate boiling ratio (DNBR) WR B-1 DNB correlation safety limit. Justification

TS 2.1.1.1 is changed to remove the s pecific reference to OFA fuel since the limit is also applicable to 422V+fuel. This change is consistent with NUREG-1431, Revision 3, "Standard Technical Specifications Westinghouse Plants."

P a g e 3 o f 9

• T S 2.1.1.2 , R e a c t o r C o r e S a f e t y L i m i t s C h a n g e t h e e x i s t i n g r e q u i r e m e n t t o r e a d: "T h e p e a k f u e l c e n t e r l i n e t e m p e r a t u r e s h a l l b e m a i n t a i n e d a s f o l l o w s: a) < 5 , 0 8 0 F , d e c r e a s i n g b y 5 8 F p e r 1 0 , 0 0 0 M W D/M T U o f b u r n u p f o r f u e l c o n t a i n i n g U O 2. b) < (5 , 0 8 0 F m i n u s 6.7 5 F p e r w/0 G d 2 O 3), d e c r e a s i n g b y 5 8 F p e r 1 0 , 0 0 0 M W D/M T U o f b u r n u p f o r f u e l c o n t a i n i n g g a d o l i n i a". J u s t i f i c a t i o n: T S 2.1.1.2 i s m o d i f i e d t o s p e c i f y t h e f u e l c e n t e r l i n e m e l t t e m p e r a t u r e f o r f u e l w i t h a h o m o g e n e o u s p o i s o n g a d o l i n i a , a n d t o b r i n g t h e P I N G P f u e l c e n t e r l i n e m e l t t e m p e r a t u r e f o r n o n-g a d o l i n i a f u e l i n l i n e w i t h t h e W e s t i n g h o u s e s t a n d a r d. T h e f u e l c e n t e r l i n e t e m p e r a t u r e m e l t i n g l i m i t s f o r b o t h g a d o l i n i a a n d n o n-g a d o l i n i a f u e l d e s i g n s a r e r e f e r e n c e d i n W C A P-8 7 2 0 , A d d e n d u m 3. T h e m e l t i n g l i m i t s a r e i n t e g r a t e d i n t o t h e p e a k f u e l c e n t e r l i n e t e m p e r a t u r e e v a l u a t i o n m e t h o d o l o g y u s e d t o c o n f i r m t h a t t h e f u e l c e n t e r l i n e m e l t d e s i g n c r i t e r i a a r e m e t. T h e n o n-g a d o l i n i a f u e l t e m p e r a t u r e c h a n g e i s c o n s i s t e n t w i t h N U R E G-1 4 3 1 , R e v i s i o n 3. F u r t h e r j u s t i f i c a t i o n i s p r o v i d e d i n W C A P-1 2 4 8 8-A , " W e s t i n g h o u s e F u e l C r i t e r i a E v a l u a t i o n P r o c e s s ," O c t o b e r 1 9 9 4.

• S R 3.5.1.4 , A c c u m u l a t o r s I n c r e a s e t h e s u r v e i l l a n c e r e q u i r e m e n t m i n i m u m b o r o n c o n c e n t r a t i o n f r o m 1 9 0 0 p p m t o 2 3 0 0 p p m. J u s t i f i c a t i o n: A c h a n g e t o i n c r e a s e t h e m i n i m u m a c c u m u l a t o r b o r o n c o n c e n t r a t i o n f r o m a T S l i m i t o f 1 9 0 0 p p m t o a l i m i t o f 2 3 0 0 p p m e n s u r e s t h a t t h e c o r e r e m a i n s s u b-c r i t i c a l d u r i n g c o r e r e f l o o d f o l l o w i n g a l a r g e-b r e a k l o s s-o f-c o o l a n t-a c c i d e n t (L B L O C A). A s p a r t o f t h e t r a n s i t i o n f r o m 4 0 0 V+ t o 4 2 2 V+ f u e l , t h e S a f e t y A n a l y s i s C h e c k l i s t l i m i t s w e r e r e c o n f i r m e d , i n c l u d i n g , a c h e c k t o s h o w t h a t t h e c o r e r e m a i n s s u b-c r i t i c a l i n t h e s h o r t t e r m d u r i n g t h e c o r e r e f l o o d p e r i o d o f a L B L O C A. C a l c u l a t i o n s s h o w e d t h a t a n i n c r e a s e i n a c c u m u l a t o r b o r o n c o n c e n t r a t i o n t o 2 3 0 0 p p m w a s s u f f i c i e n t t o p r e v e n t a r e t u r n t o c r i t i c a l i t y. T h e p r o p o s e d m i n i m u m a c c u m u l a t o r b o r o n c o n c e n t r a t i o n w i l l r e m a i n b e l o w t h e m a x i m u m b o r o n s o l u b i l i t y c o n c e n t r a t i o n a s s u m e d i n t h e s u p p o r t i n g a n a l y s i s (3 5 0 0 p p m). T h i s c h a n g e i s c o n s i s t e n t w i t h t h e m i n i m u m a c c u m u l a t o r b o r o n c o n c e n t r a t i o n i n p l a n t s p e c i f i c T e c h n i c a l S p e c i f i c a t i o n s f o r o t h e r t w o-l o o p p l a n t s , i n c l u d i n g P o i n t B e a c h N u c l e a r P l a n t a n d R. E. G i n n a N u c l e a r P o w e r P l a n t.

Page 4 of 9 TS 4.3.1.1 b, Fuel Storage - Criticality Change "Reference 1" to read "USAR Section 10.2." Justification

The change to delete Reference 1 is ju stified by referring directly to Section 10.2 of the USAR, which contains this information. Reference 1 in Specification 4.3.1.1 b is Refe rence 45 in Section 10 of the USAR.

This change is consistent with t he Fuel Storage Design Features Specification in NUREG-1431, Revisi on 3, and the Fuel Storage Design Features Specifications in Plant Specific Technical Specifications for plants such as: Beaver Valley Units 1 and 2, Callaway, Point Beach, and various other Westinghouse plants. TS 4.3.1.1 c, Fuel Storage - Criticality Change "Reference 1" to read "USAR Section 10.2." Justification

The change to delete Reference 1 is ju stified by referring directly to Section 10.2 of the USAR, which contains this information. Reference 1 in Specification 4.3.1.1 c is Refe rence 45 in Section 10 of the USAR.

This change is consistent with t he Fuel Storage Design Features Specification in NUREG-1431, Revisi on 3, and the Fuel Storage Design Features Specifications in Plant Specific Technical Specifications for plants such as: Beaver Valley Units 1 and 2, Callaway, Point Beach, and various other Westinghouse plants. TS 4.3.1.2 b, Fuel Storage - Criticality Change "Reference 2" to r ead "USAR Section 10.2." Justification

The change to delete Reference 2 is ju stified by referring directly to Section 10.2 of the USAR, which contains this information. Reference 2 in Specification 4.3.2.1 b is Refe rence 39 in Section 10 of the USAR.

This change is consistent with t he Fuel Storage Design Features Specification in NUREG-1431, Revisi on 3, and the Fuel Storage Design Features Specifications in Plant Specific Technical Specifications for plants such as: Beaver Valley Units 1 and 2, Callaway, Point Beach, and various other Westinghouse plants. TS 4.3.3, Fuel Storage - Capacity Change "Ref. 3" to read "USAR Section 10.2." Justification

The change to delete Reference 3 is ju stified by referring directly to Section 10.2 of the USAR, which contains this information.

Page 5 of 9 TS 4.3, References Delete References 1, 2 and 3. Justification

The change to delete References 1, 2 and 3 is conforming to the above changes and is justified by referring di rectly to Section 10.2 of the USAR in the text of the TS. 2.3 Changes to TS Bases Several changes will be made to the Bases for the above TS changes.

Markups of the affected pages are provided in Attachment 2. 2.4 Supporting Analyses and Evaluations Attachment 4 summarizes the eval uations and analyses that were performed to confirm the acceptable use of 422V+ fuel.

Each of the non-LOCA transients listed in Table 5.1-1 of Attachment 4 was analyzed or evaluated in support of the 422V+ Transition Program. The analyses and evaluations are discussed in detail in Section 5.1 of . The LBLOCA analyses for Units 1 and 2 were both re-analyzed. The results of these re-analyses are presented in Attachment 4, Section 5.2.1.

The small break LOCA (SBLOCA) analysis was re-analyzed, as discussed in Attachment 4, Section 5.2.2.

Primary piping breaks (LOCAs) inside c ontainment were re-analyzed, as discussed in Attachment 4, Section 5.3.

As the following events required re-analysis, NMC requests NRC review and approval. (All section numbers refer to Attachment 4): Four re-analyzed non-LOCA events (S ections 5.1.1, 5.1.8, 5.1.9 and 5.1.14). Small Break LOCA (Section 5.2.2). LOCA Containment Peak Pressure Response (Sections 5.3.1 and 5.3.2). Steam Line Break inside Containment (Section 5.4.1). 2.5 Radiological Source Term The radiological source term for 422V+ fuel was evaluated for its potential impact on dose consequences relative to the existing 400V+ fuel, using the whole body/thyroid dose and TEDE models.

The evaluation determined that the potential dose impact of the 422V+ fuel is essent ially the same as that for the 400V+ fuel. For information purposes, it should be noted that an update to the existing radiological accident analyses has been prepared and is being submitted Page 6 of 9 under separate cover for NRC approval. The re-analyses utilize a source term that reflects a bounding core inv entory for either 400V+ or 422V+ fuel. 3. Technical Evaluation 3.1 Change to Fuel Design The 400V+ fuel currently installed in PINGP Units 1 and 2 will be replaced with 422V+ fuel. Both fuel types are manufactured by Westinghouse

Nuclear. The comparison between 400V+, and 422V+ fuel is described fully

in Section 2 of Attachment 4.

Analyses and evaluations of the effe cts of the 422V+ fuel change on the reactor internals are described in Section 6 of Attachment 4. 3.2 Technical Specification Changes The justifications for the proposed TS changes are provided in Section 2.2 of this enclosure. 3.3 Supporting Analyses and Evaluations Non-LOCA transients are discussed in Section 5.1 of Attachment 4. The LBLOCA analyses are presented in Section 5.2.1 of Attachment 4. The SBLOCA analysis is discussed in Se ction 5.2.2 of A ttachment 4.

The Unit 1 and Unit 2 Main Steam Line Break Containment Response analyses are discussed in Sectio n 5.4 of Attachment 4.

Bounding LOCA containment response analyses for PINGP Unit 1 and 2 with 422V+ fuel are presented in section 5.3.1 and 5.3.2 of Attachment 4 for the long term mass and energy calculat ions and the containment response, respectively. 4. Regulatory Evaluation 4.1 Applicable Regulatory Requirements/Criteria As described in the USAR, PINGP was designed and constructed to comply with the Atomic Energy Commission General Design Criteria (AEC GDC) as proposed on July 10, 1967. The constr uction of PINGP was significantly complete prior to issuance of curr ent 10 CFR 50, Appe ndix A, General Design Criteria (GDC). Similarly, PINGP was not designed or constructed with the benefit of the Standard Review Plan (SRP

). Although PINGP is not a GDC or SRP plant, the following GDCs and SRP sections provide guidance in assessing the fuel transition. GDC 16 and GDC 50: the peak calculated containment pressure should be less than the containment design pre ssure of 46 psig, considering the most severe single failure GDC 38 and GDC50: the calculated pressure at 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> should be less than 50 percent of the peak calculated value. (This is related to the criteria for doses at 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />.)

Page 7 of 9 Applicable regulatory requirements include 10 CFR 50, 10 CFR 50.46, 10 CFR 50 Appendix K, and 10 CFR 100.

The following SRP sections have been app lied by NRC in assessing similar fuel design and utilization changes. SRP 4.2, "Fuel System Design" SRP 4.3, "Nuclear Design" SRP 4.4, "Thermal and Hydraulic Design" SRP 6.2.1.3, "Mass and Energy Re lease Analysis for Postulated Loss-of-Coolant SRP 6.3, "Emergency Core Cooling System" SRP Chapter 15, "Accident Analysis," various Review Plans, as applicable 4.2 Precedent The 422V+ fuel design has been reviewed and approved by NRC and is currently in use at Kewaunee Power Station, Point Beach Nuclear Power Plant, and R. E. Ginna Nuclear Power Plant (See References 1 through 3).

Thus, current operating experience s upports operation of the 422V+ fuel design at PINGP. 4.3 Significant Hazards Consideration Analysis The proposed amendment would make changes the Technical

Specifications that are c onforming or related to a change in fuel type from Westinghouse 0.400-inch OD Vantage+

fuel (400+) to Westinghouse 0.422-inch OD Vantage+ (422V+) fuel.

NMC has evaluated whether or not a si gnificant hazards consideration is involved with the propos ed amendment by focusing on the three standards set forth in 10 CFR 50.92, "Issuance of amendment," as discussed below: 1. Does the proposed amendment invo lve a significant increase in the probability or consequences of an accident previously evaluated? Response: No. The requested amendment is related to a change in the reload fuel design.

The design criteria for the reload fuel are consistent with those for the existing fuel and ensure that the reload fuel is compatible on the basis of coolant flow and neutronic characteri stics, as well as DNB and peak cladding temperature requirements. The reload fuel design also ensures mechanical compatibility with the existing fuel, reactor core, control rods, steam supply system, and fuel handling tools and system. The reactor fuel and its analysis are not accident initiators. Therefore, the change in reload fuel design does not affect acci dent or transient initiation.

Page 8 of 9 The minimum boron accumulator concent ration is also not an accident initiator. The proposed change to the minimum accumulator boron concentration Technical Specification limit ensures that the plant will continue to operate in a manner that provides acceptable levels of protection for health and safety of the public. Further, all design basis

accidents and transients affected by t he fuel upgrade were re-analyzed or evaluated, and the results for the existing fuel remain bounding for the transition to 422V+ fuel.

Therefore, the proposed c hanges do not involve a sign ificant increase in the probability or consequences of an a ccident previously evaluated. 2. Does the proposed amendment create the possibility of a new or different kind of accident from any accident previously evaluated? Response: No.

Use of the 422V+ fuel is consistent with current plant design bases and does not adversely affect any fission product barrier, nor does it alter the safety function of safety significant systems, structures and components or their roles in accident preventi on or mitigation. The operational characteristics of 422V+ fuel are bounded by the safety analyses (Attachment 4). The 422V+ fuel design performs within existing fuel design limits. The proposed change to the minimum accumulator boron concentration Technical Specification limit ensures that the plant wil l continue to operate in a manner that provides acceptable leve ls of protection for health and safety of the public. Further, all design basi s accidents and transients affected by the fuel upgrade were re-analyzed or evaluated, and the results for the existing fuel remain bounding fo r the transition to 422V+ fuel.

No equipment additions or modifications are included with the proposed change, and no changes to plant operating procedures are proposed.

Therefore, the proposed c hange does not create the po ssibility of a new or different kind of accident from any accident previously evaluated. 3. Does the proposed amendment in volve a significant reduction in a margin of safety? Response: No. The proposed changes do not alter the manner in which applicable design basis limits are determined, nor do they result in exceeding existing design basis limits. Thus, all licensed safety margins are maintained.

Therefore, the proposed c hanges do not involve a significant reduction in the margin of safety.

Based on the above, NMC concludes that the proposed amendment does not involve a significant hazards cons ideration under the st andards set forth Page 9 of 9 in 10 CFR 50.92(c), and, accordingly, a finding of "no significant hazards consideration" is justified.

4.4 Conclusions

Based on the considerations discussed above, (1) there is reasonable assurance that the health and safety of the public will not be endangered by operation in the proposed m anner, (2) such activiti es will be conducted in compliance with the Commission's regulations, and (3) the issuance of the amendment will not be inimical to the common defense and security or to the health and safety of the public. 5. Environmental Considerations A review has determined that the proposed amendment would change a requirement with respect to installation or use of a facility component located within the restricted area, as defined in 10 CFR 20, or would change an inspection or surveillance requirement. However, the proposed amendment does not involve (i) a signific ant hazards consideration, (ii) a significant change in the types or a si gnificant increase in the amounts of any effluents that may be released offsite, or (iii) a significant increase in individual or cumulative occupational radiation exposure. Accordingly, the proposed amendment meets the eligibility criterion for categorical exclusion set forth in 10 CFR 51.22(c)(9). Therefore, pursuant to 10 CFR 51.22(b), no environmental impact statement or environmental assessment need be prepared in connection with the proposed amendment.

6. References 1) NRC Letter, G. P. Hatchett to M. B. Sellman, "Point Beach Nuclear Plant, Units 1 and 2 - Issuance of Amendments Re: Design and Operation of Fuel Cycles with Upgraded Westinghouse Fuel," February 8, 2000 (ADAMS Accession ML003683159). 2) NRC Letter, John G. Lamb to Thomas Coutu, "Kewaunee Nuclear Power Plant - Issuance of Amendment,"

April 4, 2003 (ADAMS Accession ML030940276). 3) NRC Letter, P. D. Milano to M. G.

Korsnick, "R. E. Ginna Nuclear Power Plant - Amendment Re: 16.8 Percent Power Uprate," July 11, 2006 (ADAMS Accession ML061380103).

Attachment 1, Enclosure 1 Technical Specification Page Markups TS 2.0-1 TS 3.5.1-2 TS 4.0-2 through TS 4.0-4

6 pages follow SLs 2.0 Prairie Island Unit 1 - Amendment No. 158, 162 Units 1 and 2 2.0-1 Unit 2 - Amendment No. 149 , 153 2.0 SAFETY LIMITS (SLs) 2.1 SLs 2.1.1 Reactor Core SLs In MODES 1 and 2, the combination of THERMAL POWER, Reactor Coolant System (RCS) highest loop av erage temperature, and pressurizer pressure shall not exceed the limits sp ecified in the COLR and the following SLs shall not be exceeded:

2.1.1.1 The departure from nucleate boiling ratio (DNBR) shall be maintained > 1.17 for WRB-1 DNB correlation for OFA fuel. 2.1.1.2 The peak fuel centerline temperature shall be maintained

< 4700 F. 2.1.2 RCS Pressure SL In MODES 1, 2, 3, 4, and 5, the RCS pressure shall be maintained <

2735 psig. 2.2 SL Violations 2.2.1 If SL 2.1.1 is violated, restore compliance and be in MODE 3 within 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br />.

2.2.2 If SL 2.1.2 is violated:

2.2.2.1 In MODE 1 or 2, restore compliance and be in MODE 3 within 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br />. 2.2.2.2 In MODE 3, 4, or 5, restor e compliance within 5 minutes.

Insert A Page 3 of 13 INSERT A:

2.1.1.2 The peak fuel centerline temperature sh all be maintained as follows: a) < 5080 ºF, decreasing by 58 ºF per 10,000 MWD/MTU burnup, for fuel containing UO2; b) < (5080 ºF minus 6.75 ºF per w/o Gd 2 O 3), decreasing by 58 ºF per 10,000 MWD/MTU burnup , for fuel containing gadolinia.

Accumulators 3.5.1 Prairie Island Unit 1 - Amendment No. 158 Units 1 and 2 3.5.1-2 Unit 2 - Amendment No. 149 ACTIONS (continued) CONDITION REQUIRED ACTION COMPLETION TIME D. Two accumulators inoperable.

D.1 Enter LCO 3.0.3.

Immediately

SURVEILLANCE REQUIREMENTS SURVEILLANCE FREQUENCY SR 3.5.1.1 Verify each accumula tor isolation valve is fully open.

12 hours1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br /> SR 3.5.1.2 Verify borated wate r volume in e ach accumulator is > 1250 cubic feet (25%) and <

1290 cubic feet (91%).

12 hours1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br /> SR 3.5.1.3 Verify nitroge n cover pressure in each accumulator is

> 710 psig and <

770 psig.

12 hours1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br /> SR 3.5.1.4 Verify boron concen tration in each accumulator is 1900 ppm. 2300 ppm.

31 days

SR 3.5.1.5 Verify power is removed from each accumulator isolation valve operator when RCS pressure is

> 2000 psig.

31 days Design Features

4.0 Prairie

Island Unit 1 - Amendment No. 158 172 Units 1 and 2 4.0-2 Unit 2 - Amendment No. 149 162 4.0 DESIGN FEATURES (continued) 4.3 Fuel Storage

4.3.1 Criticality

4.3.1.1 The spent fuel storage racks are designed and shall be maintained with: a. Fuel assemblies having a maximum U-235 enrichment of 5.0 weight percent;

b. k eff < 1.0 if fully flooded with unborated water, which includes an allowance for uncertainties as described in Reference 1; USAR Section 10.2;

c. k eff < 0.95 if fully flooded with water borated to 730 ppm, which includes an allowance fo r uncertainties as described in Reference 1; USAR Section 10.2;

d. A nominal 9.5 inch center to center distance between fuel assemblies placed in th e fuel storage racks;

e. New or spent fuel assemblies with a combination of discharge burnup, initial enrichment and decay time in the "unrestricted range" of Figure 3.7.17-1 may be allowed unrestricted storage in the fuel storage racks; and

f. New or spent fuel assemblies w ith a combination of discharge burnup, initial enrichment and decay time in the "restricted range" of Figure 3.7.17-1 will be stored in compliance with Figures 4.3.1-1 through 4.3.1-4.

Design Features

4.0 Prairie

Island Unit 1 - Amendment No. 158 Units 1 and 2 4.0-3 Unit 2 - Amendment No. 149

4.0 DESIGN

FEATURES 4.3 Fuel Storage (continued) 4.3.1.2 The new fuel storage racks are designed and shall be maintained with: a. Fuel assemblies having a maximum U-235 enrichment of 5.0 weight percent;

b. k eff < 0.95 if fully flooded w ith unborated water, which includes an allowance for un certainties as described in Reference 2 ; USAR Section 10.2;

c. k eff < 0.98 if accidentally filled with a low density moderator which resulted in optimum low density moderation conditions; and

d. A nominal 21 inch center to center distance between fuel assemblies placed in the storage racks.

4.3.1.3 Fuel will not be inserted into a TN-40 spent fuel cask in the pool unless a minimum boron concentration of 1800 ppm is present.

The 1800 ppm will ensure that k eff for the spent fuel cask, including statistical unc ertainties, will be < 0.95 for all postulated arrangements of fuel within the cask. The criticality analyses for the TN-40 spent fuel storage cask were based on fresh fuel enriched to 3.85 weight percent U-235.

4.3.2 Drainage

The spent fuel storage pool is designe d and shall be maintained to prevent inadvertent draining of the pool below elevation 727 4 (Mean Sea Level).

Design Features

4.0 Prairie

Island Unit 1 - Amendment No. 172 180 Units 1 and 2 4.0-4 Unit 2 - Amendment No. 162 170 4.0 DESIGN FEATURES 4.3 Fuel Storage (continued)

4.3.3 Capacity

The spent fuel storage pool is designed and shall be maintain ed with a storage capacity limited to no more than 1386 fuel assemb lies not including those assemblies which can be retu rned to the reactor. The southeast corner of the small pool serves as the spent fuel cas k lay down area. To facilitate plant evolutions, four additional storage rack s, with a combined capacity of 196, may be temporarily installed in the cask lay down area to provide a total of 1582 storage locations (Ref.

3). USAR Section 10.2; REFERENCES

1. "Prair ie Island Units 1 and 2 Spent Fuel Pool Criticality Analysis", WCAP-16517-NP, Revision 0, Westinghouse Electric Company, November 2005. 2. "Criticality Analysis of the Prairie Island Units 1 & 2 Fresh and Spent Fuel Racks", Westinghouse Commercial Nuclea r Fuel Division, February 1993. 3. USAR, Section 10.2.

Enclosure, Attachment 2 Bases Page Markups B 3.5.1-2 B 3.5.1-4

B 3.5.1-5

B 3.7.2-2 B 3.7.16-2

B 3.7.16-3 B 3.7.16-4

B 3.7.16-5

B 3.7.17-2

B 3.7.17-4

B 3.7.17-8

11 pages follow

Accumulators B 3.5.1 Prairie Island Unit 1 - Amendment No.

Revision 158 Units 1 and 2 B 3.5.1-2 Unit 2 - Amendment No.

Revision 149 BASES BACKGROUND Each accumulator is piped into an RCS cold leg via an accumulator (continued) line and is isolated from the RCS by a motor operated isolation valve and two check valves in series. The motor operated isolation valves are MV 32071 and MV 32072 (Un it 2 - MV 32174 and MV 32175) (Westinghouse valve numbers 8 800A and 8800B respectively for both units).

The accumulator size, water volume, and nitrogen cover pressure are selected so that one of the two accu mulators is suffici ent to partially cover the core before significant clad melting or zirconium water reaction can occur following a LOCA.

The need to ensure that one accumulator is adequate for this function is consistent with the large break LOCA assumption that the entire contents of one accumulator will be lost via the RCS pipe break during the blowdown phase of the LOCA.

APPLICABLE The accumulators are assume d OPERABLE in bo th the large and SAFETY small break LOCA analyses at full power (Ref. 1). These are the ANALYSES Design Basis Accidents (DBAs) that establish the acceptance limits for the accumulators. Reference to the analyses for these DBAs is used to assess changes in the accu mulators as they relate to the acceptance limits.

In performing the LOCA calculations, conservative assumptions are made concerning the availability of ECCS flow. In the early stages of a large break LOCA, with or wi thout a loss of offsite power, the accumulators provide the sole sour ce of makeup water to the RCS. The assumption of loss of offsite power is required by regulations and conservatively imposes a delay wherein the EC CS pumps cannot deliver flow until the emergency dies el generators start, come to rated speed, and go through their tim ed loading sequence. In cold leg break scenarios, the entire contents of one accumulator are assumed to be lost through the break.

Accumulators B 3.5.1 Prairie Island Unit 1 - Revision 189 Units 1 and 2 B 3.5.1-4 Unit 2 - Revision 189 BASES APPLICABLE c. The amount of hydrogen generated by fuel element cladding SAFETY that reacts chemically with water or steam does not exceed an ANALYSES amount corresponding to interaction of 1% of the total amount (continued) of Zircaloy in the reactor; and

d. The core remains amenable to cooling during and after the break. Since the accumulators discharge during the blowdown phase of a LOCA, they do not contribute to the long term cooling requirements of 10 CFR 50.46.

The large break LOCA analysis considers a range of accumulator water volumes based on minimum and maximum volumes of 1250 cubic feet (25% indicated level) a nd 1290 cubic feet (91% indicated level).

The contained water volume is the same as the deliverable volume for the accumulators, since the accumulators are emptied, once discharged. For large breaks, an increase in water volume can be either a peak clad temperature penalty or benefit, depending on downcomer filling and subsequent spill through the break during the core reflooding portion of the transi ent. Prairie Island is a two loop plant with Upper Plenum Injecti on (UPI) LOCA analyses. For UPI plant small breaks, a decrease in water volume is a peak clad temperature penalty; thus, a minimum contained water volume is assumed. For small breaks, a nominal accumulator water volume is assumed due to lack of a consistent conservative direction. Both large and small break analyses use a nominal accumulator line water volume from the accumula tor to the ch eck valve.

The minimum boron concentration setpoint is used in the post LOCA boron concentration calcula tion. The calculation is performed to assure reactor subcriticality in a post LOCA environment including both long term sump conditions and short term reflood conditions. Of particular interest is the large break LOCA, since no credit is taken for c ontrol rod assembly insertion. A reduction in the accumulator minimum boron concentration would produce a subsequent Accumulators B 3.5.1 Prairie Island Unit 1 - Revision 174 Units 1 and 2 B 3.5.1-5 Unit 2 - Revision 174 BASES APPLICABLE reduction in the available containment sump concentration for SAFETY post LOCA shutdown, a reduction in available core boron ANALYSES concentrations during reflood and an increase in the maximum (continued) sump pH. For conservatism, the accumulators are assumed at a conservatively high boron concentration in the boron build up analyses.

The small break LOCA analyses are performed at the minimum nitrogen cover pressure, since sensitivity analyses have demonstrated that higher nitrogen cover pressure results in a computed peak clad temperature be nefit. The large break analyses utilize the nominal nitrogen cover pressure as per approved methods (Ref. 1). The maximum nitrogen cover pressure limit prevents accumulator relief valve actuati on, and ultimately preserves accumulator integrity.

The effects on containment mass and energy releases from the accumulators are accounted for in the appropriate analyses (Refs. 1 and 2).

The accumulators satisfy Criterion 3 of 10 CFR 50.36(c)(2)(ii).

LCO The LCO establishes the minimum conditions required to ensure that the accumulators are available to accomplish their core cooling safety function following a LOCA.

Two accumulators are required to ensure that 100% of the conten ts of one accumulator will reach the core during a LOCA. This is c onsistent with the assumption that the contents of one accumulator sp ill through the break. If less than one accumulator is injected during the blowdown phase of a LOCA, the ECCS acceptance criteria of 10 CFR 50.46 could be violated.

For an accumulator to be co nsidered OPERABLE, the motor-operated isolation valve must be fully open, power removed above 2000 psig, and the limits established in the SRs for contained volume, boron concentration, and nitrogen cover pressure must be met.

MSIVs B 3.7.2 Prairie Island Unit 1 - Amendment No. 158 Units 1 and 2 B 3.7.2-2 Unit 2 - Amendment No. 149 BASES BACKGROUND In addition to the fast-closing stop valve, each steam line has a (continued) downstream non-return chec k valve (NRCV). The four valves (one MSIV and one NRCV in each of two lines) preven t blowdown of more than one steam generator for any break location even if one valve fails to close. A description of the MSIVs and NRCVs is found in the USAR (Ref. 1).

APPLICABLE The design basis of the MSIVs is established by the containment SAFETY analysis for the large steam line break (SLB) inside containment, ANALYSES discussed in the USAR (Ref. 2). The design precludes the blowdown of more than one steam generator, assuming a single active component failure (e.g., the failure of one MSIV or NRCV to close). The limiting pressure case for the containment analysis is the main steam line break (MSLB) inside containment at 30% power , with offsite power available following turbine trip, and failure of a safeguards train. At lower power, the steam generator inventory and temperature are at their maximum, maximizing the analyzed mass and energy release to the containment.

With the most reactive rod cluster control assembly assume d stuck in the fully withdrawn position, there is an increased possi bility that the co re will become critical and return to power. The co re is ultimately shut down by the boric acid injection delivered by the Emergency Core Cooling System.

The analysis of several different SLB events are performed to demonstrate that the acceptance criteria listed in the USAR are satisfied.

Events evaluated include:

a. Containment response due to a large SLB inside of containment;

b. Core response due to a large SLB inside of containment;

Fuel Storage Pool Boron Concentration B 3.7.16 Prairie Island Unit 1 - Revision 194 Units 1 and 2 B 3.7.16-2 Unit 2 - Revision 194 BASES (continued)

APPLICABLE The spent fuel pool criticality analysis (Ref. 4 and 5) addresses all the fuel types currently stored in the spent fuel pool and in use in the SAFETY reactor. The fuel types consid ered in the analysis include the ANALYSES Westinghouse Standard (STD), OFA, and Vantage Plus designs (both 0.400" and 0.422" O.D. designs), and the Exxon fuel assembly types in storage in the spent fuel pool.

Accident conditions which could increase the k eff were evaluated including:

a. A new fuel assembly drop on the top of the racks;

b. A new fuel assembly misloaded between rack modules;

c. A new fuel assembly misloaded into an incorrect storage rack location;

d. Intramodule water gap reduction due to a seismic event; and

e. Spent fuel pool temperat ure greater than 150 ºF.

For an occurrence of these postula ted accident conditions, the double contingency principle of Reference 2 can be applied. This states that one is not required to assume two unlikely, independent, concurrent events to ensure protection agains t a criticality accident. Thus, for these postulated accident conditi ons, the presence of additional soluble boron in the spent fuel pool water (above the 464 ppm required to maintain keff less than 0.95 under normal conditions) can be assumed as a realistic initial condition since not assuming its presence would be a second unlikely event.

Calculations were performed (Ref. 4 and 5) to determine the amount of soluble boron required to offset the highest reactivity increase caused by these postula ted accidents and to maintain k eff less than or equal to 0.95. It was found that a spent fuel pool boron concentration of 730 ppm was ade quate to mitigate these postulated criticality related accidents and to maintain k eff less than or equal to 0.95. This specification ensures the spent fuel pool contains adequate dissolved boron to compen sate for the increased reactivity caused by these accidents. The 1800 ppm spent fuel pool boron concentration limit in this specification was chosen to be consistent with the boron concentration limit required for a spent fuel cask Fuel Storage Pool Boron Concentration B 3.7.16 Prairie Island Unit 1 - Revision 194 Units 1 and 2 B 3.7.16-3 Unit 2 - Revision 194 BASES APPLICABLE containing fuel. The 1800 ppm limit will ensure that k eff for the SAFETY spent fuel cask, including sta tistical probabilities, will be less than ANALYSES or equal to 0.95 for all postulated arrangements of fuel within the (continued) cask.

A spent fuel pool boron dilution analysis was performed which confirmed that sufficient time is av ailable to detect and mitigate a dilution of the spent fuel pool before the 0.95 k eff design basis is exceeded. The spent fuel pool boron dilution analysis concluded that an unplanned or inadverten t event which could result in the dilution of the spent fuel pool boron concentration from 1800 ppm to 750 ppm is not a credible event.

A spent fuel pool boron concentration of 750 ppm was required by the previous spent fuel rack criticality analysis to ensure that the spent fuel rack k eff would be less than or equal to 0.95 for the allowable storage configurations, ex cluding accidents. Therefore the spent fuel pool diluti on analysis utilized 750 ppm as the endpoint of the analysis to determine the d ilution time and volume of water required to dilute the spent fuel pool from the 1800 ppm Technical Specification limit. The current spen t fuel rack criticality analysis (Ref. 4 and 5) only requires a boron concentration of 464 ppm to ensure that the sp ent fuel rack k eff will be less than or equal to 0.95 for the allowable storage configur ation, excluding accidents. Therefore the spent fuel pool boron dilution analysis which assumes 750 ppm as the endpoint of the anal ysis is conservative with respect to the endpoint of 464 ppm since a larger volume of water would be required, which would take more time to dilute the spent fuel pool to 464 ppm. The concentration of dissolved boron in the fuel storage pool satisfies Criterion 2 of 10 CFR 50.36(c)(2)(ii).

Fuel Storage Pool Boron Concentration B 3.7.16 Prairie Island Unit 1 - Revision 194 Units 1 and 2 B 3.7.16-4 Unit 2 - Revision 194 BASES (continued)

LCO The fuel storage pool boron concentration is required to be 1800 ppm. The specified concentration of dissolved boron in the fuel storage pool preserves the assu mptions used in the analyses of the potential critical accident scenar ios as described in Reference 4 and 5. This concentration of dissolved boron is the minimum required concentration for fuel assembly storage, movement within the fuel storage pool, and for loading and unloading a spent fuel storage cask.

APPLICABILITY This LCO applies whenever fuel assemblies are st ored in the spent fuel storage pool.

ACTIONS A.1 and A.2 The Required Actions are modifi ed by a Note indicating that LCO 3.0.3 does not apply.

When the concentration of boron in the spent fuel storage pool is less than required, immediate action must be taken to preclude the occurrence of an accident or to mitigate the consequences of an accident in progress. This is most efficiently achieved by immediately suspending the movement of fuel assemblies. The concentration of boron is restored simultaneously with suspending movement of fuel assemblies. This does not preclude movement of a fuel assembly to a safe position.

If the LCO is not met while moving irradiated fuel assemblies in MODE 5 or 6, LCO 3.0.3 would not be applicable. If moving irradiated fuel assemblies while in MODE 1, 2, 3, or 4, the fuel movement is independent of reactor operation. Therefore, inability to suspend movement of fuel asse mblies is not sufficient reason to require a reactor shutdown.

Fuel Storage Pool Boron Concentration B 3.7.16 Prairie Island Unit 1 - Revision 194 Units 1 and 2 B 3.7.16-5 Unit 2 - Revision 194 BASES (continued)

SURVEILLANCE SR 3.7.16.1 REQUIREMENTS This SR verifies that the concentration of boron in the spent fuel storage pool is within the required lim it. As long as this SR is met, the analyzed accidents are fully addressed. The 7 day Frequency is appropriate because no major replenishment of pool water is expected to take place over such a short period of time.

REFERENCES 1. USAR, Section 10.2.

2. ANSI/ANS-8.1-1983.

3. Nuclear Regulatory Commission, Letter to All Power Reactor Licensees from B. k. Grimes, "OT Position for Review and Acceptance of Spent Fuel Storage and Handling Applications", April 14, 1978.

4. "Prairie Island Units 1 and 2 Spent Fuel Pool Criticality Analysis", WCAP-16517-NP, Revision 0, Westinghouse Electric Company, November 2005.

5. Addendum 1 to WCAP-16517-NP, Revision 0, "Prairie Island Units 1 and 2 Spent Fuel Pool Criticality Analysis", Bishop, T.C., February 2008

Spent Fuel Pool Storage B 3.7.17 Prairie Island Unit 1 - Revision 1 82 Units 1 and 2 B 3.7.17-2 Unit 2 - Revision 1 82 BASES BACKGROUND assembly in accordan ce with the accompanying LCO and (continued) maintaining boron concen tration in accordance with LCO 3.7.16.

APPLICABLE The hypothetical criticality accidents can only take place during SAFETY or as a result of the movement of an assembly (Ref. 4 and 5). For ANALYSES these accident occurrences, the presence of soluble boron in the spent fuel storage pool (controlled by LCO 3.7.16, "Fuel Storage Pool Boron Concentration") prevents criticality. By closely controlling the movement of each assembly and by verifying the appropriate checkerboarding after each fuel handling campaign, the time period for potential accidents may be limited to a small fraction of the total operating time. During the remaining time period with no potential for criticality accident s, the operation may be under the auspices of the accompanying LCO.

The spent fuel storage racks have be en analyzed in accordance with the methodology contained in Reference 4. That methodology ensures that the spent fuel rack multiplication factor, k eff , is less than 0.95 as recommended by ANSI 57.2-1983 (Ref. 6) and NRC guidance (Ref. 3). The codes, methods and techniques contained in the methodology are used to satisfy this criterion on k eff. The resulting Prairie Island spent fuel ra ck criticality analysis allows for the storage of fuel assemblies with enrichments up to a maximum of 5.0 (nominal 4.95% + 0.05%) weight percent U-235 while maintaining k eff 0.95 including uncertaintie s and credit for soluble boron. In addition, sub-criticality of the pool (k eff < 1.0) is assured on a 95/95 basis, without the presence of the soluble boron in the pool. Credit is taken for radioactive decay time of the spent fuel and for the presence of fuel rods containing gadolinium burnable poison.

The criticality analysis (Ref. 4 and 5) utilized the following storage configurations to ensure that the spent fuel p ool will remain subcritical during the storage of fu el assemblies w ith all possible combinations of burnup and initial enrichment:

Spent Fuel Pool Storage B 3.7.17 Prairie Island Unit 1 - Revision 182 Units 1 and 2 B 3.7.17-4 Unit 2 - Revision 182 BASES APPLICABLE b. The interface must be configured such that there is one row SAFETY carryover of the pattern of burned assemblies from the ANALYSES checkerboard region into the first row of the unrestricted region (continued) (Figure 4.3.1-2).

Specification 3.7.17 and Section 4.3 ensure that fu el is stored in the spent fuel racks in accordance with the storage configurations assumed in the spent fuel rack criticality analysis (Ref. 4 and 5). The spent fuel pool criticality anal ysis addresses all the fuel types currently stored in the sp ent fuel pool and in us e in the reactor. The fuel types considered in the an alysis include the Westinghouse Standard (STD), OFA, a nd Vantage Plus designs (both 0.400" and 0.422" O.D. designs), and the Exxon fuel asse mbly types in storage in the spent fuel pool.

Accident conditions which could increase the k eff were evaluated including:

a. A new fuel assembly drop on the top of the racks;

b. A new fuel assembly misloaded between rack modules;

c. A new fuel assembly misloaded into an incorrect storage rack location;

d. Intramodule water gap reduc tion due to a seismic event; and

e. Spent fuel pool temp erature greater than 150 F. For an occurrence of these postula ted accident conditions, the double contingency principle of Reference 2 can be applied. This states that one is not required to assume two unlikely, independent, concurrent events to ensure protection agains t a criticality accident. Thus, for these postulated accident conditi ons, the presence of additional soluble boron in the spent fuel pool water (above the 464 ppm required to maintain keff less than 0.95 under normal conditions) can Spent Fuel Pool Storage B 3.7.17 Prairie Island Unit 1 - Revision 18 6 Units 1 and 2 B 3.7.17-8 Unit 2 - Revision 18 6 BASES SURVEILLANCE SR 3.7.17.2 (continued) REQUIREMENTS The 7 day allowance for completion of this SR provides adequate time for completion of the spent fuel pool invent ory verification while minimizing the time a fuel assembly may be misloaded in the spent fuel pool. If a fuel assemb ly is misloaded during the fuel handling campaign, the minimum boron concentration required by LCO 3.7.16 will ensure that the spent fuel rack k eff remains within limits until the spent fuel invent ory verification is performed

REFERENCES 1. USAR, Section 10.2.

2. ANSI/ANS-8.1-1983.

3. Nuclear Regulatory Commission, Letter to All Power Reactor Licensees from B. K. Grimes, "OT Position for Review and Acceptance of Spent Fuel Storage and Handling Applications", April 14, 1978.

4. "Prairie Island Units 1 and 2 Spent Fuel Pool Criticality Analysis", WCAP-16517-NP, Revision 0, Westinghouse Electric Company, November 2005.

5. Not Used.

Addendum 1 to WCAP-16517-NP, Revision 0, "Prairie Island Units 1 and 2 Spent Fuel Pool Criticality Analysis", Bishop, T.C., February 2008

6. American Nuclear Society, "American National Standard Design Requirements for Light Water Reactor Fuel Storage Facilities at Nuclear Power Pl ants", ANSI/ANS-57.2-1983, October 7, 1983.

Attachment 3, Enclosure 1 Retyped Technical Specification Pages TS 2.0-1 TS 3.5.1-2 TS 4.0-2 through TS 4.0-4

5 Pages follow SLs 2.0 Prairie Island Unit 1 - Amendment No. 158 162 Units 1 and 2 2.0-1 Unit 2 - Amendment No. 149 153 2.0 SAFETY LIMITS (SLs) 2.1 SLs 2.1.1 Reactor Core SLs In MODES 1 and 2, the combination of THERMAL POWER, Reactor Coolant System (RCS) highest loop av erage temperature, and pressurizer pressure shall not exceed the limits sp ecified in the COLR and the following SLs shall not be exceeded:

2.1.1.1 The departure from nucleate boiling ratio (DNBR) shall be maintained >

1.17 for WRB-1 DNB correlation for OFA fuel.

2.1.1.2 The peak fuel centerline te mperature shall be maintained as follows: a) < 5080 ºF, decreasing by 58 ºF per 10,000 MWD/MTU burnup, for fuel containing UO2; b) < (5080 ºF minus 6.75 ºF per w/o Gd 2 O 3), decreasing by 58 ºF per 10,000 MWD/MTU burnup, fo r fuel containing gadolinia.

2.1.2 RCS Pressure SL In MODES 1, 2, 3, 4, and 5, the RCS pressure shall be maintained <

2735 psig. 2.2 SL Violations 2.2.1 If SL 2.1.1 is violated, restore comp liance and be in MODE 3 within 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br />.

2.2.2 If SL 2.1.2 is violated:

2.2.2.1 In MODE 1 or 2, restore compliance and be in MODE 3 within 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br />. 2.2.2.3 In MODE 3, 4, or 5, restor e compliance within 5 minutes.

Accumulators 3.5.1 Prairie Island Unit 1 - Amendment No. 158 Units 1 and 2 3.5.1-2 Unit 2 - Amendment No. 149 ACTIONS (continued) CONDITION REQUIRED ACTION COMPLETION TIME D. Two accumulators inoperable.

D.1 Enter LCO 3.0.3.

Immediately

SURVEILLANCE REQUIREMENTS SURVEILLANCE FREQUENCY SR 3.5.1.1 Verify each accumul ator isolation valve is fully open.

12 hours1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br /> SR 3.5.1.2 Verify borated wate r volume in each accumulator is > 1250 cubic feet (25%) and <

1290 cubic feet (91%).

12 hours1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br /> SR 3.5.1.3 Verify nitroge n cover pressure in each accumulator is

> 710 psig and <

770 psig.

12 hours1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br /> SR 3.5.1.4 Verify boron concen tration in each accumulator is

> 2300 ppm.

31 days

SR 3.5.1.5 Verify power is removed from each accumulator isolation valve operator when RCS pressure is

> 2000 psig.

31 days Design Features

4.0 Prairie

Island Unit 1 - Amendment No. 158 172 Units 1 and 2 4.0-2 Unit 2 - Amendment No. 149

.162 4.0 DESIGN FEATURES (continued) 4.3 Fuel Storage

4.3.1 Criticality

4.3.1.1 The spent fuel storage racks are designed and shall be maintained with: a. Fuel assemblies having a maximum U-235 enrichment of

5.0 weight

percent;

b. k eff < 1.0 if fully flooded w ith unborated water, which includes an allowance for unce rtainties as described in USAR Section 10.2;

c. k eff < 0.95 if fully flooded with water borated to 730 ppm, which includes an allowance fo r uncertainties as described in USAR Section 10.2;

d. A nominal 9.5 inch center to center distance between fuel assemblies placed in the fuel storage racks;

e. New or spent fuel assemb lies with a combination of discharge burnup, initial enrich ment and decay time in the "unrestricted range" of Figure 3.7.17-1 may be allowed unrestricted storage in the fuel storage racks; and

f. New or spent fuel assemb lies with a combination of discharge burnup, initial enrich ment and decay time in the "restricted range" of Figure 3.

7.17-1 will be stored in compliance with Figures 4.3.1-1 through 4.3.1-4.

Design Features

4.0 Prairie

Island Unit 1 - Amendment No. 158 Units 1 and 2 4.0-3 Unit 2 - Amendment No. 149

4.0 DESIGN

FEATURES 4.3 Fuel Storage (continued) 4.3.1.2 The new fuel storage racks are designed and shall be maintained with: a. Fuel assemblies having a maximum U-235 enrichment of

5.0 weight

percent;

b. k eff < 0.95 if fully flooded w ith unborated water, which includes an allowance for unce rtainties as described in USAR Section 10.2;

c. k eff < 0.98 if accidentally filled with a low density moderator which resulted in optimum low density moderation conditions; and

d. A nominal 21 inch center to center distance between fuel assemblies placed in the storage racks.

4.3.1.3 Fuel will not be inserted into a TN-40 spent fuel cask in the pool unless a minimum boron concentration of 1800 ppm is present.

The 1800 ppm will ensure that k eff for the spent fuel cask, including statistical un certainties, will be < 0.95 for all postulated arrangements of fuel within the cask. The criticality analyses for the TN-40 spent fuel storage cask were based on fresh fuel enriched to 3.85 weight percent U-235.

4.3.2 Drainage

The spent fuel storage pool is designe d and shall be maintained to prevent inadvertent draining of the pool below elevation 727 4 (Mean Sea Level).

Accumulators 3.5.1 Prairie Island Unit 1 - Amendment No. 172 180 Units 1 and 2 4.0-4 Unit 2 - Amendment No. 162 1 70 4.0 DESIGN FEATURES 4.3 Fuel Storage (continued)

4.3.3 Capacity

The spent fuel storage pool is desi gned and shall be maintained with a storage capacity limited to no more than 1386 fuel assemblies not including those assemblies which can be returned to the reactor. The southeast corner of the small pool serves as the spent fuel cask lay down area. To facilitate plant evolutions, four additional storage racks, with a combined capacity of 196, may be temporarily installed in the cask lay down area to provide a total of 1582 storage locations. (USAR Section 10.2)

Enclosure 1, Attachment 4 Prairie Island Units 1 and 2 422V+ Reload Transition Licensing Report

262 Pages follow

Prairie Island Units 1&2 422V+ Reload Transition Licensing Report

May 2008

i TABLE OF CONTENTS LIST OF TABLES.......................................................................................................................................iv LIST OF FIGURES.....................................................................................................................................vi LIST OF ACRONYMS................................................................................................................................xi 1 INTRODUCTION AND

SUMMARY

..........................................................................................1-1

1.1 INTRODUCTION

...........................................................................................................1-1 1.2 FUEL FEATURES (422V+)............................................................................................1-2

1.3 PEAKING

FACTORS.....................................................................................................1-3 1.4 RTDP UNCERTAINTIES................................................................................................1-3

1.5 PERFORMANCE

CAPABILITIES WORKING GROUP PARAMETERS...................1-4

1.6 GENERAL

ANALYSIS ASSUMPTIONS......................................................................1-4

1.7 CONCLUSION

S.............................................................................................................1-5

1.8 REFERENCES

................................................................................................................1-5 2 MECHANICAL DESIGN FEATURES........................................................................................2-1

2.1 INTRODUCTION

AND

SUMMARY

............................................................................2-1

2.2 COMPATIBILITY

OF FUEL ASSEMBLIES.................................................................2-1 2.2.1 Fuel Rods.........................................................................................................2-2 2.2.2 Grid Assemblies...............................................................................................2-3

2.2.3 Guide

Thimble and Instrumentation Tubes.....................................................2-4

2.2.4 Reconstitutable

Top Nozzle.............................................................................2-4

2.2.5 Debris

Filter Bottom Nozzle............................................................................2-4

2.3 MECHANICAL

PERFORMANCE................................................................................2-4 2.4 FUEL ROD PERFORMANCE........................................................................................2-4 2.5 SEISMIC/LOCA IMPACT ON FUEL ASSEMBLIES....................................................2-9 2.5.1 Fuel Assembly and Reactor Core Models.......................................................2-9 2.5.2 Grid Load Analysis........................................................................................2-10 2.5.3 Conclusions...................................................................................................2-10

2.6 REFERENCES

..............................................................................................................2-11 3 NUCLEAR DESIGN....................................................................................................................3-1

3.1 INTRODUCTION

AND

SUMMARY

............................................................................3-1

3.2 DESIGN

BASIS..............................................................................................................3-1 3.3 METHODOLOGY..........................................................................................................3-2

3.4 DESIGN

EVALUATION - PHYSICS CHARACTERISTICS AND KEY SAFETY PARAMETERS...............................................................................................................3-2

3.5 DESIGN

EVALUATION - POWER DISTRIBUTIONS AND PEAKING FACTORS........................................................................................................................

3-3 3.6 NUCLEAR DESIGN EVALUATION CONCLUSIONS................................................3-3

3.7 REFERENCES

................................................................................................................3-4

ii TABLE OF CONTENTS (cont.)

4 THERMAL AND HYDRAULIC DESIGN..................................................................................4-1

4.1 INTRODUCTION

AND

SUMMARY

............................................................................4-1 4.2 METHODOLOGY..........................................................................................................4-1

4.3 HYDRAULIC

COMPATIBILITY...................................................................................4-2

4.4 EFFECTS

OF FUEL ROD BOW ON DNBR.................................................................4-3 4.5 FUEL TEMPERATURE/PRESSURE ANALYSIS.........................................................4-3

4.6 TRANSITION

CORE EFFECT......................................................................................4-4

4.7 BYPASS

FLOW..............................................................................................................4-5 4.8 THERMAL-HYDRAULIC DESIGN PARAMETERS...................................................4-5

4.9 CONCLUSION

................................................................................................................4-5 4.10 REFERENCES................................................................................................................4-5 5 ACCIDENT ANALYSIS..............................................................................................................

5-1 5.1 NON-LOCA TRANSIENTS...........................................................................................5-1

5.1.1 Uncontrolled

RCCA Withdrawal from a Subcritical Condition (USAR Section 14.4.1)....................................................................................5-3

5.1.2 Uncontrolled

RCCA Withdrawal at Power (USAR Section 14.4.2)................5-6 5.1.3 RCCA Misalignment (USAR Section 14.4.3).................................................5-7

5.1.4 Chemical

and Volume Control System Malfunction (USAR Section 14.4.4)....................................................................................5-7

5.1.5 Startup

of an Inactive Reactor Coolant Loop (USAR Section 14.4.5)............5-8

5.1.6 Feedwater

Malfunction (USAR Section 14.4.6)..............................................5-8

5.1.7 Excessive

Load Increase Incident (USAR Section 14.4.7).............................5-9 5.1.8 Loss of Reactor Coolant Flow - Flow Coastdown (USAR Section 14.4.8)....................................................................................5-9 5.1.9 Loss of Reactor Coolant Flow - Locked Rotor (USAR Section 14.4.8)..................................................................................5-14 5.1.10 Loss of External Electrical Load (USAR Section 14.4.9).............................5-17 5.1.11 Loss of Normal Feedwater (USAR Section 14.4.10)....................................5-17 5.1.12 Loss of All AC Power to the Station Auxiliaries (USAR Section 14.4.11)................................................................................5-18 5.1.13 Rupture of a Steam Pipe (USAR Section 14.5.5)..........................................5-18 5.1.14 RCCA Ejection (USAR Section 14.5.6)........................................................5-19 5.1.15 ATWS (USAR Section 14.8).........................................................................5-23 5.2 LOSS-OF-COOLANT ACCIDENTS............................................................................5-70 5.2.1 Large-Break Best-Estimate LOCA................................................................5-70 5.2.2 Small-Break LOCA.......................................................................................5-70 5.2.3 Post-LOCA Long-Term Subcriticality Cooling Evaluation...........................5-74 5.2.4 Post-LOCA Boron Buildup Analysis and Long-Term Post-LOCA Cooling..........................................................................................................5-75 5.2.5 LOCA Hydraulic Forces................................................................................5-76

iii TABLE OF CONTENTS (cont.) 5.3 RUPTURES IN PRIMARY SIDE PIPING.................................................................5-126 5.3.1 Long-Term LOCA Mass and Energy Releases Inside Containment...........5-126 5.3.2 LOCA Containment Peak Pressure Response Analysis...............................5-133 5.3.3 Short-Term Mass and Energy Release and Containment Subcompartment Pressurization Evaluation................................................5-136 5.4 RUPTURES IN SECONDARY SIDE PIPING...........................................................5-148 5.4.1 Steam Line Break Inside Containment........................................................5-148 5.4.2 Mass and Energy Releases Outside Containment........................................5-154

5.5 REFERENCES

............................................................................................................5-160 6 MECHANICAL ANALYSIS.......................................................................................................6-1 6.1 Reactor Internals Structural Analysis..............................................................6-1 6.2 References.......................................................................................................6-6 7 NRC CONDITIONAL REQUIREMENTS FOR THE USE OF 422V+ FUEL, AND COMPUTER CODES UTILIZED...............................................................................................7-1 7.1 VANTAGE+.....................................................................................................................7-1 7.2 LOCA/NOTRUMP-EM: ZIRLO MODIFICATIONS.....................................................7-2 7.3 VIPRE..............................................................................................................................7-4 7.4 FACTRAN.......................................................................................................................7-7 7.5 RETRAN.......................................................................................................................7-11 7.6 LOFTRAN.....................................................................................................................7-14 7.7 TWINKLE.....................................................................................................................7-16 7.8 ADVANCED NODAL CODE (ANC)...........................................................................7-17 7.9 GOTHIC........................................................................................................................7-18 iv LIST OF TABLES Table 1-1 RTDP Uncertainties.........................................................................................................1-7 Table 1-2 422V+ Fuel Upgrade NSSS Design Parameters - Prairie Island Units 1&2...................1-8 Table 1-3 422V+ Fuel Upgrade NSSS Design Parameters - Prairie Island Unit 2..........................1-9 Table 2-1 Comparison of 14x14 STD, 400V+, 422V+ Fuel Assembly Mechanical Design Parameters......................................................................................................................2-13 Table 3-1 Key Safety Parameters.....................................................................................................3-5 Table 4-1 Prairie Island Thermal-Hydraulic Design Parameters Comparison.................................4-7 Table 4-2 Hot Channel Factors (1).....................................................................................................4-9 Table 4-3 RTDP Uncertainties.......................................................................................................4-10 Table 4-4 RTDP DNBR Margin Summary (1).................................................................................4-11 Table 4-5 Limiting Parameter Direction........................................................................................4-12 Table 5.1-1 Non-LOCA Transients Evaluated or Analyzed...............................................................5-2 Table 5.1.1-1 Assumptions and Results - Uncontrolled RCCA Withdrawal from a Subcritical Condition.......................................................................................................................5-24 Table 5.1.1-2 Sequence of Events - Uncontrolled RCCA Withdrawal from a Subcritical Condition.......................................................................................................................5-25 Table 5.1.8-1 Sequence of Events - Partial Loss of Reactor Coolant Flow........................................5-25 Table 5.1.8-2 Sequence of Events - Complete Loss of Reactor Coolant Flow...................................5-25 Table 5.1.9-1 Sequence of Events - Reactor Coolant Pump Locked Rotor.........................................5-26 Table 5.1.14-1 Assumptions and Results-RCCA Ejection.....................................................................5-26 Table 5.1.14-2 Prairie Island Rod Ejection Results with and Without Gadolinium...............................5-27 Table 5.1.14-3 Sequence of Events - RCCA Ejection...........................................................................5-28 Table 5.2.1-1 Prairie Island Unit 1 Best-Estimate UPI Large-Break LOCA Results...........................5-77 Table 5.2.1-2 Prairie Island Unit 1 Conditions Analyzed with WCOBRA/TRAC Compared to Best-Estimate UPI Test Conditions................................................................................5-77 Table 5.2.1-3 Prairie Island Unit 2 Best-Estimate UPI Large-Break LOCA Results...........................5-77 Table 5.2.1-4 Prairie Island Unit 2 Conditions Analyzed with WCOBRA/TRAC Compared to Best-Estimate UPI Test Conditions...............................................................................5-77 Table 5.2.2-1 Input Assumptions and Initial Conditions for Units 1&2..............................................5-78 Table 5.2.2-2 Steam Generator Safety Valve Flows per Steam Generator...........................................5-80

v LIST OF TABLES (cont.) Table 5.2.2-3 HHSI Flow for One HHSI Pump with Faulted Loop Spilling to RCS Pressure (Break sizes less than 5.187 inches)..............................................................................5-80 Table 5.2.2-4 HHSI Flows for One HHSI Pump with Faulted Loop Spilling to Containment Pressure (0 psig) (Break sizes greater than 5.187 inches)............................................................5-81 Table 5.2.2-5 RHR Flows for One RHR Pump with No Spilling During Injection from RWST........5-81 Table 5.2.2-6 Time Sequence of Events for Unit 1..............................................................................5-82 Table 5.2.2-7 Beginning-of-Life Fuel Rod Heatup Results for Unit 1................................................5-82 Table 5.2.2-8 Time Sequence of Events for Unit 2..............................................................................5-83 Table 5.2.2-9 Beginning-of-Life Fuel Rod Heatup Results for Unit 2................................................5-83 Table 5.3-1 System Parameters Initial Conditions..........................................................................5-137 Table 5.3-2 SI Flow Minimum Safeguards.....................................................................................5-138 Table 5.3-3 SI Flow Maximum Safeguards....................................................................................5-139 Table 5.3-4 LOCA Containment Response Analysis Parameters...................................................5-140 Table 5.3-5 DEHL Break Sequence of Events................................................................................5-142 Table 5.3-6 DEPS Break Sequence of Events (Minimum Safeguards)..........................................5-143 Table 5.3-7 DEPS Break Sequence of Events (Maximum Safeguards)..........................................5-144 Table 5.3-8 LOCA Containment Integrity Response Results (Loss-of-Offsite Power Assumed).....................................................................................................................5-145 Table 5.4-1 Initial Containment Conditions....................................................................................5-155 Table 5.4-2 Sequence of Events for Steam Line Break 30% Initial Power, Containment Safeguards Failure (Peak Containment Pressure Cases)..............................................5-155 Table 5.4-3 Sequence of Events for Steam Line Break Full Initial Power, Containment Safeguards Failure (Peak Containment Vapor Temperature Cases)............................5-156

vi LIST OF FIGURES Figure 2-1 14x14 400V+ and 422V+ Fuel Assembly Designs

• ...................................................2-15 Figure 2-2 14x14 400V+ and 422V+ Fuel Rod Comparison.......................................................2-16 Figure 3-1 First Transition Cycle Loading Pattern with BOC and EOC Assembly Burnups.........3-6 Figure 3-2 First Transition Cycle BOC, MOC, and EOC Assembly Power Distributions.............3-7 Figure 3-3 Second Transition Cycle Loading Pattern with BOC and EOC Assembly Burnups....3-8 Figure 3-4 Second Transition Cycle BOC, MOC, and EOC Assembly Power Distributions........3-9 Figure 3-5 Third Transition Cycle, Pattern A, Loading Pattern with BOC and EOC Assembly Burnups......................................................................................................3-10 Figure 3-6 Third Transition Cycle, Pattern A, BOC, MOC, and EOC Assembly Power Distributions...............................................................................................................3-1 1 Figure 3-7 Third Transition Cycle, Pattern B, Loading Pattern with BOC And EOC Assembly Burnups......................................................................................................3-12 Figure 3-8 Third Transition Cycle, Pattern B, BOC, MOC, and EOC Assembly Power Distributions...............................................................................................................3-1 3 Figure 3-9 Critical Boron Concentration Comparison Versus Exposure......................................3-14 Figure 3-10 Axial Offset Comparison Versus Exposure................................................................3-15 Figure 3-11 Radial Peaking Factor (F NH) Comparison Versus Exposure.....................................3-16 Figure 3-12 Total Peaking Factor ()Z (F N Q) Comparison Versus Exposure...................................3-17 Figure 5.1.1-1 Uncontrolled RCCA Withdrawal from a Subcritical Condition - Reactor Power Versus Time................................................................................................................5-29 Figure 5.1.1-2 Uncontrolled RCCA Withdrawal from a Subcritical Condition - Heat Flux Versus Time................................................................................................................5-30 Figure 5.1.1-3 Uncontrolled RCCA Withdrawal from a Subcritical Condition - Hot Spot Fuel Centerline Temperature Versus Time..........................................................................5-31 Figure 5.1.1-4 Uncontrolled RCCA Withdrawal from a Subcritical Condition - Hot Spot Fuel Average Temperature Versus Time.............................................................................5-32 Figure 5.1.1-5 Uncontrolled RCCA Withdrawal from a Subcritical Condition - Hot Spot Cladding Temperature Versus Time............................................................................5-33 Figure 5.1.8-1 Total Core Inlet Flow Versus Time - Partial Loss of Flow (PLOF), One Pump Coasting Down...........................................................................................................5-34 Figure 5.1.8-2 RCS Faulted Loop Flow Versus Time - PLOF, One Pump Coasting Down..............5-35 Figure 5.1.8-3 Nuclear Power Versus Time - PLOF, One Pump Coasting Down.............................5-36 vii LIST OF FIGURES (cont.) Figure 5.1.8-4 Core Average Heat Flux Versus Time - PLOF, One Pump Coasting Down...............5-37 Figure 5.1.8-5 Pressurizer Pressure Versus Time - PLOF, One Pump Coasting Down.....................5-38 Figure 5.1.8-6 RCS Faulted Loop Temperature Versus Time - PLOF, One Pump Coasting Down..5-39 Figure 5.1.8-7 Hot Channel Heat Flux Versus Time - PLOF, One Pump Coasting Down................5-40 Figure 5.1.8-8 DNBR Versus Time - PLOF, One Pump Coasting Down..........................................5-41 Figure 5.1.8-9 Total Core Inlet Flow Versus Time - Complete Loss of Flow (CLOF) - Two Pumps Coasting Down...............................................................................................5-42 Figure 5.1.8-10 RCS Loop Flow Versus Time - CLOF - Two Pumps Coasting Down......................5-43 Figure 5.1.8-11 Nuclear Power Versus Time - CLOF - Two Pumps Coasting Down.........................5-44 Figure 5.1.8-12 Core Average Heat Flux Versus Time - CLOF - Two Pumps Coasting Down..........5-45 Figure 5.1.8-13 Pressurizer Pressure Versus Time - CLOF - Two Pumps Coasting Down................5-46 Figure 5.1.8-14 RCS Faulted Loop Temperature Versus Time - CLOF - Two Pumps Coasting Down..........................................................................................................................5

-47 Figure 5.1.8-15 Hot Channel Heat Flux Versus Time - CLOF - Two Pumps Coasting Down...........5-48 Figure 5.1.8-16 DNBR Versus Time - CLOF - Two Pumps Coasting Down.....................................5-49 Figure 5.1.9-1 Total Core Inlet Flow Versus Time - Locked Rotor/Shaft Break - RCS Pressure/Peak Cladding Temperature Case................................................................5-50 Figure 5.1.9-2 RCS Loop Flow Versus Time - Locked Rotor/Shaft Break - RCS Pressure/Peak Cladding Temperature Case........................................................................................5-51 Figure 5.1.9-3 Nuclear Power Versus Time - Locked Rotor/Shaft Break - RCS Pressure/Peak Cladding Temperature Case........................................................................................5-52 Figure 5.1.9-4 Core Average Heat Flux Versus Time - Locked Rotor/Shaft Break - RCS Pressure/Peak Cladding Temperature Case................................................................5-53 Figure 5.1.9-5 Pressurizer Pressure Versus Time - Locked Rotor/Shaft Break - RCS Pressure/Peak Cladding Temperature Case................................................................5-54 Figure 5.1.9-6 Vessel Lower Plenum Pressure Versus Time - Locked Rotor/Shaft Break - RCS Pressure/Peak Cladding Temperature Case................................................................5-55 Figure 5.1.9-7 RCS Loop Temperature Versus Time - Locked Rotor/Shaft Break - RCS Pressure/Peak Cladding Temperature Case................................................................5-56 Figure 5.1.9-8 Hot Spot Cladding Inner Temperature Versus Time - Locked Rotor/Shaft Break - RCS Pressure/Peak Cladding Temperature Case........................................................5-57 Figure 5.1.14-1 RCCA Ejection - BOC Full-Power Reactor Power Versus Time...............................5-58 viii LIST OF FIGURES (cont.) Figure 5.1.14-2 RCCA Ejection - BOC Full-Power Fuel and Cladding Temperatures Versus Time for U0 2 Fuel Case..............................................................................................5-59 Figure 5.1.14-3 RCCA Ejection - BOC Full-Power Fuel and Cladding Temperatures Versus Time for Gadolinium Fuel Case...........................................................................................5-60 Figure 5.1.14-4 RCCA Ejection - BOC Zero-Power Reactor Power Versus Time..............................5-61 Figure 5.1.14-5 RCCA Ejection - BOC Zero-Power Fuel and Cladding Temperatures Versus Time for U0 2 Fuel Case..............................................................................................5-62 Figure 5.1.14-6 RCCA Ejection - BOC Zero-Power Fuel and Cladding Temperatures Versus Time for Gadolinium Fuel Case.................................................................................5-63 Figure 5.1.14-7 RCCA Ejection - EOC Full-Power Reactor Power Versus Time...............................5-64 Figure 5.1.14-8 RCCA Ejection - EOC Full-Power Fuel and Cladding Temperatures Versus Time for U0 2 Fuel Case.......................................................................................................5-65 Figure 5.1.14-9 RCCA Ejection - EOC Full-Power Fuel and Cladding Temperatures Versus Time for Gadolinium Fuel Case...........................................................................................5-66 Figure 5.1.14-10 RCCA Ejection - EOC Zero-Power Reactor Power Versus Time..............................5-67 Figure 5.1.14-11 RCCA Ejection - EOC Zero-Power Fuel and Cladding Temperatures Versus Time............................................................................................................................5-68 Figure 5.1.14-12 RCCA Ejection - EOC Zero-Power Fuel and Cladding Temperatures Versus Time for Gadolinium Fuel Case.................................................................................5-69 Figure 5.2.2-1 Code Interface Description for Small Break Model...................................................5-84 Figure 5.2.2-2 Hot Rod Axial Power Shape.......................................................................................

5-85 Figure 5.2.2-3 Reactor Coolant System Pressure Inch Break (Unit 1)........................................5-86 Figure 5.2.2-4 Core Mixture Level Inch Break (Unit 1)..............................................................5-87 Figure 5.2.2-5 Core Exit Vapor Temperature Inch Break (Unit 1)...............................................5-88 Figure 5.2.2-6 Broken Loop and Intact Loop Pumped SI Flow Rates Inch Break (Unit 1)........5-89 Figure 5.2.2-7 Total Break Flow and Total Pumped SI Flow Rates Inch Break (Unit 1)............5-90 Figure 5.2.2-8 Cladding Temperature at PCT Elevation and Local Oxidation at MLO Elevation Inch Break (Unit 1)..............................................................................5-91 Figure 5.2.2-9 Reactor Coolant System Pressure Inch Break (Unit 2)........................................5-92 Figure 5.2.2-10 Core Mixture Level Inch Break (Unit 2)..............................................................5-93 Figure 5.2.2-11 Core Exit Vapor Temperature Inch Break (Unit 2)...............................................5-94 Figure 5.2.2-12 Broken Loop and Intact Loop Pumped SI Flow Rates Inch Break (Unit 2)........5-95 ix LIST OF FIGURES (cont.) Figure 5.2.2-13 Total Break Flow and Total Pumped SI Flow Rates Inch Break (Unit 2)............5-96 Figure 5.2.2-14 Cladding Temperature at PCT Elevation and Local Oxidation at MLO Elevation Inch Break (Unit 2)..............................................................................5-97 Figure 5.2.2-15 Reactor Coolant System Pressure - 1.5-Inch Break (Unit 1).....................................5-98 Figure 5.2.2-16 Core Mixture Level - 1.5-Inch Break (Unit 1)...........................................................5-99 Figure 5.2.2-17 Reactor Coolant System Pressure Inch Break (Unit 1)......................................5-100 Figure 5.2.2-18 Core Mixture Level Inch Break (Unit 1)............................................................5-101 Figure 5.2.2-19 Cladding Temperature at PCT Elevation and Local Oxidation at MLO Elevation Inch Break (Unit 1)............................................................................................5-102 Figure 5.2.2-20 Reactor Coolant System Pressure Inch Break (Unit 1)......................................5-103 Figure 5.2.2-21 Core Mixture Level Inch Break (Unit 1)...........................................................5-104 Figure 5.2.2-22 Cladding Temperature at PCT Elevation and Local Oxidation at MLO Elevation Inch Break (Unit 1).............................................................................................5-105 Figure 5.2.2-23 Reactor Coolant System Pressure 6-Inch Break (Unit 1).........................................5-106 Figure 5.2.2-24 Core Mixture Level Inch Break (Unit 1)............................................................5-107 Figure 5.2.2-25 Reactor Coolant System Pressure Inch Break (Unit 1)......................................5-108 Figure 5.2.2-26 Core Mixture Level Inch Break (Unit 1)............................................................5-109 Figure 5.2.2-27 Reactor Coolant System Pressure - 10.126-Inch Break (Unit 1).............................5-110 Figure 5.2.2-28 Core Mixture Level - 10.126-Inch Break (Unit 1)...................................................5-111 Figure 5.2.2-29 Reactor Coolant System Pressure - 1.5-Inch Break (Unit 2)...................................5-112 Figure 5.2.2-30 Core Mixture Level - 1.5-Inch Break (Unit 2).........................................................5-113 Figure 5.2.2-31 Reactor Coolant System Pressure Inch Break (Unit 2)......................................5-114 Figure 5.2.2-32 Core Mixture Level Inch Break (Unit 2)............................................................5-115 Figure 5.2.2-33 Cladding Temperature at PCT Elevation and Local Oxidation at MLO Elevation Inch Break (Unit 2)............................................................................................5-116 Figure 5.2.2-34 Reactor Coolant System Pressure Inch Break (Unit 2)......................................5-117 Figure 5.2.2-35 Core Mixture Level Inch Break (Unit 2)............................................................5-118 Figure 5.2.2-36 Cladding Temperature at PCT Elevation and Local Oxidation at MLO Elevation Inch Break (Unit 2)............................................................................................5-119 Figure 5.2.2-37 Reactor Coolant System Pressure Inch Break (Unit 2)......................................5-120 Figure 5.2.2-38 Core Mixture Level Inch Break (Unit 2)............................................................5-121 x LIST OF FIGURES (cont.) Figure 5.2.2-39 Reactor Coolant System Pressure Inch Break (Unit 2)......................................5-122 Figure 5.2.2-40 Core Mixture Level Inch Break (Unit 2)............................................................5-123 Figure 5.2.2-41 Reactor Coolant System Pressure - 10.126-Inch Break (Unit 2).............................5-124 Figure 5.2.2-42 Core Mixture Level - 10.126-Inch Break (Unit 2)...................................................5-125 Figure 5.3-1 LOCA Containment Integrity Analysis - Containment Pressure Response.............5-146 Figure 5.3-2 LOCA Containment Integrity Analysis - Containment Gas Temperature Response...................................................................................................................5-14 6 Figure 5.3-3 LOCA Containment Integrity Analysis - Containment Sump Temperature.............5-147 Figure 5.3-4 LOCA Containment Integrity Analysis: - Containment Liner Temperature............5-147 Figure 5.4-1 Steam Line Break Flow Rate, 30% Power, Containment Safeguards Failure (Peak Containment Pressure Cases).........................................................................5-157 Figure 5.4-2 Steam Line Break Enthalpy, 30% Power, Containment Safeguards Failure (Peak Containment Pressure Cases).........................................................................5-157 Figure 5.4-3 Steam Line Break Flow Rate, Full-Power, Containment Safeguards Failure (Peak Containment Temperature Cases)...................................................................5-158 Figure 5.4-4 Steam Line Break Enthalpy, Full-Power, Containment Safeguards Failure (Peak Containment Temperature Cases)...................................................................5-158 Figure 5.4-5 Containment Pressure, 30% Power, Containment Safeguards Failure (Peak Containment Pressure Cases).........................................................................5-159 Figure 5.4-6 Containment Temperature, Full-Power, Containment Safeguards Failure (Peak Containment Temperature Cases)...................................................................5-159

xi LIST OF ACRONYMS AEC Atomic Energy Commission AFW auxiliary feedwater ALARA as low as reasonably achievable AMSAC anticipated transient without scram mitigation system actuation circuitry ANS American Nuclear Society ASME American Society of Mechanical Engineers ASTRUM automated statistical treatment of uncertainty method ATWS anticipated transient without scram B&PV Boiler and Pressure Vessel Code BAST boric acid storage tank BEBF best-estimate bypass flow BL broken loop BOC beginning of cycle BOL beginning of life CCW component cooling water CFCU containment fan coil units CFR Code of Federal Regulations CHF critical heat flux CLOF complete loss of flow COLR Core Operating Limits Report CRDM control rod drive mechanism CVCS chemical and volume control system DEHL double-ended hot let DEPS double-ended pump suction DEPS double-ended pump suction DER double-ended rupture DFBN debris filter bottom nozzle DLM diffusion layer model DNB departure from nucleate boiling DNBR departure from nucleate boiling ratio DSS diverse scram system ECCS emergency core cooling system EDG emergency diesel generator ELI excessive load increase EM evaluation model EOC end of cycle EOL end of life EQ equipment qualification ESF engineered safety feature ESFAS engineered safety feature actuation system FCEP Fuel Criteria Evaluation Process FIV flow-induced vibration FLI feedwater line isolation FPS full-power second xii FRV feedwater regulating valve FU fuel upgrade FW flow-induced vibration FWI feedwater isolation GDC General Design Criteria HFP hot full power HHSI high-head safety injection HI heat input HPO hot pump overspeed HZP hot zero power ID inner diameter IFM intermediate flow mixing ITC isothermal temperature coefficient KNPP Kewaunee Nuclear Power Plant LBB leak before break LBLOCA large-break loss-of-coolant accident LOCA loss-of-coolant accident LOOP loss of offsite power LPD low-pressure drop MCO moisture carryover MDC moderator density coefficient MI mass input MMF minimum measured flow MOC middle of cycle MSS main steam system MTC moderator temperature coefficient MUR-PU measurement uncertainty recapture power uprate NMC Nuclear Management Company NMV non-mixing vane NRC Nuclear Regulatory Commission NRS narrow-range span NSSS nuclear steam supply system OBE operating basis earthquake OD outer diameter OFA optimized fuel assembly OPPS over-pressure protection system OPT overpower T OSG Original Steam Generator OTT overtemperature T PCT peak cladding temperature PCWG Performance Capabilities Working Group PLOF partial loss of flow PORV power-operated relief valve PTLR Pressure-Temperature Limit Report PWR pressurized water reactor RCCA rod cluster control assembly xiii RCL reactor coolant loop RCP reactor coolant pump RCS reactor coolant system RHR residual heat removal RPS reactor protection system RPV reactor pressure vessel RSAC Reload Safety Analysis Checklist RSE Reload Safety Evaluation RSG replacement steam generator RTDP Revised Thermal Design Procedure RTN removable top nozzle RTSR Reload Transition Safety Report RWST refueling water storage tank SAC Safety Analysis Checklist SBLOCA small-break loss-of-coolant accident SER Safety Evaluation Report SGTP steam generator tube plugging SI safety injection SIS safety injection system SLI steam line isolation SRP Standard Review Plan SRSS square root of sum of square SSE safety shutdown earthquake STD STANDARD STDP Standard Thermal Design Procedure TDBF thermal design bypass flow TDF thermal design flow UPI upper plenum injection USAR Updated Safety Analysis Report V+ Vantage+

1-1 1 INTRODUCTION AND

1.1 INTRODUCTION

The Prairie Island Nuclear Generating Plant plans to refuel and operate with upgraded Westinghouse fuel, commencing with Cycle 26 for both Units 1&2. The upgraded fuel is 0.422-inch outer diameter (OD), 14x14 VANTAGE+ fuel with PERFORMANCE+ features - hereafter referred to as 422V+. This fuel is similar to the 422V+ fuel assemblies in operation in Point Beach Units 1&2 and Kewaunee Nuclear Power Plant (KNPP). The current fuel in operation at Prairie Island is 0.400-inch OD, 14x14 VANTAGE+ fuel with PERFORMANCE+ features hereafter referred to as 400V+. PERFORMANCE+ features can include pre-oxidized cladding, a protective bottom grid, ZIRLOTM(1) mid/intermediate flow mixing (IFM) grids, low cobalt top/bottom nozzle, and the debris mitigating bottom end plug. Specific PERFORMANCE+ features that are in use in Prairie Island are described in Section 2. VANTAGE+ fuel is defined by the use of ZIRLO TM cladding, which was approved in WCAP-12610-P-A, (Reference 1). Fuel with a 0.400-inch OD was first introduced with the optimized fuel assembly (OFA) in WCAP-9500-A, (Reference 2). This fuel diameter was utilized in all initial uses of VANTAGE+ fuel, including the introduction of VANTAGE+ in Prairie Island, prior to the introduction of the 0.422-inch OD VANTAGE+ in Point Beach. Therefore, it is significant to note that the fuel herein referred to as 400V+ is the same as that which is referred to as VANTAGE+ in the Prairie Island Updated Safety Analysis Report (USAR). The term 400V+ is being used in this report to further emphasize that the primary function of the fuel change is a change in diameter from 0.400 to 0.422 inches.The significant new mechanical features of the 422V+ design that are a new product feature for Prairie Island include the following: 0.422-inch OD fuel rod 0.422-inch OD instrumentation tube New 400V+ style (2.25-inch tall, vertical springs, horizontal dimples) mid-grid, designed to be compatible with the 0.422-inch OD fuel rods and 14x14 400V+ mixed fuel cores This report summarizes the evaluations and analyses that were performed to confirm the acceptable use of these features for Prairie Island Units 1&2 operations. In general, an evaluation was performed for events that the transient conditions of the RCS and main steam system (MSS) were determined to be insignificantly impacted by the change in fuel, whereas analysis was performed for those events in which the change in fule has a significant impact on the transient conditions of the RCS and MSS.. Sections 2.0 through 6.0 of this report provide the results of the mechanical (fuel), nuclear, thermal-hydraulic, accident, and reactor internals mechanical evaluations and analyses, respectively. The analyses and evaluations included in this report support the necessary Technical Specification changes. Nuclear

1. ZIRLO TM trademark property of Westinghouse Electric Company LLC.

1-2 Regulatory Commission (NRC) Safety Evaluation Report (SER) conditional requirements for the use of 422V+ fuel are provided in Section 7.0. This report serves as a reference report for the region-by-region reload transition from the Prairie Island Units 1&2 Cycle 26 core to subsequent Westinghouse cores containing the upgraded features. Representative loading patterns were evaluated to provide assurance that future reloads would meet the reload-specific criteria. Thus, this report will be used as a basic reference document in support of future Prairie Island Reload Safety Evaluations (RSEs) for upgraded fuel reloads. For the analyses, key safety parameters have been chosen to maximize the applicability of the results for future reload cycle evaluations, which will be performed utilizing the Westinghouse standard reload methodology (Reference 3). The objective of subsequent cycle-specific RSEs will be to verify that the applicable safety limits are not exceeded based on the reference analyses currently in the USAR (Reference 4) or as established in this report.

1.2 FUEL FEATURES (422V+) Prairie Island Units 1&2 Cycle 26 and subsequent cores will have the following changes relative to the existing design: Fuel assemblies that incorporate the 0.422-inch OD fuel rods and 0.422-inch OD instrumentation tubes A 400V+ style low-pressure drop (LPD) ZIRLO TM mid-grid for the 0.422-inch OD rod The 0.422-inch OD fuel rod and associated mid-grid are features that were reviewed by the NRC and are in use at Point Beach, Ginna, and KNPP. The NRC was notified of the change back to a larger OD (0.422-inch) fuel rod with VANTAGE+ fuel during the Point Beach transition (Reference 5). The NRC was then later notified of a revision to the 422V+ design with R.E. Ginna that introduced a balanced vane pattern to the mid-grid (Reference 6). The remaining VANTAGE+ features are already in use in the existing fuel at Prairie Island. For historical reference, VANTAGE+ features have been submitted to the NRC in the licensing topical report, "VANTAGE+ Fuel Assembly Reference Core Report," WCAP-12610 Appendices A through D (Reference 1), Appendix E (Reference 7), Appendices F and G (Reference 8), and associated Addenda 1 through 4 (References 9 through 12). VANTAGE+ has received generic NRC approval (References 13 through 15) for lead rod burnups up to 62,000 MWD/MTU. PERFORMANCE+ features included pre-oxidized cladding, protective bottom grids, ZIRLO TM grids, low cobalt nozzles, and a debris mitigating bottom end plug. PERFORMANCE+ features noted in Section 2.2 are currently in use in Prairie Island with the 0.400-inch OD VANTAGE+ fuel. This 422V+ fuel assembly skeleton is similar to the 14x14 400V+ fuel and 14x14 STANDARD (STD) fuel, which have been in operation for many cycles in two-loop Westinghouse plants, except for those modifications necessary to accommodate higher burnup and those modifications necessary to accommodate the 0.422-inch OD fuel rod. Since 422V+ fuel is intended to replace the Westinghouse Standard and Optimized fuel designs, the 422V+ exterior assembly envelope is similar in design dimensions (refer to Table 2-1 in Section 2), and 1-3 the functional interface with the reactor internals is equivalent to those of the Westinghouse 400V+ fuel design for which Prairie Island Units 1&2 is currently licensed. Also, the 422V+ fuel assembly is designed to be mechanically and hydraulically compatible with the Westinghouse OFA designs in full or transition cores, and the same functional requirements and design criteria previously established for the Westinghouse Optimized Fuel (References 1 and 2) remain valid for the 422V+ fuel assembly. Those functional requirements and design criteria were approved generically for all standard fuel arrays including 14x14 arrays. Table 2-1 compares the 422V+ fuel assembly to the current Westinghouse 400V+ fuel design. The Prairie Island 422V+ fuel contains fully enriched axial blankets in the top and bottom of the fuel stack. Gadolinia is dispersed within the fuel in select rods as a burnable absorber. The gadolinium rods utilize solid pellets in the blankets, whereas the non-gadolinium rods utilize annular pellets in the blanket region. The Prairie Island Units 1&2 422V+ fuel rod design, for lead rod burnups through 62,000 MWD/MTU, is based on the ZIRLO TM fuel performance models given in Reference 1 and as modified in Reference 16.

1.3 PEAKING

FACTORS The Technical Specification maximum peaking factors remain unchanged for this fuel type change. The full-power F NH peaking factor limit is 1.77. The full-power F Q N(Z) peaking factor limit is 2.50. Analyses have been completed on representative core designs that have provided assurance that future core designs are feasible with these peaking factors. These peaking factors will maintain flexibility in developing fuel management schemes for future fuel cycles that will achieve acceptable fuel economy and neutron utilization.

1.4 RTDP UNCERTAINTIES Although Prairie Island Units 1&2 are currently licensed for Revised Thermal Design Procedure (RTDP) methodology (Reference 17), revised RTDP uncertainties were considered in conjunction with the Fuel Upgrade (FU) Program. This RTDP uncertainty evaluation requires a review of temperature, pressure, power, and flow uncertainties used in the safety analysis. These uncertainties are calculated based on installed plant instrumentation or special test equipment, and on calibration and calorimetric procedures. The method of uncertainty analysis is discussed in Reference 17. The uncertainty analysis statistically combines the individual uncertainties using the square root of the sum of the squares (SRSS) method. The analysis includes uncertainties for the method of measurement, the type of field device (that is, RTDs, transmitters, special test measurements), and the calibration of the instrumentation. These uncertainties for temperature, pressure, power and flow are then used in the development of the reactor core limits and the departure from nucleate boiling ratio (DNBR) limits. The T reactor trip setpoints are then developed from the new core limits for use in the Core Operating Limits Report (COLR). Using analyses on representative cores, it has been concluded that the existing T reactor trip setpoints will be adequate for use of 422V+ fuel. Not all analyses use RTDP methodology. For those analyses that do not conform to the RTDP methodology requirements, Standard Thermal Design Procedures (STDP) are still employed. The 1-4 difference between the two methodologies is in the initial conditions for reactor coolant system (RCS) pressure, RCS temperature, and reactor power. Minimum measured flow (MMF), which is used with RTDP, is equivalent to the thermal design flow (TDF), which is used with STDP, plus a flow uncertainty. For those events using RTDP methodology, the uncertainties are statistically combined with the DNBR correlation uncertainties to obtain the overall DNBR uncertainty factor used to define the design DNBR limit. Thus, nominal values for RCS pressure, RCS temperature, and power are used for the initial conditions for the non-loss-of-coolant-accident (LOCA) events. This methodology is consistent with the reference licensing basis analyses found in the Prairie Island Units 1&2 USAR (Reference 4). Uncertainties are applied in a manner that is consistent with the analysis and is in the most conservative direction for a specific event. Table 1-1 is a summary of the RTDP uncertainties that were determined by the Nuclear Management Company (NMC) for the FU Program as compared to the original RTDP uncertainties. The uncertainties actually used in the safety analysis are equivalent or more conservative than the existing uncertainties (with the exception of the power uncertainty as noted in Table 1-1). The rationale of using slightly larger values for the uncertainties ensures conservatism in the overall analysis and allows margin for unanticipated equipment issues or other emergent issues.

1.5 PERFORMANCE

CAPABILITIES WORKING GROUP PARAMETERS The analyses bases for the Prairie Island Units 1&2 422V+ fuel upgrade are the parameters specified on the Performance Capabilities Working Group (PCWG) parameter sheets, which contain the basic thermal-hydraulic information used to complete the analysis work. These PCWG parameter sheets are used by all the analysis groups in performing their analysis to ensure a consistent set of parameters for the analyses. Note that analysts modify their actual inputs in accord ance with their procedures to obtain conservative or appropriate results. A copy of the Prairie Island Units 1&2 422V+ fuel upgrade parameter sheets are attached in Tables 1-2 and 1-3. These parameter sheets provide four cases for analysis purposes. The low and high steam generator tube plugging (SGTP) levels for the Framatome Model 56/19 Replacement Steam Generator (RSG) that will be in Unit 2 and is in Unit 1 are shown in Table 1-2 (cases 1 and 2, respectively.) The low SGTP levels (case 3) and high SGTP levels (cases 4 and 5) for the Westinghouse Model 51 original steam generator (OSG) in Unit 2 are shown in Table 1-3. The high SGTP limit represents the maximum allowable SGTP level for any single steam generator. Refer to the footnotes in Tables 1-2 and 1-3 for more details.

1.6 GENERAL

ANALYSIS ASSUMPTIONS The plant configurations assumed in the analyses are based on the existing plant configuration as confirmed by NMC. The core physics assumed in the analyses are based upon representative core loading patterns determined by Westinghouse using reasonable cycle energy requirements provided by NMC. These assumptions and analyses are used to establish the plant licensing basis as delineated in the USAR and Technical Specifications. The core physics assumptions are confirmed to be bounding for each specific reload as documented in the Reload Safety Analysis Checklists as described in Reference 3. NMC and Westinghouse mutually participate in cycle planning meetings to assure any planned plant configuration changes are considered prior to each reload. Any changes requiring reanalysis or changes to the licensing basis are thus identified with sufficient time for implementation before each reload.

1-5

1.7 CONCLUSION

S The results of the evaluations summarized within this report show that the Prairie Island Units 1&2 can operate safely and within the licensing basis as delineated in the USAR and Technical Specifications with 422V+ fuel including the transition to 422V+. This is contingent upon operating within the assumptions identified within this report and consistent with the plant configuration assumed in the supporting analysis.

1.8 REFERENCES

1. WCAP-12610/WCAP-14342 and Appendices A through D, "VANTAGE + Fuel Assembly Reference Core Report," Davidson, S. L., Nuhfer, D. L. (Eds.), June 1990. 2. WCAP-9500-A, "Reference Core Report 17x17 Optimized Fuel Assembly," Davidson, S. L. (Ed.), et al., May 1982. 3. WCAP-9272-P-A/WCAP-9273-NP-A, "Westinghouse Reload Safety Evaluation Methodology," Davidson, S. L. (Ed.), et al., July 1985. 4. "Prairie Island Updated Safety Analysis Report," Revision 29, May 2007. 5. Letter from M. F. Baumann, (WEPCO) to Document Control Desk (NRC), "14x14, 0.422" OD VANTAGE + (422V+) Fuel Design," NPL 97-0538, November 1997. 6. Letter from B. F. Maurer, (Westinghouse) to J. S. Wermiel, (NRC), "Fuel Criterion Evaluation Process (FCEP) Notification of Revision to 14x14 422 VANTAGE+ Design (Proprietary/Non-proprietary)," LTR-NRC-05-34 Rev. 1, October 2005. 7. WCAP-12610, Appendix E, "Appendix E - ZIRLOŽ High Temperature Oxidation Tests," Burman, D. L., August 1990. 8. WCAP-12610, Appendices F and G, "Appendix F - LOCA NOTRUMP Evaluation Model: ZIRLOŽ Modifications; Appendix G - Accident Evaluations LOCA Plant Specific," Kachmar, M. P., Iyengar, J., and Shimeck, D. J., December 1990. 9. WCAP-12610, Addendum 1, "Westinghouse Responses to NRC Request for Additional Information on WCAP-12610 VANTAGE + Fuel Assembly Reference Core Report," Davidson, S. L., Nuhfer, D. L. (Eds.), February 1991. 10. WCAP-12610, Addendum 2 "Westinghouse Responses to NRC Second Request for Additional Information on WCAP-12610 VANTAGE + Fuel Assembly Reference Core Report," Davidson, S. L., Nuhfer, D. L. (Eds.), February 1991. 11. WCAP-12610, Addendum 3, "Westinghouse Responses to NRC Additional Issues on WCAP-12610 VANTAGE + Fuel Assembly Reference Core Report," Davidson, S. L., Nuhfer, D. L. (Eds.), February 1991.

1-6 12. WCAP-12610, Addendum 4, "Additional Information for Appendices F and G of WCAP-12610 Appendix F - LOCA NOTRUMP Evaluation Model: ZIRLOŽ Modifications; Appendix G - Accident Evaluations LOCA Plant Specific," Kachmar, M. P., Nissley, M., and Tauche, W., May 1991. 13. Letter from A. C. Thadani (NRC) to S. R. Tritch (Westinghouse): "Acceptance for Referencing of Topical Report WCAP-12610 'VANTAGE + Fuel Assembly Reference Core Report',"

July 1, 1991. 14. Letter from A. C. Thadani (NRC) to S. R. Tritch (Westinghouse): "Acceptance for Referencing of Topical Report WCAP-12610, Appendices F, 'Appendix F - LOCA NOTRUMP Evaluation Model: ZIRLOŽ Modifications,' and G, 'Appendix G - Accident Evaluations LOCA Plant Specific'," October 9, 1991. 15. Letter from J. D. Peralta (NRC) to B.F. Maurer (Westinghouse), "Approval for Increase in Licensing Burnup Limit to 62,000 MWD/MTU (TAC No. MD1486)," May 2006. 16. Letter from N. J. Liparulo (Westinghouse) to R. C. Jones (NRC), "WCAP-12610, Appendix B, 'Extended Burnup Fuel Design Methodology and ZIRLOŽ Fuel Performance Models,' Addendum 1 [Proprietary]," ET-NRC-92-3732, August 21, 1992. 17. WCAP-11397-P-A, "Revised Thermal Design Procedure," Friedland, A. J. and Ray, S., April 1989. 18. WCAP-16206-P, "Safety Analysis Transition Program Engineering Report for the Prairie Island Nuclear Power Plant - Volume 1 Engineering Analyses," Brown, L. (Ed.), February 2004.

1-7 Table 1-1 RTDP Uncertainties Parameter Original RTDP Uncertainty Used in Analyses (1) Uncertainty Used for 422V+ Analysis Power (2) +/- 2.0% power +/- 0.5% power Reactor Coolant System Flow +/- 3.0% flow +/- 3.0% flow Pressure +/- 45 psi +/- 60.0 psi Inlet Temperature +/- 4.0 °F - 0.5°F (bias) +/- 4.0°F

- 0.5°F (bias)

Notes: 1. From WCAP-16206-P (Reference 18). 2. The power uncertainty was reduced to account for installation of a more accurate flow measurement system used in the power measurement. The RTDP analyses completed within this report were thus completed at a bounding high power level to confirm acceptable operation at any power level, including measurement uncertainties of 0.5% or more, up to 1,683 MWt.

1-8 Table 1-2 422V+ Fuel Upgrade NSSS Design Parameters - Prairie Island Units 1&2 Case 1 Case 2 Thermal Design Parameters MWt 1,690 1,690 10 6 Btu/hr 5,767 5,767 Reactor Power, MWt 1,683 (1) 1,683 (1) 10 6 Btu/hr 5,743 5,743 Thermal Design Flow, Loop gpm 89,000 89,000 Reactor 10 6 lb/hr 68.9 68.9 Reactor Coolant Pressure, psia 2,250 2,250 Core Bypass, % 6.0 (2) 6.0 (2) Reactor Coolant Temperature, °F Core Outlet 596.5 596.5 Vessel Outlet 592.6 592.6 Core Average 563.3 563.3 Vessel Average 560.0 560.0 Vessel/Core Inlet 527.4 527.4 Steam Generator Outlet 527.1 527.1 Steam Generator Steam Temperature, °F 513.2 (3) 511.0 Steam Pressure, psia 765 (3) 751 Steam Flow, 10 6 lb/hr total 7.36 (3) 7.36 Feedwater Temperature, °F 437.5 437.5 Moisture, % max. 0.10 0.10 Tube Plugging, % 0 10 Zero Load Temperature, °F 547 547 Hydraulic Design Parameters Mechanical Design Flow, gpm per loop 106,000 Minimum Measured Flow, gpm total 183,400 Notes: 1. Represents the upper limit on core thermal power with consideration of measurement uncertainty. 2. Core bypass flow accounts for thimble plug removal. 3. For analyses limited by high steam pressure, conditions corresponding to a maximum steam pressure of 778 psia, steam temperature of 515.0°F, and steam flow of 7.37 x 10 6 lb/hr are assumed. This covers the possibility that the plant could operate with better-than-expected steam generator performance.

1-9 Table 1-3 422V+ Fuel Upgrade NSSS Design Parameters - Prairie Island Unit 2 Case 3 Case 4 (1) Case 5 (2) Thermal Design Parameters MWt 1,690 1,690 1,690 10 6 Btu/hr 5,767 5,767 5,767 Reactor Power, MWt 1,683 (3) 1,683 (3) 1,683 (3) 10 6 Btu/hr 5,743 5,743 5,743 Thermal Design Flow, Loop gpm 89,000 89,000 89,000 Reactor 10 6 lb/hr 68.9 68.9 68.9 Reactor Coolant Pressure, psia 2,250 2,250 2,250 Core Bypass, % 6.0 (4) 6.0 (4) 6.0 (4) Reactor Coolant Temperature, °F Core Outlet 596.5 596.5 596.5 Vessel Outlet 592.6 592.6 592.6 Core Average 563.3 563.3 563.3 Vessel Average 560.0 560.0 560.0 Vessel/Core Inlet 527.4 527.4 527.4 Steam Generator Outlet 527.1 527.1 527.1 Steam Generator Steam Temperature, °F 505.0 (5) 499.8 495.1 Steam Pressure, psia 712 (5) 680 651 Steam Flow, 10 6 lb/hr total 7.36 (5) 7.35 7.35 Feedwater Temperature, °F 437.5 437.5 437.5 Moisture, % max. 0.25 0.25 0.25 Tube Plugging, % 0 15 (6) 25 Zero Load Temperature, °F 547 547 547 Hydraulic Design Parameters Mechanical Design Flow, gpm per loop 106,000 Minimum Measured Flow, gpm total 183,400 Notes: 1. Applicable to systems and components analysis only. 2. Applicable to safety analysis only. 3. Represents upper limit on core thermal power with consideration of measurement uncertainty. 4. Core bypass flow accounts for thimble plug removal. 5. For analyses limited by high steam pressure, conditions corresponding to a maximum steam pressure of 730 psia, steam temperature of 507.7°F, and steam flow of 7.36 x 10 6 lb/hr are assumed. This covers the possibility that the plant could operate with better-than-expected steam generator performance. 6. Supports 15% average/25% peak tube plugging.

2-1 2 MECHANICAL DESIGN FEATURES

2.1 INTRODUCTION

AND

SUMMARY

This section evaluates the mechanical design of the 14x14 422 Vantage+ (V+) fuel design and its compatibility with the currently used 14x14 400V+ fuel assembly design during the transition through mixed-fuel type core populations to cores with only 422V+ type fuel. The 422V+ fuel assembly has been designed to be compatible with the 400V+ fuel assembly, reactor internals interfaces, the fuel handling equipment, and refueling equipment. The 422V+ design dimensions are essentially equivalent (that is, the grid and nozzle envelopes, the guide thimbles internal diameters and locations, and the instrumentation tube internal diameters and locations are unchanged) to the current Prairie Island Units 1&2 400V+ assembly design from an exterior assembly envelope and reactor internals interface standpoint (refer to Table 2-1). Significant new mechanical features of the 422V+ design relative to the current 400V+ design include the following: 0.422-inch outer diameter (OD) fuel rod 0.422-inch OD instrumentation tube New 400V+ style (2.25-inch tall, vertical springs, horizontal dimples) ZIRLOŽ mid-grid; designed to be compatible with the 0.422-inch OD fuel rods and 14x14 400V+ mixed fuel cores The ZIRLO TM 422V+ mid-grid design combines enhanced anti-snag geometry and reduced pressure drop performance in an 400V+ style package. This mid-grid design evolution started with the original STANDARD (STD) Inconel mid-grid. This design was modified to an optimized fuel assembly (OFA) style Zircaloy vaneless design that has been used extensively in the Zorita Nuclear Power Plant and in Point Beach. From the Zorita design, the mid-grid was adapted for Point Beach and the Kewaunee Nuclear Power Plant (KNPP) by adding mixing vanes, changing the material composition from Zircaloy-4 to ZIRLO TM, and other minor design changes. Additional discussions of this mid-grid design are covered in subsection 2.2.2. Based on the evaluation of the 422V+ and 400V+ design differences, described in Section 2.2, it is concluded that the two designs are mechanically compatible with each other. The 422V+ fuel rod mechanical design bases remain relatively unchanged from that used for the 400V+ assemblies currently operating in Prairie Island Units 1&2 Cycle 25, other than the change in the fuel rod radial dimension (that is, 0.422-inch OD versus 0.400-inch OD). In addition, the 422V+ fuel rod mechanical design is similar to those in operation at Point Beach Units 1&2 and KNPP.

2.2 COMPATIBILITY

OF FUEL ASSEMBLIES Table 2-1 compares the STD, 400V+, and 422V+ design parameters for Prairie Island Units 1&2. Figure 2-1 depicts the 400V+ and 422V+ assembly designs noting respective overall height and grid elevation dimensions. The 422V+ assembly skeleton is similar to that previously described for 400V+ (Reference 1), except for the new mid-grid design. The 400V+ top and bottom nozzles will be used in the 422V+ fuel assembly. The grid centerline elevations of the 422V+ are the same as those of the 400V+ assembly.

2-2 Crossflows can result from a mismatch in grid loss coefficients if different grid types are used in adjacent assemblies. The difference in grid loss coefficients between the two fuel assembly designs is within what Westinghouse has evaluated in the past for various fu el designs/transitions and will not result in any significant crossflows between the two assembly designs. The hydraulic compatibility effects are discussed further in Section 4.0.

Since the 422V+ fuel is intended to replace either the Westinghouse STD or the 400V+ fuel assembly designs, the 422V+ exterior assembly envelope is essentially equivalent in design dimensions, and the functional interface with the reactor internals is also equivalent to those of previous Westinghouse fuel designs. The 422V+ fuel assembly is designed to be mechanically and hydraulically compatible with the STD and 400V+ in full or transition cores, and the same functional requirements and design criteria, as previously established for the Westinghouse 400V+ fuel assembly, remain valid for the 422V+ fuel assembly.

2.2.1 Fuel Rods Figure 2-2 shows a comparison of the 400V+ fuel rod to that of the 422V+ fuel rod design. The 422V+ fuel rod has the same clad-wall thickness as the 400V+ fuel rod. The 422V+ fuel rod length is longer to provide the additional plenum volume space for fission gas release at the extended burnup. To provide additional volume, the 422V+ fuel rod has a shorter stack height. The bottom end plug has an internal grip feature to facilitate rod loading on both designs (422V+ and 400V+) and provides appropriate lead in for the removable top nozzle reconstitution feature. The Prairie Island 422V+ fuel rod also has a zirconium oxide coating at the bottom end of the fuel rod. A metallurgical-bonded layer of ZrO 2 uniformly covers a minimum of 4.5 inches of the bottom of the end of the fuel rod. The minimum coating length was chosen to ensure that the coating would extend through the top of the current bottom Inconel structure grid, independent of fuel rod loading position or fuel assembly design. The extra layer of zirconium oxide provides additional rod fretting wear protection early in life, before the natural oxide layer can be built up from in reactor operations. Th e 400V+ and 422V+ fuel also have fully-enriched annular pellets in the axial blankets. The 422V+ design for the gadolinium (containing reduced central enrichment and axial blankets with reduced enrichment solid pellets) and non-gadolinium fuel (containing axial blankets with fully-enriched annular pellets) will behave in a similar manner to 400V+ fuel of comparable enrichments and burnups, with regards to xenon stability, load follow capability, peaking factors, rod worths, and trip reactivity. The annular pellets of the non-gadolinium fuel provide additional plenum volume for fission gas releases. The key design difference between enriched annular pellets in axial blankets and the enriched solid fuel pellets is the annulus itself. The annulus volume is approximately 25 percent of the total pellet volume.

Annular pellets in axial blankets have the same chamfer as the enriched fuel pellets, but no dish on the pellet ends. Pellet length-to-diamete r ratio is maintained at approximately 1.4. This ratio has been adjusted for the use of an even multiple of pellet lengths to obtain appropriate axial blanket zone lengths in fabrication. A reduction in the fuel stack length of 0.75 inches along with 6 inches of annular pellets in axial blankets will be used in the 422V+ fuel design for the non-gadolinium fuel. This stack reduction is incorporated to provide additional plenum volume space for gas releases.

2-3 The relatively low range of linear heat rate that the annular pellets in axial bl ankets will experience, and the modest fraction of the fuel volume that it occupies, ensures that their use will not have any significant effect on the limiting fuel temperature or rod internal pressure, other than that due to the additional void volume provided by the axial blanket pellet annulus (Reference 2). As noted in Table 2-1, the fuel pellet-to-clad gap has increased from 3.5 mils for the 400V+ fuel product to 3.75 mils for the 422V+. The impact of the slightly larger gap would be to delay the point in time when the cladding would make contact with the fuel pellet.

The gadolinium burnable absorber rod assemblies have Gd 2 O 3 dispersed in the UO 2 matrix. The gadolinium fuel rods provide the advantages of permitting longer fuel cycles and higher batch discharge exposures, permitting shaping of the radial power distribution, and increasing the flexibility of fuel management schemes. It should be noted that the axial blankets will be 5.9 inches for the gadolinium rods. The mechanical design analysis for the gadolinium burnable absorber rods is enveloped by the analysis for the non-gadolinium fuel rods. The enrichment in these rods is reduced so that temperatures, stresses and strains in the absorber rods are not limiting The 422V+ fuel rod design bases and evaluation are given in Section 2.0 of Reference 3.

2.2.2 Grid Assemblies The top and bottom grids of the 422V+ fuel assemblies are Inconel (non-mixing vane) and are similar in design to the current Inconel grids used with the 400V+ fuel assemblies, except for the cell sizing changes required to accommodate the larger diameter fuel rod and instrumentation tube. The remainder of the grids are manufactured with ZIRLO TM and the interlocking strap and grid

/sleeve joints are laser welded, whereas the top and bottom Inconel grid joints are brazed. See Table 2-1 for additional information on the grids. The other significant change between the two designs, other than ZIRLO TM, is the development of a low-pressure drop (LPD) mid-grid. This mid-grid design evolution started with the original STD Inconel mid-grid. This design was modified to an OFA-style Zircaloy vaneless design that has been used extensively in the Zorita nuclear power plant. From the Zorita design, the mid-grid was adapted to the Point Beach and KNPP design by adding mixing vanes, changing the material composition from Zircaloy-4 to ZIRLO TM, and other minor design changes. This new mid-grid design has been developed and installed under the guidelines of the Fuel Criteria Evaluation Process (FCEP) (References 4 and 5). By complying with the requirements of the FCEP, it has been demonstrated that the new mid-grid meets all of the design criteria of the existing tested mid-grids that form the basis of the WRB-1 correlation database and that the WRB-1 correlation with a 95/95 correlation limit of 1.17 applies to the new mid-grid. This FCEP applicability demonstration was presented to the Nuclear Regulatory Commission (NRC) in a meeting held on August 5, 1997. The slides associated with this presentation, which provide more information on this change, were formally transmitted to the NRC via Reference 6. The 422V+ grid assembly design bases and evaluation are given in Section 2.3 of Reference 3 and confirmed in Reference 7.

2-4 2.2.3 Guide Thimble and Instrumentation Tubes The guide thimble diameter and dashpot length of the 422V+ guide thimbles are identical to those of the 400V+ design. The 422V+ guide thimble tube inside diameter (ID), as in the 400V+, provides a nominal diametral clearance of 0.057 inches for the control rods. The 422V+ instrumentation tube diameter has been increased from 0.400 inches to 0.422 inches when compared to the 400V+ fuel assembly design. The general design bases for the 422V+ guide thimble and instrumentation tubes remain the same as for those given in Reference 8.

2.2.4 Reconstitutable

Top Nozzle The 422V+ reconstitutable top nozzle (RTN) will be fabri cated of low cobalt 304 stainless steel, to reduce as-low-as-reasonably-achievable (ALARA) concerns. The RTN and low cobalt 304 stainless steel are currently used in existing 400V+ fuel assembly design. The 422V+ RTN is in use in Point Beach.

2.2.5 Debris

Filter Bottom Nozzle The debris filter bottom nozzle (DFBN) will also be fa bricated of low cobalt 304 stainless steel to reduce ALARA concerns. The low cobalt 304 stainless steel and the bottom nozzle geometries (that is, flow holes, thimble holes, chamfers, counterbores, width, and height) are identical to those of the current 14x14 DFBN, with the exception of the increased counterbore diameter in the instrument sheath and through-hole diameter to accommodate the 0.422-inch OD instrumentation tube.

2.3 MECHANICAL

PERFORMANCE Design changes associated with the 14x14 422V+ design do not significantly influence the fuel assembly structural characteristics that were determined by prior mechanical testing, as previously addressed in Section 2.2. Therefore, the 422V+ fuel assembly, with expected structural behavior and projected performance, will meet design requirements throughout the fuel's life.

2.4 FUEL ROD PERFORMANCE Fuel rod performance for 422V+ Prairie Island Units 1&2 fuel is similar to 422V+ fuel in use at Point Beach Units 1&2, which has previously been shown to satisfy the NRC Standard Review Plan (SRP) (Reference 6) fuel rod design criteria on a region-by-region basis. The design bases for Westinghouse 422V+ fuel are discussed in Reference 3. There is no impact from a fuel rod design standpoint of having fuel with more than one type of geometry simultaneously residing in the core during the transition cycles. The mechanical fuel rod design evaluation for 422V+ fuel incorporates all appropriate design features of the region, including any changes to the fuel rod or pellet geometry from that of previous fuel regions (such as changes in the fuel rod diameter and plenum length). Fuel performance evaluations have been completed for 422V+ fuel to demonstrate that the design criteria will be satisfied in the core under the planned operating conditions using representative core loading patterns. Any additional changes from the plant operating conditions originally evaluated for the mechanical design will be addressed for all affected 422V+ fuel regions as 2-5 part of the cycle-specific reload safety evaluation process when the plant changes are to be implemented. The impacts of using cycle-specific loading patterns will also be addressed in the cycle-specific reload safety evaluation process. Fuel rod design evaluations for the 422V+ fuel for both the gadolinium and non-gadolinium fuel rod designs were performed using NRC approved models (References 3 and 9) and NRC-approved design criteria and methods (References 10 through 13) to demonstrate that all fuel rod design criteria are satisfied. The fuel rod design criteria given below are verified by evaluating the predicted performance of the limiting fuel rod, defined as the rod that gives the minimum margin to the design limit. In general no single rod is limiting with respect to all the design criteria. Generic evaluations have identified which rods are most likely to be limiting for each criterion, and exhaustive screening of fuel rod power histories to determine the limiting rod is typically not required. The NRC-approved PAD 4.0 code, which incorporates models (References 3 and 9) for in-reactor behavior, is used to calculate the fuel rod performance over its irradiation history. PAD is the principal design tool for evaluation of fuel rod performance. PAD iteratively calculates the interrelated effects of temperature, pressure, clad elastic and plastic behavior, fission gas release and fuel densification and swelling as a function of time and linear power. PAD 4.0 is a best-estimate plus uncertainties approach. A statistical convolution of individual uncertainties due to design model uncertainties, fabrication uncertainties, and dimensional uncertainties is used. Fuel Rod Design Criteria The criteria pertinent to the fuel rod design as documented in Reference 5 are as follows: Rod internal pressure

• Cladding flattening Cladding stress and strain

• Fuel rod axial growth Cladding oxidation and hydriding

• Plenum cladding support Fuel temperature

• Cladding free-standing Cladding fatigue

• End-plug weld integrity Each of these key fuel rod design criteria has been evaluated for use of the Westinghouse 422V+ fuel assembly design in Prairie Island Units 1&2. Based on these evaluations, it is concluded that each design criterion can be satisfied through transition cycles to a full core of the 422V+ design. The design criteria are described in more detail below.

2-6 Rod Internal Pressure The internal pressure of the lead fuel rod in the r eactor will be limited to a value below that could cause: The diametral gap to increase due to outward cladding creep during steady-state operation Extensive departure from nucleate boiling (DNB) propagation to occur. The rod internal pressure for the Prairie Island 422V+ fuel rods has been evaluated by modeling the gas inventories, gas temperature, and rod internal volumes through the rods' life. The resulting rod internal pressure is compared to the design limit on a case-by-case basis of current operating conditions to end of life (EOL). This evaluation showed that the rod internal pressure satisfies the design limit. The second part of the rod internal pressure design basis precludes extensive DNB propagation and associated fuel failure. The basis for this criterion is that no significant additional fuel failures due to DNB propagation will occur in cores that have fuel rods operating with rod internal pressure in excess of system pressure. The design limit for Condition II events is that DNB propagation is not extensive, that is, the process is shown to be self-limiting and the number of additional rods in DNB due to propagation is relatively small. For Condition III/IV events, it is shown that the total number of rods in DNB, including propagation effects, is consistent with the assumptions used in radiological dose licensing bases for the event under consideration. Cladding Stress and Strain The standard cladding stress criterion is applied for the non-gadolinium fuel, which is defined as the volume average effective cladding stress with the Von Mises equation considering interference due to uniform cylindrical pellet-cladding contact caused by pellet thermal expansion, pellet swelling and uniform cladding creep, and pressure differences, is less than the 0.2 percent offset yield stress with due consideration to temperature and irradiation effects under Condition I and II events. The revised cladding stress criterion approved by the NRC, as documented in Reference 12, is applied for the gadolinium fuel, which is defined as: The maximum cladding stress intensities, excluding pellet cladding interaction, but accounting for cladding corrosion as a loss-of-load carrying metal, be less than the stress limit, as defined based on the American Society of Mechanical Engineers (ASME) code calculations. The 1-percent transient cladding strain criterion is met (established as the most limiting design criterion for transient cladding stress/strain criteria). An additional steady-state cladding strain criterion based on the total (plastic plus elastic strain) is met. No centerline fuel melting occurs. The effect of the plastic deformation is accounted for in all fuel rod design criteria as appropriate.

2-7 The design limit for cladding strain during steady-state operation is that the total plastic tensile creep strain due to uniform cladding creep and uniform cylindrical fuel pellet expansion associated with fuel swelling and thermal expansion is less than 1 percent from the unirradiated condition. The design limit for fuel rod cladding strain during Condition II events is that the total tensile strain due to uniform cylindrical pellet thermal expansion is less than 1 percent from the pre-transient value. The Westinghouse PAD 4.0 fuel performance methods and models (References 3, 5, 9, 10, 12, and 13) were used to evaluate fuel rod cladding stress and strain limits for the 422V+ designs using representative loading patterns. The local power duty during Condition II events was a key factor in evaluating the margin to fuel cladding stress and strain limits. The fuel duty for the FU Program conditions was more limiting, resulting in an increase in the cladding stress and strain levels. The fuel rod cladding stress and strain evaluations showed that the 400V+ and 422V+ fuel rod designs at the FU analyzed operating conditions for the partial core transition cycles and the longer cycle length equilibrium cycle meet the cladding stress and strain fuel rod design limits. The overall fuel analyses results show that the calculated fuel cladding strain is less than the limit, thereby demonstrating that the 400V+ and 422V+ fuel rod designs with the appropriately designed fuel cycling scheme will not preclude the fuel's capability to meet the cladding stress and strain limits. Cladding Oxidation and Hydriding The design criteria related to cladding corrosion require that the ZIRLO TM cladding metal-oxide interface temperature be maintained below specified limits to prevent a condition of accelerated oxidation which would lead to cladding failure. The calculated cladding temperature (metal-oxide interface temperature) will be less than 780°F during steady-state operation and for Condition II transients, the calculated cladding temperature will not exceed 850°F for ZIRLO TM cladding. The cladding surface temperatures were evaluated at the FU analyzed operating conditions and the resulting metal-oxide interface temperatures satisfy the above temperature limits. The best-estimate hydrogen pickup level in the ZIRLO TM cladding will be less than or equal to 600 ppm on a volumetric average basis at EOL. The hydrogen pickup criterion, which limits the loss of ductility due to hydrogen embrittlement that occurs upon the formation of zirconium hydride platelets, has been met with the current approved model for the Prairie Island FU Program. Fuel Temperature For Condition I and II events, the fuel and reactor protection systems are designed to ensure that the calculated centerline fuel temperature does not exceed the fuel melting temperature criterion. The intent of this criterion is to avoid a condition of gross fuel melting, which can result in severe duty on the cladding. The concern here is based on the large volume increase associated with the phase change in the fuel and the potential loss-of-cladding integrity as a result of molten fuel/cladding interaction. The temperature of the fuel pellets was evaluated as a function of local power and burnup by modeling the fuel rod geometry, thermal properties, heat flux, and temperature differences in order to calculate fuel surface, average, and centerline temperatures. The fuel surface and average temperatures with associated rod internal pressure are provided to transient analysis and loss-of-coolant-accident (LOCA) for accident 2-8 analysis of the 422V+ fuel design (see Section 4.5 for additional information). The fuel centerline temperature is used to show that fuel melt will not occur.

Cladding Fatigue Acceptable cladding fatigue performance requires that, for a given strain range, the number of strain fatigue cycles is less than those required for failure, considering a factor of safety of 2.0 on stress amplitude and a factor of 20.0 on the number of cycles. Cladding fatigue for the Prairie Island 422V+ fuel was evaluated by using a limiting fatigue duty cycle consisting of daily load follow maneuvers. The evaluation showed that the cumulative fatigue usage factor is less than the design limit. Cladding Flattening Cladding, fuel pellet, and fuel rod initial backfill pressure parameters are defined to prevent fuel rod failures due to long-term creep collapse of the fuel rod cladding into axial gaps formed within the fuel stack. Current fuel rod designs employing fuel with improved in-pile stability provide adequate assurance that axial gaps large enough to allow cladding flattening will not form within the fuel stack. The NRC has approved WCAP-13589-A (Reference 11), which provided data to confirm that significant axial gaps in the fuel column due to densification (and therefore cladding flattening) will not occur in current Westinghouse fuel designs. The Prairie Island 422V+ fuel meets the criteria for applying the Reference 11 methodology and, therefore, cladding flattening will not occur. Fuel Rod Axial Growth This criterion assures that sufficient axial space exists to accommodate the maximum expected fuel rod growth without degradation of the assembly function. Fuel rods are designed with adequate clearance between the fuel rod and the top and bottom nozzles to accommodate the differences in the growth of fuel rods and the growth of the fuel assembly to preclude interference between these members. The Prairie Island fuel rod growth evaluation demonstrates that there is adequate margin to the fuel rod growth design limit for the 422V+ fuel. Plenum Cladding Support This criterion assures that the fuel cladding in the plenum region of the fuel rod will neither collapse during normal operating conditions, nor distort so as to degrade fuel rod performance. The helical coil spring used in the 422V+ fuel design prevents potential cladding collapse by providing cladding support. Cladding Free Standing The cladding free-standing criterion requires that the cladding shall be short-term free standing at beginning of life, at power, and during hot hydrostatic testing. This criterion precludes the instantaneous collapse of the cladding onto the fuel pellet caused by the pressure differential across the clad wall.

2-9 Evaluations of the cladding-free-standing criterion have shown that instantaneous collapse of the cladding onto the fuel will be precluded for differential pressures well in excess of the maximum expected differential pressure for the 400V+ and 422V+ fuel rod designs for the transition cycles and the equilibrium 422V+ cycles evaluated. End Plug Weld Integrity The fuel rod end plug weld shall maintain its integrity during Condition I and II events and shall not contribute to any additional fuel failures above those already considered for Condition III and IV events. The intent of this criterion is to ensure that fuel rod failures will not occur due to tensile pressure differential loads which can exist across the weld.

For Prairie Island Units 1&2, the results of the evaluation confirm that the end plug integrity will be maintained and the fuel system will not be damaged due to excessive end plug weld tensile pressure differential loads for the 400V+ and 422V+ fuel rod designs for the transition cycles and the equilibrium 422V+ cycles evaluated.

2.5 SEISMIC/LOCA IMPACT ON FUEL ASSEMBLIES The 14x14 422V+ fuel assembly design is comparable to the 14x14 400V+ fuel assembly design. Seismic/LOCA analyses demonstrated adequate grid load margin on all fuel assemblies except for the unrodded fuel assemblies on the periphery of the 3 fuel assembly array. Evaluations have demonstrated that the core coolable geometry and control rod insertion requirements are met. An evaluation of the 14x14 422V+ fuel assembly structural integrity, considering the lateral effects of two LOCA auxiliary line breaks (accumulator and pressurizer surge line) and both a safe shutdown earthquake (SSE) and an operating basis earthquake (OBE) seismic accident have been performed. The OBE/SSE/LOCA analysis results were obtained using the time-history numerical integration technique. The maximum grid impact forces obtained from both the SSE and the LOCA transients were combined using the square root of the sum of the square (SRSS) method. The maximum loads were compared with the allowable grid crush strengths. This analysis is discussed in more detail in the following paragraphs.

2.5.1 Fuel Assembly and Reactor Core Models Based on the assembly vibration frequencies and mode shapes, a parametric study was performed using NKMODE. NKMODE calculates a set of equivalent spring-mass elements representing an individual fuel assembly structural system. Based on this model, it has been shown that the mode shapes agree well with the predominate fuel assembly vibration frequencies. The lumped mass-spring fuel assembly model was further verified using the WECAN finite element code. With the appropriate analysis parameters such as grid impact stiffness and damping, number of fuel assemblies in a planar array and gap clearance established, the WEGAP reactor core model was used for analyzing the transient loads.

2-10 2.5.2 Grid Load Analysis Homogeneous Core The maximum SSE and LOCA results for the 14x14 422V+ fuel assembly occur in the peripheral assemblies of the 3 fuel assembly arrays. The maximum fuel assembly deflection was 0.6334 inches. The maximum grid impact forces obtained from the SSE and LOCA analyses were combined using the SRSS method and were compared to the 2-directional la teral allowable grid strengths associated with the 14x14 422V+ fuel assembly. The results indicate that the combined maximum impact force is less than the allowable grid strength (except for the outer two assemblies of the 3 fuel assembly array that experience grid crush) and 87.1 percent of the respective allowable grid strengths. The allowable grid strengths are established at a 95 percent confidence level on the true mean from the distribution of experimentally determined grid crush data at the operating temperature. Mixed Core The maximum SSE and LOCA results for the 14x14 422V+ fuel assembly occur in the peripheral assemblies of the 3 fuel assembly arrays. The maximum fuel assembly deflection was 0.7065 inches. The maximum grid impact forces obtained from the SSE and LOCA analyses were combined using the SRSS method and were compared to the 2-directional la teral allowable grid strengths associated with the 14x14 422V+ fuel assembly. The results indicate that the combined maximum impact force is less than the allowable grid strength (except for the outer two assemblies of the 3 fuel assembly array that experience grid crush), and 93.2 percent of the respective allowable grid strengths. The allowable grid strengths are established at a 95 percent confidence level on the true mean from the distribution of experimentally determined grid crush data at the operating temperature. Stress The fuel assembly displacement is limited by the total accumulated gap clearances plus the grid deformations. Fuel assembly stresses were calculated based on the most limiting case. Stresses were calculated based on the maximum vertical impact load and the maximum fuel assembly lateral deflection, and operating load conditions. The results indicate that adequate margins for both the fuel rods and the thimble tubes exist, so that fragmentation of the fuel rods should not occur. The reactor can be safely shutdown under faulted condition loading. In conclusion, the 14x14 422V+ assembly design is structurally acceptable under the OBE event and for the combined SSE seismic and LOCA loadings for both Prairie Island plants at the FU analyzed conditions.

2.5.3 Conclusions

The maximum horizontal input motion congruent with the core principal axis is used to determine dynamic fuel responses. The reactor core is analyzed as a de-coupled system with respect to the two lateral directions. The input forcing function is obtained from a se parate reactor pressure vessel and reactor internals system analysis. The evaluation of the 422V+ fuel assembly in accordance with NRC requirements as given in Standard Review Plan (SRP) 4.2 (Reference 14), Appendix A(4), shows that the 422V+ fuel is structurally 2-11 acceptable for the Prairie Island reactors at the FU analyzed conditions. The grid loads evaluated for the LOCA and SSE seismic events and combined by the SRSS method identified in SRP 4.2 are less than the allowable limits (except for the 3 fuel assembly arrays). An evaluation was performed that demonstrated that the core coolable geometry requirement is satisfied in the presence of the few crushed grids in unrodded peripheral assemblies. The same conclusion is true for a transition core composed of both 422V+ fuel assemblies and 400V+. Thus, the requirement of ensuring that a core coolable geometry is maintained is met. This conclusion is also valid for the OBE event as no grid crush is predicted. The stresses in the fuel assembly components resulting from seismic and LOCA-induced deflections are within acceptable limits. The reactor can be safely shut down under the combined faulted condition loads. In conclusion, the 14x14 422V+ assembly design is structurally acceptable under the OBE event and for the combined SSE seismic and LOCA loadings for Prairie Island Units 1&2 at the FU analyzed conditions.

2.6 REFERENCES

1. WCAP-9500-A, "Reference Core Report 17x17 Optimized Fuel Assembly," Davidson, S. L. (Ed.), et al., May 1982. 2. WCAP-14710, "1-D Heat Conduction Model for Annular Fuel Pellets," Shimeck, D. J., September 1996. 3. WCAP-12610-P-A, "VANTAGE + Fuel Assembly Reference Core Report," Davidson, S. L., Nuhfer, D. L. (Eds.), April 1995. 4. WCAP-12488-A (Proprietary), WCAP-14204-A (Non-Proprietary), "Westinghouse Fuel Criteria Evaluation Process," Davidson, S. L. (Ed.), et al., October 1994. 5. WCAP-12488-A, Addendum 1-A, Revision 1, "Addendum 1 to WCAP-12488-A Revision to Design Criteria (Westinghouse Fuel Criteria Evaluation Process)," January 2002. 6. Letter from M. F. Baumann, (WEPCO) to Document Control Desk (NRC), "14x14, 0.422" OD VANTAGE + (422V+) Fuel Design," NPL 97-0538, November 1997. 7. Letter from B. F. Maurer, (Westinghouse) to J. S. Wermiel, (NRC), "Fuel Criterion Evaluation Process (FCEP) Notification of Revision to 14x14 422 VANTAGE+ Design (Proprietary/Non-proprietary)," LTR-NRC-05-34 Rev. 1, October 2005. 8. "Prairie Island Updated Safety Analysis Report," Revision 29, May 2007. 9. WCAP-15063-P-A with errata, "Westinghouse Improved Performance Analysis and Design Model (PAD 4.0)," Foster, J. P., Sidener, S., July 2000. 10. WCAP-10125-P-A (Proprietary), "Extended Burnup Evaluation of Westinghouse Fuel," Davidson, S. L. (Ed.), et al., December 1985.

2-12 11. WCAP-13589-A, "Assessment of Clad Flattening and Densification Power Spike Factor Elimination in Westinghouse Nuclear Fuel," Kersting, P. J., et al., March 1995. 12. WCAP-10125-P-A Addendum 1-A, Revision 1-A, "Extended Burnup Evaluation of Westinghouse Fuel, Revision to Design Criterion," May 2005. 13. WCAP-9272-P-A, "Westinghouse Reload Safety Evaluation Methodology" Davidson, S. L. (Ed.), et al., July 1985. 14. NUREG-0800, "Standard Review Plan - 4.2 - Fuel System Design," Revision 3, March 2007.

2-13 Table 2-1 Comparison of 14x14 STD, 400V+, 422V+ Fuel Assembly Mechanical Design Parameters STD 400V+ 422V+ Fuel Assembly Overall Length, inch 159.975 159.775 159.775 Fuel Rod Overall Length, inch 151.850 (1) 151.914 152.563 Nominal Assembly Envelope at Bottom Nozzle, inch 7.761 7.761 7.761 Fuel Rod Pitch, inch 0.556 0.556 0.556 Number of Fuel Rods/Assembly 179 179 179 Number of Guide Thimbles/Assembly 16 16 16 Number of Instrumentation Tubes/Assembly 1 1 1 Fuel Tube Material Zircaloy-4 ZIRLOŽ ZIRLO Fuel Tube Cladding OD, inch 0.422 0.400 0.422 Fuel Rod Cladding Thickness, inch 0.0243 0.0243 0.0243 Fuel Cladding Gap, mil 3.75 3.5 3.75 Enriched Fuel Pellet diameter, inch 0.3659 0.3444 0.3659 Enriched Fuel Pellet length, inch 0.4390 0.4130 0.4390 Annular Axial Blanket Pellet Diameters ID, inch N/A 0.1720 0.1830 OD, inch N/A 0.3444 0.3659 Annular Axial Blanket Pellet Length, inch N/A 0.5000 0.5000 Fuel Rod End Plugs Standard Tapered and radiused Tapered and radiused Fuel Stack Height (cold, undensified), inch 144 144 143.25 Annular Axial Blanket Length (top and bottom), inch N/A 6 6 Plenum volume, inch 3 0.7017 0.6126 0.8546 Guide Thimble Material Zircaloy-4 ZIRLOŽ ZIRLO Guide Thimble OD, inch 0.539 0.526 0.526 Guide Thimble Wall Thickness, inch 0.017 0.017 0.017 Grid Material, Inner Mid-grid (5) Inconel Zircaloy-4 ZIRLO Edges Modified No Yes Yes 2-14 Table 2-1 Comparison of 14x14 STD, 400V+, 422V+ Fuel Assembly Mechanical Design Parameters (cont.) STD 400V+ 422V+ Grid Material, End Grids (2) Inconel Inconel Inconel Inner Spring (Mid-Grids) Vertical Vertical Vertical Grid Fabrication Inconel Grids Brazed joining of interlocking stamped straps Brazed joining of interlocking stamped

straps Brazed joining of interlocking stamped

straps Zircaloy-4 Mid-Grids None Laser weld joining of interlocking stamped straps None ZIRLO Mid-Grids None None Laser weld joining of interlocking stamped straps Grid/Guide Thimble Attachment Inconel Grids Thimbles bulged together with sleeve prebrazed Thimbles bulged together with sleeve prebrazed Thimbles bulged together with sleeve prebrazed Zircaloy-4/ZIRLO Mid-Grids None Thimbles bulged together with sleeves laser prewelded to grid straps Thimbles bulged together with sleeves laser prewelded to grid straps Relative Cladding Thickness/ Diameter Ratio 0.948 1.00 0.948 H 20/UO 2 Volume Ratio for Assembly 1.902 2.302 1.902 Relative UO 2/Rod 1.129 1.00 1.123 Top Nozzle Welded stainless steel standard Reconstitutable

stainless steel, reduced height, removable design, low cobalt Reconstitutable stainless steel, reduced height, removable design, low cobalt Compatible with Fuel Handling Equipment Yes Yes Yes Note: 1. Typical value.

2-15 Figure 2-1. 14x14 400V+ and 422V+ Fuel Assembly Designs*

• All units are in inches.

400V+ 422V+

2-16 Figure 2-2. 14x14 400V+ and 422V+ Fuel Rod Comparison 151.914 IN 3-1 3 NUCLEAR DESIGN

3.1 INTRODUCTION

AND

SUMMARY

The effects of using the Westinghouse 422 Vantage+ (V+) fuel features on the nuclear design bases and methodologies for the Prairie Island Units 1&2 are evaluated in this section. The 422V+ fuel design differs from that of 400V+ as described in Section 2 of this Licensing Report. Two key features in the 422V+ design relative to the current 0.400V+ optimized fuel assembly (OFA) design are: 1) an increase in the fuel rod cladding outer diameter from 0.400 inches to 0.422 inches; and 2) a change in the fuel stack height within the assembly. The 422V+ fuel assembly will have a fuel stack height reduction of 0.75 inches to accommodate fission gas release from the extended burnups of the 422V+ design. The specific values of core safety parameters, such as power distributions, peaking factors, rod worths, and reactivity parameters, are primarily loading-pattern dependent. The variations in the loading-pattern-dependent safety parameters are expected to be typical of the normal cycle-to-cycle variations for the standard fuel reloads. Key safety parameter limits (that is, departure from nucleate boiling [DNB], peak cladding temperature, peak linear heat rate) are maintained for the 422V+ fuel design. Standard nuclear design analytical models and methods (References 1 through 3) accurately describe the neutronic behavior of the 422V+ fuel design. Storage of the 422V+ upgraded fuel at Prairie Island Units 1&2 was reviewed with respect to criticality effects, as well as the spent fuel pool gamma heating. Spent fuel pool criticality analyses for the 422V+ fuel were reviewed by Westinghouse and confirmed bounding for storage of 422V+ fuel (Reference 4). The evaluation concluded that the spent fuel pool Keff remains below the 0.95 limit for 422V+ fuel. Westinghouse also reviewed criticality evaluations for the new fuel vault (Reference 5) for 422V+ fuel with a maximum enrichment of 5.0 w/o U-235. The evaluation concluded the new fuel vault Keff remains below the 0.95 limit for the fully flooded condition, and 0.98 if moderated by aqueous foam. Therefore, the storage of 422V+ fuel meets the required criteria for spent fuel and new fuel storage.

3.2 DESIGN

BASIS The specific design bases and their relation to the Ge neral Design Criteria (GDC) in the Code of Federal Regulations (CFR) Section 10 CFR 50, Appendix A, for the 422V+ design are the same as those of the 400V+ design (Section 3.1 of Reference 6).

For the 422V+ product, the fuel burnup design was analyzed to a lead rod burnup of up to 75,000 MWD/MTU (Note: V+ is currently licensed to 60,000 MWD/MTU by the Nuclear Regulatory Commission (NRC) (Reference 6) with extension to 62,000 MWD/MTU (Reference 7). The effects of extended burnup on nuclear design parameters have been previously discussed in Reference 8. That discussion is valid for the anticipated 422V+ design discharge burnup level. In accordance with the NRC recommendation made in their review of Reference 8, Westinghouse will continue to monitor predicted versus measured physics parameters for extended burnup applications.

3-2 3.3 METHODOLOGY The purpose of this reload transition core analysis is to determine, prior to the cycle-specific reload design, if the assumed values for the key safety parameters are adequate for the transition to 422V+ fuel upgrade using representative core designs. This will provide assuran ce that future reload specific analyses validation efforts will not result in any unanticipated conditions that require extensive re-analysis or alternate core loading patterns just prior to installing the cycle specific core design. It also provides assurance the plant licensing basis as defined in Technical Specifications and the updated Safety Analysis Report (USAR) are adequate for anticipated operations with 422V+ fuel. No changes to the Westinghouse nuclear design philosophy, methods, or models are necessary for the transition to the 422V+ fuel design. Key safety parameters are established as part of the reference safety analysis. These key safety parameters are then compared to the record of analysis for a reload cycle. The reference safety analysis or record of analysis is established for major changes to either the fuel assembly design or major changes to the plant, or changes to the safety analyses to recapture margins. If one or more of the key safety parameters fall outside of the bounds assumed in the reference safety analysis, the affected transients will be re-evaluated or re-analyzed using standard methods, and results documented in the Reload Safety Evaluation (RSE) for that cycle. The 0.422 inch outer diameter (OD) fuel rod has had extensive nuclear design and operating experience with the 14x14 STANDARD (STD) fuel assembly design, which has extensive history in the Point Beach Units 1&2, Prairie Island Units 1&2, and Kewaunee. ZIRLOŽ material has also had extensive nuclear design and operating experience with other fuel assembly designs including the existing fuel in Prairie Island Units 1&2. These changes have a negligible effect on the use of standard nuclear design analytical models and methods to accurately describe the neutronic behavior of the 422V+ fuel (Reference 6).

3.4 DESIGN

EVALUATION - PHYSICS CHARACTERISTICS AND KEY SAFETY PARAMETERS Multiple cycles of representative core designs were established to model th e transition to a full 422V+ fueled core. These models incorporate 422V+ style low-pressure drop (LPD) ZIRLOŽ mid-grids; ZIRLOŽ clad fuel rods; ZIRLOŽ fabricated guide thimble tubes and instrumentation tubes; and assembly dimensional and fuel rod design modifications. Typical loading patterns were developed based on projected energy requirements for Prairie Island Units 1&2. For the third transition cycle, an additional loading pattern wa s developed based upon a longer than normal cycle length. This loading pattern was intended to be evaluated for an extended cycle length that utilizes higher than typical boron letdown curves to identify any potential analyses limitations with minimal margin to acceptance criteria. These models are not intended to represent limiting loading patterns. They were developed with the intent to show that enough margin exists between typical safety parameter values and the corresponding limits to allow flexibility in designing actual reload cores. Four representative core models were developed and used for the majority of calculations performed here.

3-3 The first "transition" cycle model is used to capture the initial and predominant transition effects. A second transition core and third "all 422V+" core models were developed and used to capture the core characteristics when a full core of the 422V+ fuel is present. Figures 3-1, 3-3, 3-5, and 3-7 provide the fuel loading and assembly exposures at beginning of cycle (BOC) and end of cycle (EOC) are summarized for all three transition cycles, including both the short (Pattern A) and long cycle lengths (Pattern B) for the third transition cycle. Assembly power distributions at BOC, middle of cycle (MOC), and EOC are also provided for each model in Figures 3-2, 3-4, 3-6, and 3-8. Comparisons of key core parameters versus cycle length for the models are provided in Figures 3-9 through 3-12. These include critical boron concentration, axial offset, hot rod (F NH), and total peaking factor ()Z (F N Q), respectively. Table 3-1 provides the key safety parameter ranges that were used or confirmed as part of the transition core analysis. The changes in fuel design were accounted for in the transition core analysis.

3.5 DESIGN

EVALUATION - POWER DISTRIBUTIONS AND PEAKING FACTORS Figure 3-11 shows a comparison of the radial peaking factors (F NH) between the core models used. The limit can remain at 1.77 for the transition to 422V+ fuel with consideration of any transition core penalties necessary as desc ribed in Section 4.6. A comparison of the )

Z (F N Q without uncertainty versus cycle length for each of the core models used is provided in Figure 3-12. The )

Z (F N Q (total peaking factor) limit can remain at 2.50 for the transition to 422V+ fuel. Beyond the power distribution impacts already mentioned, other changes to the core power distributions and peaking factors are the result of the normal cycle-to-cycle variations in core loading patterns. The normal methods of feed enrichment variation and fresh burnable absorbers will be employed to control peaking factors. Compliance with the peaking factor Technical Specifications can be assured using these methods. 3.6 NUCLEAR DESIGN EVALUATION CONCLUSIONS The nuclear core design analysis completed for the transition to 422V+ has confirmed that peaking factors and key safety parameter limits can be maintained identical or similar to the existing accident analysis assumptions for 400V+ (see Table 3-1). The revised key safety parameters calcula ted for the transition to 422V+ have been used in applicable analyses evaluated in other sections of this report. All analyses were able to meet the required acceptance criteria with these revised key safety parameters for 422V+ fuel.

3-4

3.7 REFERENCES

1. WCAP-9272-P-A/WCAP-9273-NP-A, "Westinghouse Reload Safety Evaluation Methodology," Davidson, S. L. (Ed.), et al., July 1985. 2. WCAP-11596-P-A, "Qualification of the PHOENIX-P/ANC Nuclear Design System for Pressurized Water Reactor Cores," Nguyen, T. Q., et al., June 1988. 3. WCAP-10965-P-A, "ANC: A Westinghouse Advanced Nodal Computer Code," Liu, Y. S., et al., September 1986. 4 Addendum 1 to WCAP-16517-NP, Revision 0, "Prairie Island Units 1 and 2 Spent Fuel Pool Criticality Analysis," Bishop, T. C., February 2008. 5. Letter from J. W. Baker (Westinghouse) to H. Hoelscher (Nuclear Management Company), "Prairie Island Units 1 & 2 Criticality Analysis," NF-NMC-08-29, February 22, 2008. 6. WCAP-12610-P-A, "VANTAGE + Fuel Assembly Reference Core Report," Davidson, S. L. (Ed.), et al., April 1995. 7. Letter from J. D. Peralta (NRC) to B.F. Maurer (Westinghouse), "Approval for Increase in Licensing Burnup Limit to 62,000 MWD/MTU (TAC No. MD1486)," NRC/RCPL-06-055, May 2006. 8. WCAP-10125-P-A (Proprietary), "Extended Burnup Evaluation of Westinghouse Fuel," Davidson, S. L. (Ed.), et al., December 1985.

3-5 Table 3-1 Key Safety Parameters Safety Parameter All 400V+ All 422V+ Reactor Core Power (MWt) 1,683 1,683 Core Average Coolant Temperature Hot Full-Power (HFP) (°F) 563.3 563.3 Coolant System Pressure (psia) 2,250 2,250 Most Positive Isothermal Temperature Coefficient (ITC) (pcm/°F) 0 0 Most Positive Moderator Density Coefficient (MDC) (K/g/cm 3) 0.43 0.43 Doppler Temperature Coefficient (pcm/degF) -2.90 to -0.91 -2.90 to -0.91 Doppler Only Power Coefficient (pcm/%Power, Q = power in %) Least Negative, HFP to Hot Zero-Power (HZP) -9.50 + 0.035*Q -12.0 + 0.045*Q Most Negative, HFP to HZP -24.0 + 0.100*Q -24.0 + 0.100*Q Beta-Effective 0.0043 to 0.0072 0.0043 to 0.0072 Normal Operation F NH 1.77 1.77 Shutdown Margin (%) 1.70 - Units 1 &2 1.70 - Unit 1 1.90 - Unit 2 (1) Normal Operation F Q(Z) 2.5 2.5 Note: 1. Unit 2 with original steam generators. A value of 1.70 % for Unit 2 may be supported when the replacement steam generators are in place.

3-6 1 l 2 l 3 l 4 l 5 l 6 l 7 l 1 - Feed A Feed 0 26225 Once B (5,4) 270° 25909 51160 Once B (3,1) 0° 30145 52445 Once B (4,4) 270° 29979 52295 Once B (3,2) 270° 30095 53259 Feed C Feed 0 24150 Once B (2,3) 0° 30124 411942 - Once B (5,4) 0° 25909 51160 Feed B Feed 0 26888 Once B (6,2) 180° 24860 49578 Once B (5,3) 180° 29214 53042 Feed C Feed 0 27350 Feed C Feed 0 23811 Thrice B (6,4) 270° 45569 535353 - Once B (3,1) 90° 30145 52445 Once B (2,6) 180° 24836 49609 Once B (4,5) 180° 25910 51246 Feed B Feed 0 27623 Feed B Feed 0 26177 Once B (2,5) 180° 30650 45879 4 - Once B (4,4) 0° 29979 52295 Once B (3,5) 180° 29217 Feed B Feed 0 27642 Once B (5,1) 270° 30714 Feed C Feed 0 22901 Twice B (6,3) 180° 45653 53811 5 - Once B (3,2) 0° 30095 53259 Feed C Feed 0 Feed B Feed 0 261 82 Feed C Feed 0 22898 Twice B (5,5) 180° 42496 52607 6 - Feed C Feed 0 24150 Feed C Feed 0 Once B (5,2) 180° 30650 45883 Twice B (3,6) 180° 45642 53800 7 - Once B (2,3) 90° 30124 41194 Thrice B (4,6) 90° 45568 53534 Region (Shuffle) Rotation BOC Burnup EOC Burnup Figure 3-1. First Transition Cycle Loading Pattern with BOC and EOC Assembly Burnups 3-7 1234567lllllllFeed AOnce BOnce BOnce BOnce BFeed COnce B1.2191.1691.0581.0371.0591.1800.4531.3411.1461.0010.9961.0361.2030.488 1.3311.0950.9600.9761.0171.2630.555Once BFeed BOnce BOnce BFeed CFeed CThrice B1.1691.3141.1791.0931.3081.2440.3311.1461.3621.1171.0701.3741.1960.3531.0951.3521.0671.0441.3981.1890.400Once BOnce BOnce BFeed BFeed BOnce B1.0581.1831.1841.3171.2610.659 1.0011.1201.1451.3971.3160.6770.9601.0681.0971.3971.3360.710Once BOnce BFeed BOnce BFeed CTwice B1.0371.0961.3191.0981.2240.345 0.9961.0721.3981.0531.1530.361 0.9761.0451.3971.0081.1250.397Once BFeed CFeed BFeed CThrice B1.0591.3101.2621.2240.456 1.0361.3751.3171.1530.449 1.0171.3981.3361.1250.470Feed CFeed COnce BTwice B1.1801.2450.6600.3461.2031.1970.6780.3611.2631.1890.7100.397Once BThrice B0.4530.331 0.4880.353 0.5550.400EOC PowerMOC PowerBOC PowerRegion 7 5 3 -

2 Figure 3-2. First Transition Cycle BOC, MOC, and EOC Assembly Power Distributions 3-8 1 l 2 l 3 l 4 l 5 l 6 l 7 l 1 - Once B (1,1) 0° 26225 43866 Once B (3,4) 0° 27623 47055 Feed A Feed 0 24864 Once B (2,2) 90° 26888 48511 Once B (1,6) 0° 24150 Feed B Feed 0 23939 Twice B (1,7) 180° 41194 502462 - Once B (3,4) 90° 27623 47055 Once B (2,5) 90° 27350 47712 Once B (5,4) 0° 22898 45404 Once B (3,5) 90° 26176 48509 Feed B Feed 0 26420 Feed B Feed 0 22117 Twice B1 (4,3) 0° 48503 553583 - Feed A Feed 0 24864 Once B (4,5) 0° 22901 4540 5 Once B (5,2) 90° 27369 49349 Once B (6,2) 180° 23817 47447 Feed B Feed 0 25483 Twice B (3,6) 180° 45879 57803 4 - Once B (2,2) 180° 26888 48511 Once B (5,3) 270° 26192 48509 Once B (2,6) 180° 23811 47441 Feed A Feed 0 26209 Feed B Feed 0 22553 Twice B (3,2) 0° 49609 57143 5 - Once B (6,1) 0° 24150 47458 Feed B Feed 0 26416 Feed B Feed 0 25483 Feed B Feed 0 22552 Once B (4,3) 0° 27672 38959 6 - Feed B Feed 0 23939 Feed B Feed 0 22117 Twice B (6,3) 180° 45883 57806 Twice B (2,3) 0° 49578 57115 7 - Twice B (7,1) 180° 41194 50246 Twice B1 (3,4) 0° 48509 55365 Region (Shuffle) Rotation BOC Burnup EOC Burnup Figure 3-3. Second Transition Cycle Loading Pattern with BOC and EOC Assembly Burnups 3-9 1234567lllllllOnce BOnce BFeed AOnce BOnce BFeed BTwice B1.0181.1141.2821.1821.2261.2090.3700.8980.9921.2831.1171.2121.2450.4180.8760.9621.2971.0731.1831.2710.471Once BOnce BOnce BOnce BFeed BFeed BTwice B1.1141.1801.3091.2321.2591.1050.2720.9921.0391.1601.1551.3801.1540.3190.9620.9961.0981.1051.4121.1860.365Feed AOnce BOnce BOnce BFeed BTwice B1.2821.3081.2691.3021.2520.4971.2831.1601.1281.2321.3310.5531.2971.0981.0551.1711.3530.599Once BOnce BOnce BFeed AFeed BTwice B1.1821.2321.3021.2611.1400.3101.1171.1541.2321.3731.1800.3511.0731.1041.1711.3811.1870.389Once BFeed BFeed BFeed BOnce B1.2261.2591.2521.1400.5461.2121.3791.3311.1800.5901.1831.4121.3531.1870.615Feed BFeed BFeed BTwice B1.2091.1050.4970.3101.2451.1540.5530.3521.2711.1860.5990.390Twice BTwice B0.3700.2720.4180.3190.4710.365 4 EOC Power 6 -

7 -RegionBOC PowerMOC Power 1 -

2 -

3 - Figure 3-4. Second Transition Cycle BOC, MOC, and EOC Assembly Power Distributions 3-10 1 l 2 l 3 l 4 l 5 l 6 l 7 l 1 - Twice B (1,1) 0° 43866 58042 Once B (2,5) 0° 26420 45384 Once B (5,2) 270° 26416 46315 Once B (3,1) 90° 24864 45507 Once B (1,6) 0° 23939 45800 Feed B Feed 0 21953 Twice B (5,1) 90° 47459 54718 2 - Once B (2,5) 90° 26420 45384 Feed A Feed 0 21868 Once B (4,5) 0° 22553 43699 Once B (2,6) 180° 22117 43936 Feed A Feed 0 23389 Feed B Feed 0 20313 Twice B (2,4) 90° 48509 54397 3 - Once B (5,2) 0° 26416 46315 Once B (5,4) 0° 22552 43701 Once B (4,4) 0° 26209 46492 Once B (5,3) 90° 25483 46698 Feed B Feed 0 23149 Twice B (3,2) 180° 45405 55741 4 - Once B (3,1) 180° 24864 45507 Once B (6,2) 180° 22117 43940 Once B (3,5) 270° 25483 46699 Feed A Feed 0 22822 Feed B Feed 0 20183 Twice B (4,3) 0° 47441 53977 5 - Once B (6,1) 0° 23939 45800 Feed A Feed 0 23390 Feed B Feed 0 23150 Feed B Feed 0 20184 Twice B (5,5) 180° 38959 47770 6 - Feed B Feed 0 21953 Feed B Feed 0 20313 Twice B (2,3) 180° 45404 55740 Twice B (3,4) 0° 47447 53983 7 - Twice B (5,1) 180° 47459 54718 Twice B (4,2) 270° 48509 54397 Region (Shuffle) Rotation BOC Burnup EOC Burnup Figure 3-5. Third Transition Cycle, Pattern A, Loading Pattern with BOC and EOC Assembly Burnups 3-11 1234567lllllllTwice BOnce BOnce BOnce BOnce BFeed BTwice B0.7931.1131.2301.2621.2751.2210.3730.7931.0511.0941.1421.2331.2440.4130.8171.0431.0451.0831.1921.2610.450Once BFeed A Once BOnce BFeed A Feed BTwice B1.1131.2221.3151.3291.2421.1190.2991.0511.2291.1721.2191.3421.1560.3361.0431.2431.1161.1591.3431.1750.367Once BOnce BOnce BOnce BFeed BTwice B1.2301.3151.2351.2421.2320.5381.0941.1721.1211.1901.3210.5931.0451.1161.0701.1551.3570.630Once BOnce BOnce BFeed A Feed BTwice B1.2621.3291.2421.1951.0920.3291.1421.2201.1901.3081.1510.376 1.0831.1591.1551.3301.1820.411Once BFeed A Feed BFeed BTwice B1.2751.2421.2321.0920.454 1.2331.3421.3211.1510.502 1.1921.3431.3571.1820.537Feed BFeed BTwice BTwice B1.2211.1190.5380.3291.2441.1560.5930.3761.2611.1750.6300.411Twice BTwice B0.3730.299 0.4130.3360.4500.367 4 EOC Power 6 -

7 -BOC PowerMOC PowerRegion 1 3 - Figure 3-6. Third Transition Cycle, Pattern A, BOC, MOC, and EOC Assembly Power Distributions 3-12 1 l 2 l 3 l 4 l 5 l 6 l 7 l 1 - Feed A Feed 0 27466 Once B (4,4) 0° 26209 50830 Feed A Feed 0 27695 Once B (3,1) 90° 24864 49457 Once B (1,6) 0° 23939 49477 Feed C Feed 0 25685 Twice B (5,1) 90° 47459 56066 2 - Once B (4,4) 90° 26209 50830 Feed A Feed 0 27930 Once B (5,3) 90° 25483 51027 Once B (2,6) 180° 22117 48258 Feed B Feed 0 29351 Feed C Feed 0 24334 Twice B (3,2) 180° 45405 52939 3 - Feed A Feed 0 27695 Once B (3,5) 270° 25483 51023 Feed B Feed 0 30290 Once B (4,5) 0° 22553 49423 Feed C Feed 0 28345 Once B (5,2) 0° 26416 41547 4 - Once B (3,1) 180° 24864 49457 Once B (6,2) 180° 22117 48255 Once B (5,4) 0° 22552 49417 Feed B Feed 0 29463 Feed C Feed 0 24310 Twice B (4,3) 0° 47441 55540 5 - Once B (6,1) 0° 23939 49477 Feed B Feed 0 29345 Feed C Feed 0 28328 Feed C Feed 0 24254 Twice B (5,5) 270° 38959 49317 6 - Feed C Feed 0 25685 Feed C Feed 0 24328 Once B (2,5) 0° 26420 41537 Twice B (3,4) 0° 47447 55514 7 - Twice B (5,1) 180° 47459 56066 Twice B (2,3) 180° 45404 52936 Region (Shuffle) Rotation BOC Burnup EOC Burnup Figure 3-7. Third Transition Cycle, Pattern B, Loading Pattern with BOC and EOC Assembly Burnups 3-13 1234567lllllllFeed AOnce BFeed AOnce BOnce BFeed CTwice B1.0110.9971.1291.1601.2361.2080.3731.2631.1281.2671.1051.1421.1340.3741.2731.1011.2461.0611.1141.1770.438Once BFeed AOnce BOnce AFeed BFeed CTwice B0.9971.0861.1441.2471.3161.1440.3281.1281.2841.1651.1841.3161.0800.3301.1011.2691.1101.1261.3101.1140.384Feed AOnce BFeed BOnce BFeed COnce B1.1291.1431.3181.2791.2720.6751.2671.1641.3811.2171.2680.6661.2461.1101.3241.1551.2770.713Once BOnce BOnce BFeed BFeed CTwice B1.1601.2471.2781.3061.1190.3481.1051.1841.2171.3281.0870.3581.0611.1261.1551.3081.1060.403Once BFeed BFeed CFeed CTwice B1.2361.3151.2701.1160.4521.1421.3151.2671.0840.4571.1141.3101.2771.1040.501Feed CFeed COnce BTwice B1.2081.1430.6740.3471.1341.0790.6660.3571.1771.1140.7120.402Twice BTwice B0.3730.3280.3740.3300.4380.384 4 EOC Power 6 -Region 7 -BOC PowerMOC Power 1 -

2 -

3 - Figure 3-8. Third Transition Cycle, Pattern B, BOC, MOC, and EOC Assembly Power Distributions 3-14 0 500 1000 1500 2000 25000100200300400500600700Exposure (EFPD)Critical Concentration (ppm)Cycle 26Cycle 27Cycle 28ACycle 28B Figure 3-9. Critical Boron Concentration Comparison Versus Exposure 3-15 4-2 0 2 4 6 8 100100200300400500600700Exposure (EFPD)Axial Offset (%)Cycle 26Cycle 27Cycle 28ACycle 28B Figure 3-10. Axial Offset Comparison Versus Exposure 3-16 1.41.51.61.70100200300400500600700Exposure (EFPD)Radial Peaking FactorCycle 26Cycle 27Cycle 28ACycle 28B Figure 3-11. Radial Peaking Factor (F NH) Comparison Versus Exposure 3-17 1.61.71.81.9 22.10100200300400500600700Exposure (EFPD)Total Peaking FactorCycle 26Cycle 27Cycle 28ACycle 28B Figure 3-12. Total Peaking Factor ()Z (F N Q) Comparison Versus Exposure

4-1 4 THERMAL AND HYDRAULIC DESIGN

4.1 INTRODUCTION

AND

SUMMARY

This section describes the calculational methods used for the thermal-hydraulic analysis, the departure from nucleate boiling (DNB) performance, and the hydraulic compatibility during the transition from 14x14 VANTAGE+ (V+) with a 0.400-inch diameter rod (400V+, also referred to as optimized fuel assembly - OFA) to 14x14 V+ with a 0.422-inch diameter rod (422V+). Based on design differences and hydraulic testing of the fuel assemblies, it has been shown that the 400V+ fuel design and 422V+ fuel assembly designs are hydraulically compatible. Table 4-1 illustrates a comparison between the previous thermal-hydraulic design parameters and the new thermal-hydraulic design parameters for Prairie Island Units 1&2 that were used in this analysis. A discussion of the thermal-hydraulic parameters is provided in Section 4.8. The thermal-hydraulic design criteria and methods are the same as those approved for Prairie Island (Reference 1) and are described in the following sections. All of the current Updated Safety Analysis Report (USAR) thermal-hydraulic design criteria are satisfied.

4.2 METHODOLOGY

The thermal-hydraulic analysis of the 14x14 422V+ fuel in Prairie Island Units 1&2 is based on the Revised Thermal Design Procedure (RTDP) (Reference 2) and the WRB-1 DNB correlation (Reference 3) as described in the Prairie Island Units 1&2 USAR (Reference 1). The DNB analysis of the core containing both 14x14 422V+ and 400V+ fuel assemblies has been shown to be valid with the WRB-1 DNB correlation (References 3 and 4), RTDP (Reference 2), and the VIPRE-W Modeling (Reference 4). The W-3 correlation and Standard Thermal Design Procedure (STDP) are used when any of the conditions are outside the range of the WRB-1 correlation (such as pressure, local mass velocity, local quality, heated length, grid spacing, equivalent hydraulic diameter, equivalent heated hydraulic diameter, and distance from last grid to critical heat flux site) and RTDP (such as the statisti cal variance is exceeded on power, T IN, pressure, flow, bypass, F NH , F EH,1 , and F E Q). The WRB-1 DNB correlation is based entirely on rod bundle data and takes credit for the significant improvements in the accuracy of the critical heat flux predictions over previous DNB correlations. The Nuclear Regulatory Commission (NRC) approval that a 95/95 correlation limit DNB ratio (DNBR) of 1.17 is appropriate for the 14x14 400V+ and 422V+ fuel assemblies has been documented (Reference 3). With RTDP methodology, uncertainties in plant operating parameters, nuclear and thermal parameters, fuel fabrication parameters, computer codes, and DNB correlation predictions are combined statistically to obtain the overall DNB uncertainty factor. This factor is used to define the design limit DNBR that satisfies the DNB design criterion (that is, a plant-specific design limit accounts for the RTDP uncertainties above the correlation DNBR limit). The criterion is that the probability that DNB will not occur on the most limiting fuel rod is at least 95 percent at 95-percent confidence level for any Condition I or II event (such as normal operation or anticipated operational o ccurrences). Since the parameter uncertainties are considered in determining the RTDP design limit DNBR values, the plant safety analyses are performed using input parameters at their nominal values. For cases where conditions fall outside the bounds of the RTDP methodology (such as the statistical variance is exceeded on power, 4-2 T IN, pressure, flow, bypass, F NH , F EH,1 , and F E Q or the W-3 correlation is used), STDP is used and the associated analyses are performed using input parameters with their uncertainties included. The increase in DNB margin is realized when nominal values of the peaking factors/hot channel engineering factors and operating parameters are used in the RTDP DNB safety analyses. Table 4-2 provides a listing and description of these hot channel factors. Table 4-3 provides a listing of the uncertainties used in the safety analysis. These bounding values were used during the RTDP design limit DNBR calculation. The design limit DNBR values for the 422V+ fuel are 1.23/1.23 for typical/thimble cells. The design limit DNBR values for 400V+ are 1.22/1.22 for typical/thimble cells. For use in the DNB safety analyses, the design limit DNBR is conservatively increased to provide DNB margin to offset the effect of rod bow, transition core, DNBR penalty due to instrument bias, any other DNB penalties that may occur, and to provide flexibility in design and operation of the plant. This increase in the design limit to account for various penalties and operational issues is the plant-specific margin retained between the design limit and the safety analysis limit. After accounting for the plant-specific margin, the safety analysis limit for the 422V+ fuel is 1.415/1.415 (typical/thimble) and 1.340/1.340 (typical/thimble) for the 400V+ fuel. These safety analysis limits are employed in the DNB analyses. With the safety analysis limit set, the core limit lines, axial offset limit lines, and dropped rod limit lines are generated. In generating the various limits, the maximum F NH is determined that yields acceptable results based upon the safety analysis limits. Based on generating these limits, the maximum F NH limit that can be supported is 1.77 (including uncertainties) for the 422V+ fuel. A lower F NH limit is required for 400V+ fuel as shown in Table 4-1 to offset the transition core DNBR penalty. The measurement uncertainty of 4 percent (Reference 5) is accounted for in the above F NH limits . Table 4-4 summarizes the available DNBR margin for Prairie Island for the transition core. It should be noted that the DNBR margin summaries are cycle dependent and may vary from that documented here in future reload designs.

4.3 HYDRAULIC

COMPATIBILITY The 14x14 422V+ and 400V+ fuel assembly designs have been shown to be hydraulically compatible (References 6 through 8), based on a comparison of the component loss coefficients. The difference in loss coefficients causes fuel assembly crossflow. Refer to Section 2.2 for more discussion of crossflow. The axial grid locations, grid heights, and fuel assembly pitch and envelope for the 14x14 422V+ have remained consistent with the 14x14 400V+ design, minimizing assembly-to-assembly crossflow. By maintaining grid-to-grid overlap between the 400V+ design and the 422V+ design, excessive crossflow between assemblies is prevented. The small difference in loss coefficients between the two designs and the respective grid locations of the two designs has been analyzed to demonstrate that no crossflow-induced vibration will result in a condition in which fretting or whirling would be induced. The fuel assembly crossflow that exists for the transition core is well within the bounding Westinghouse experience basis of transition core analysis (that is, transition cores with intermediate flow mixing (IFM) vane grids will experience the maximum crossflow situation).

4-3 A second area of hydraulic compatibility associated with higher resistance fuel assemblies (the 400V+ design) is the associated impact on lift forces. When a fuel assembly with a different hydraulic resistance is loaded into a core, it changes the flow distribution in the surrounding assemblies. In particular, if this fuel assembly has a higher value of fuel assembly loss coefficient, the surrounding assemblies (such as lower resistance fuel assemblies; the 422V+ assemblies) would see a higher average flow through them than they would in a full core situation. Therefore, the lift force on these surrounding assemblies can be expected to increase. The larger the number is of high resistance fuel assemblies loaded in the core, the greater the lift force is on the lower resistance assemblies. The transition core effect and lift force calculations show acceptable margin for the top nozzle spring design of the 422V+ fuel such that fuel assembly lift off will not occur.

4.4 EFFECTS

OF FUEL ROD BOW ON DNBR The concern with regards to fuel rod bow phenomenon is the potential effects on bundle power distribution and on the margin of fuel rods to DNB. Therefore, the phenomenon of fuel rod bowing must be accounted for in the DNBR safety analysis of Condition I and Condition II events. Fuel rod bow is the phenomenon of fuel rods bowing between mid-grids. The effect of the rod bow is to impact the channel spacing between adjacent fuel rods. With a reduced channel spacing, the potential of DNB occurring increases. To determine the impact of rod bow on DNB, Westinghouse conducted tests to determine the impact of rod bow on DNB performance. These tests and subsequent analyses were documented in Reference 9. Currently, the maximum rod bow penalties applicable to Prairie Island at an assembly average burnup of 24,000 MWD/MTU (References 10 and 11) are 2.9 percent DNBR for the 422V+ fuel and 3.0 percent for the 400V+ fuel. For burnups greate r than 24,000 MWD/MTU, credit is taken for the effect of F NH burndown, due to the decrease in fissionable isotopes and the buildup of fission product inventory (Reference 10 and 11). Therefore, no additional rod bow penalty is required at burnups greater than 24,000 MWD/MTU. For both the 400V+ and 422V+ fuel types, the rod bow penalty will be offset with DNB margin retained between the safety analysis and design DNBR limits (refer to Table 4-4).

4.5 FUEL TEMPERATURE/PRESSURE ANALYSIS Fuel temperatures and associated rod internal pre ssures have been generated (Reference 12) for the 400V+ and 422V+ fuel types. The characteristics of the gadolinium fuel are such that the gadolinium rods would exhibit higher fuel temperatures due to an inherent lower thermal conductivity of the gadolinium-bearing fuel pellet. In addition, increasing gadolinium enrichment results in a corresponding decrease in the fuel melting temperature. The performance criteria employed by Westinghouse for gadolinium rods is to ensure that these rods are less limiting than the non-gadolinium rods, throughout life, in terms of fuel temperatures, rod internal pressures, and core stored energy. This is achieved by reducing the U235 enrichment in the gadolinium rods so that the gadolinium rods maintain a sufficiently lower power throughout life. Therefore, the fuel performance parameters for the 422V+ fuel without gadolinium bound those for the 422V+ fuel with gadolinium, and likewise for 400V+. Higher fuel rod average and surface temperatures are conservative for the loss-of-coolant-accident (LOCA) and transient analyses. In addition, minimum fuel average and fuel surface temperatures are required by transient analysis. Therefore, the 422V+ and 400V+ non-gadolinium fuel minimum temperatures are generated, which with the maximum fuel temperatures, form a consistent basis for transient analysis.

4-4 Fuel centerline temperatures were also generated for the 400V+ and 422V+ fuel types. These have been provided to the Core Design group for future verification, during reload design validation, that fuel melt will not occur. The maximum kW/ft limits for fuel melt are 22.65 for the 400V+ fuel and 22.76 kW/ft for the 422V+ fuel. In addition to the fuel temperatures and pressures, the core stored energy for the 422V+ fuel has been determined for use in containment analysis (refer to Sections 5.3 and 5.4). Core stored energy is defined as the amount of energy in the fuel rods in the core above the local coolant temperature. The local core stored energy is normalized to the local linear power level. The units for the core stored energy are in full-power seconds (FPS). The value of the core stored energy for the 422V+ fuel product is 4.60 FPS. A bounding value of 5.77 FPS continues to be used for the 400V+ fuel.

4.6 TRANSITION

CORE EFFECT Redistribution of flow in pressurized water reactor cores is a well documented and modeled phenomenon that occurs generally because of differences in loss coefficients and rod diameters. In a mixed core, with assemblies having different hydraulic resistance, the local hydraulic resistance differences are also a mechanism for flow redistribution. This redistribution results in the fluid velocity vector having a lateral component as well as the dominant axial component. The lateral component is commonly referred to as crossflow. The crossflow induced by local hydraulic resistance differences will typically impact the mechanical design of the fuel assemblies, as well as the safety analyses of the core. Refer to Section 2.2 for more discussion of crossflow. The mechanical design of the fuel assemblies in the core could be affected in two ways: 1) excitation of peripheral rods in the fuel assemblies such that wear mechanisms of fretting or whirling could exist, and 2) introduction of higher resistance assemblies will influence the lift forces on the remaining assemblies. The hydraulic compatibility of the 422V+ and 400V+ fuel assemblies has been addressed in Sections 2.2 and 4.3 and found to be acceptable. In the safety analysis, crossflow affects both LOCA and DNB. The primary consideration for the LOCA analysis is the reduction of the normalized mass velocity as compared to a full core of that assembly type. DNB is affected because the flow redistribution affects both mass velocity and enthalpy distributions. With the current DNB correlations, WRB-1 and W-3, the flow redistribution affects the prediction of minimum DNBR. As such, the design procedure is based on the principle that once the transition core DNBR penalty is determined, all further plant-specific analysis may proceed as if it were a full core of 422V+. Transition cores are analyzed as if they were full cores of one assembly type (full 422V+ or full 400V+) by applying the applicable transition core penalty. Transition core penalties are calculated according to the methodology documented in Reference 5. This methodology is used to calculate the maximum 400V+ to 422V+ transition core penalties. There is a maximum 9.0-percent transition core DNBR penalty for the 400V+ fuel which will be offset by a 6.0-percent F NH reduction in burned 400V+ fuel based on a conservative 1.5-percent DNBR: 1-percent FH sensitivity. There is a maximum 2.0-percent transition core DNBR penalty on the 422V+ which will be offset by available DNBR margin.

4-5 4.7 BYPASS FLOW Two different bypass flow rates are used in the thermal-hydraulic design analysis-thermal design bypass flow (TDBF) and best-estimate bypass flow (BEBF). These two bypass flows are used in non-statistical and statistical analyses, respectively. The TDBF is the conservatively high core bypass flow used in calculations where the results are adversely affected by low core flow. Specifically, TDBF is used with the vessel thermal design flow (TDF) in power capability analyses that use standard (non-statistical) methods. The TDBF is also used with the vessel best-estimate flow (BEF) to calculate core and fuel assembly pressure drops. The BEBF is the flow that would be expected using nominal values for dimensions and operating parameters that affect bypass flow without applying any uncertainty factors. The BEBF is used in conjunction with the vessel minimum measured flow (MMF) for power capability analyses that use RTDP (statistical methodology). It is also used to calculate fuel assembly lift forces. The maximum permissible TDBF is 6.0 percent and the maximum permissible BEBF is 4.5 percent for both 400V+ and 422V+ fuel.

4.8 THERMAL-HYDRAULIC DESIGN PARAMETERS Table 4-1 lists numerous thermal-hydraulic parameters for the current design basis at 1,677 MWt for both the 400V+ and 422V+ fuel types. Some of the parameters listed in Table 4-1 are used in the analysis basis as VIPRE-W input parameters while others are simply provided since they are listed in the USAR. This section will identify those parameters that are used as input parameters to the VIPRE-W model and will also identify the limiting direction of each parameter. The following parameters from Table 4-1 are used in the VIPRE-W model: Reactor core heat output (MWt) F NH, nuclear enthalpy rise hot-channel factor Core pressure for RTDP analyses (psia) Pressurizer/core pressure (psia) Heat generated in fuel (%) Thermal design flow for non-RTDP analyses (gpm) Average heat flux (Btu/hr-ft

2) Minimum measured flow for RTDP analyses (gpm) Nominal vessel/core inlet temperature (°F) Bypass flow (%) The limiting direction for these parameters is as shown in Table 4-5.

4.9 CONCLUSION

The thermal-hydraulic evaluation of the fuel upgrade for Prairie Island has shown that 400V+ and 422V+ fuel assemblies are hydraulically compatible. More than sufficient DNBR margin in the safety limit DNBR exists to cover any rod bow and transition core penalties. All current thermal-hydraulic design criteria are satisfied.

4.10 REFERENCES 1. Prairie Island Updated Safety Analysis Report, Revision 29, May 2007. 2. WCAP-11397-P-A, "Revised Thermal Design Procedure," Friedland, A. J. and Ray, S., April 1989.

4-6 3. Letter from C. O. Thomas (NRC) to E. P. Rahe (Westinghouse), "SER on the Applicability of WRB-1 to Westinghouse 14x14 and 15x15 OFA," June 29, 1984. 4. WCAP-14565-P-A (Proprietary), "VIPRE-01 Modeling and Qualification for Pressurized Water Reactor Non-LOCA Thermal-Hydraulic Safety Analysis," Sung, Y. X., et al., October 1999. 5. WCAP-11837-P-A, "Extension of Methodology for Calculating Transition Core DNBR Penalties," Schueren, P. and McAtee, K. R., January 1990. 6. Letter from B. W. Gergos (Westinghouse) to M. Baumann (WEPCO), "14x14 Heavy Fuel Assembly Loss Coefficients," 99WE-G-0025, May 19, 1999. 7. Letter from B. W. Gergos (Westinghouse) to M. Baumann (WEPCO), "FACTS Vibration Test Report - 14 Heavy Fuel Assembly," 99WE-G-0026, May 19, 1999. 8. Letter from B. W. Gergos (Westinghouse) to M. Baumann (WEPCO), "FACTS Hydraulic Test Report - Westinghouse 14x14 Heavy Fuel Assembly," 99WE-G-0023, May 10, 1999. 9. WCAP-8691, Revision 1, "Fuel Rod Bow Evaluation," Skaritka, J., July 1979. 10. Letter from C. Berlinger, (NRC) to E. P Rahe, Jr. (Westinghouse), "Request for Reduction in Fuel Assembly Burnup Limit for Calculation of Maximum Rod Bow Penalty," June 18, 1986. 11. Letter from E. P. Rahe, Jr. (Westinghouse) to J. R. Miller, (NRC), "Partial Response to Request Number 1 for Additional Information on WCAP-8691, Revision 1," NS-EPR-2515, October 9, 1981; and Letter from E. P. Rahe, Jr. (Westinghouse) to J. R. Miller, (NRC), "Remaining Response to Request Number 1 for Additional Information on WCAP-8691, Revision 1," NS-EPR-2572, March 16, 1982. 12. WCAP-15063-P-A with errata, "Westinghouse Improved Performance Analysis and Design Model (PAD 4.0)," Foster, J. P., Sidener, S., July 2000. 13. WCAP-16206-P-A., "Safety Analysis Transition Engineering Report for the Prairie Island Nuclear Power Plant - Volume 1 Engineering Analysis," Brown, L. (Ed.), February 2004.

4-7 Table 4-1 Prairie Island Thermal-Hydraulic Design Parameters Comparison Thermal-Hydraulic Design Parameters Design Value Reactor Core Heat Output, MWt (1) 1,677 Reactor Core Heat Output, 10 6, Btu/hr 5,722 Heat Generated in Fuel, % 97.4 Core Pressure, Nominal - RTDP, psia 2,265 Pressurizer Pressure, Nominal, psia 2,250 Radial Power Distribution(2)(4) 1.77[1+0.3(1-P)]

where P =

Power Thermal Rated Power Thermal Hot Full-Power Nominal Coolant Conditions Vessel TDF Rate (including bypass) 10 6 lbm/hr 68.85 GPM 178,000 Core Flow Rate (excluding bypass, (3) based on TDF) 10 6 lb m/hr 64.71 GPM 167,320 Core Flow Area, ft 2 29.2 (full-core 400V+) 27.1 (full-core 422V+)Core Inlet Mass Velocity, (Based on TDF) 10 6 lb m/hr-ft 2 2.22 2.39 Thermal-Hydraulic Design Parameters (Based on TDF, 6.0% Bypass, 1,677 MWt) Design Value Nominal Vessel/Core Inlet Temperature, °F 527.4 Vessel Average Temperature, °F 560.0 Core Average Temperature, °F 563.3 Vessel Outlet Temperature, °F 592.6 Core Outlet Temperature, °F 596.4 Average Temperature Rise in Vessel, °F 65.2 Average Temperature Rise in Core, °F 69.0 Heat Transfer 400V+ Design Value 422V+ Design Value Active Heat Transfer Surface Area, ft 2 27,161 28,507 Average Heat Flux, BTU/hr-ft 2 210,670 200,723 4-8 Table 4-1 Prairie Island Thermal-Hydraulic Design Parameters Comparison (cont.) Average Linear Power, kw/ft 6.47 6.50 Peak Linear Power for Normal Operation, (5) kW/ft 16.18 16.25 Pressure Drop Across Core, psi (6) Full core of 400V+ 26.1 ----- Full core of 422V+ ----- 25.2 Notes: 1. A power level of 1,677 MWt was been used for all RTDP thermal-hydraulic design analyses. For analyses explicitly modeling parameter uncertainties, a power level of 1,683 MWt was used. 2. Includes 4% measurement uncertainty. 3. Based on design bypass flow of 6% for current design value. 4. Based on a bounding transition core penalty calculation, a 6% reduction in FdH for the 400V+ fuel is required to offset the transition core DNBR penalty. There is sufficient retained DNBR margin to offset the transition core DNBR penalty on 422V+ fuel. See Table 4-4 for more information. 5. Based on maximum F Q of 2.50. 6. Based on BEF rate of 201,400 gpm (OFA)/202,000 gpm (422V+) and 4% bypass flow.

4-9 Table 4-2 Hot Channel Factors (1) FH = F NH x F EH Where: F NH Nuclear enthalpy rise hot channel factor - The ratio of the relative power of the hot rod, which is one of the rods in the hot channel, to the average rod power. The normal operation value of this is given in the plant Technical Specifications or a Core Operating Limits Report.

F EH Engineering enthalpy rise hot channel factor - The nominal enthalpy rise in an isolated hot channel can be calculated by dividing the nominal power into this channel by the core average inlet flow per channel. The engineering enthalpy rise hot channel factor accounts for the effects of flow conditions and fabrication tolerances. It can be written symbolically as:

F EH = f (F EH,1 , F EH,2 , F EH inlet maldist , F EH redist , F EH mixing) Where: F EH,1: Accounts for rod-to-rod variations in fuel enrichment and weight.

F EH,2: Accounts for variations in fuel rod outer diameter, rod pitch, and bowing.

F EH inlet maldist: Accounts for the nonuniform flow distribution at the core inlet.

F EH redist: Accounts for flow redistribution between adjacent channels due to the different thermal-hydraulic conditions between channels.

F EH mixing: Accounts for thermal diffusion energy exchange between adjacent channels caused by both natural turbulence and forced turbulence due to the mixing vane grids.

Note: 1. The value of these factors and the way in which they are combined depends upon the design methodology used, that is, STDP or RTDP. Note that no actual combined effect value is calculated for F EH, 1. These factors are accounted for by using the VIPRE-W code.

4-10 Table 4-3 RTDP Uncertainties Parameter Original RTDP Uncertainty (WCAP-16206-P) Uncertainty Used for 422V + Analysis Power +/- 2.0% +/- 0.5% power Reactor Coolant System Flow +/- 3.0% +/- 3.0 % flow Pressure +/- 45.0 psi +/- 60.0 psi Inlet Temperature +/- 4.0°F - 0.5°F (bias) +/- 4.0°F

- 0.5°F (bias)

4-11 Table 4-4 RTDP DNBR Margin Summary (1) 400V+ 422V+ DNB Correlation WRB-1 DNBR Correlation Limit 1.17 DNBR Design Limit (TYP)

(2) 1.22 1.23 (THM)(3) 1.22 1.23 DNBR Safety Limit (TYP) 1.34 1.415 (THM) 1.34 1.415 DNBR Retained Margin (4) (TYP) 8.9% 13.1% (THM) 8.9% 13.1% Rod Bow DNBR Penalty -3.0% -2.9% Instrumentation Bias Penalty (TYP) -0.82% -0.67% (THM) -0.82% -0.69% MUR Penalty -2.8% - Transition Core DNBR Penalty (5) -9.0% -2.0%

DNBR Credit for FH reduction in 400V+ 9.0% - Available DNBR Margin (TYP) 2.3% 7.5% (THM) 2.3% 7.5%

Notes: 1. Steam line break is analyzed using the W-3 correlation with STDP. The correlation limit DNBR is 1.45 in the range of 500 to 1,000 psia and the safety limit DNBR is 1.593 which provides 8.9% margin to cover the rod bow penalty and retain generic margin for operational issues. Rod withdrawal from subcritical is also analyzed using the W-3 correlation (w/o spacer factor) with STDP below the bottom non-mixing vane (NMV) grid. The correlation limit DNBR is 1.30 above 1,000 psia and the safety limit DNBR is 1.428, which also provides 8.9% margin. WRB-1 with STDP is used for rod withdrawal from subcritical above the bottom NMV grid with a correlation limit DNBR of 1.17 and a safety limit DNBR of 1.285 which also provided 8.9% margin. 2. TYP Typical Cell

3. THM Thimble Cell 4. DNBR margin is the margin that exists between the safety limit and the design limit DNBRs. 5. Transition core penalty on 400V+ is offset by FdH reduction in once and twice-burned 400V+ fuel during transition cycles.

4-12 Table 4-5 Limiting Parameter Direction Parameter Limiting Direction for DNB F NH, nuclear enthalpy rise hot channel factor Maximum Heat generated in fuel (%) Maximum Reactor core heat output (MWt) Maximum Average heat flux (Btu/hr-ft

2) Maximum Nominal vessel/core inlet temperature (°F) Maximum Core pressure (psia) Minimum Pressurizer pressure (psia) Minimum Thermal design flow for non-RTDP analyses (gpm) Minimum Minimum measured flow for RTDP analyses (gpm) Minimum

5-1 5 ACCIDENT ANALYSIS This section presents the results of analyses a nd evaluations of non-loss-of-coolant accident (LOCA) transients, large-and small-break LOCAs, and containment integrity.

5.1 NON-LOCA TRANSIENTS Non-LOCA transient analyses and evaluations were performed to support implementation of the 422V+ fuel transition at the Prairie Island Units 1&2. All non-LOCA events analyzed in the Prairie Island Updated Safety Analysis Report (USAR) were analyzed or evaluated. Program features that were considered include: A nuclear steam supply system (NSSS) power level of 1,690 MWt (including 7 MWt of reactor coolant pump (RCP) heat in addition to the core power of 1,683 MWt) A nominal, full-power reactor coolant vessel average temperature (T avg) of 560°F Reactor coolant system (RCS) thermal design flow (TDF) of 178,000 gpm Westinghouse Model 51 original steam generators (OSGs) with a maximum steam generator tube plugging (SGTP) of 25 percent and Framatome ANP Model 56/19 replacement steam generators (RSGs) with a maximum SGTP of 10 percent (A maximum loop-to-loop tube plugging asymmetry of 10 percent has been addressed.) Nominal operating RCS (pressurizer) pressure of 2,250 psia Computer codes used in the analyses are FACTRAN (Reference 1), RETRAN (Reference 2), LOFTRAN (Reference 3), ANC (Reference 4), TWINKLE (Reference 5), and VIPRE (Reference 6). Each of the USAR transients listed in Table 5.1-1 were evaluated or analyzed in support of the 422V+ Transition Program as specified in Reference 23. In general, an evaluation was performed for events that the transient conditions of the RCS and main steam system (MSS) were determined to be insignificantly impacted by the change in fuel, whereas analysis was performed for those events in which the change in fuel has a significant impact on the transient conditions of the RCS and MSS. These transient evaluations and analyses demonstrate that all applicable safety analysis acceptance criteria are satisfied for Prairie Island as discussed below. Specifically, for the safety analysis events evaluated in support of the fuel transition, the existing results related to departure from nucleate boiling (DNB) concerns remain conservative as the benefit associated with the increase in the nominal DNB ratio (DNBR) outweighs any penalties associated with the fuel transition. For the events re-analyzed in support of the fuel transition, it is concluded that all applicable criteria are met. The following subsections provide an assessment of the impact of the transition to 422V+ fuel for transients that were re-analyzed or evaluated.

5-2 Table 5.1-1 Non-LOCA Transients Evaluated or Analyzed Transient USAR Section Notes Uncontrolled RCCA Withdrawal from a Subcritical Condition 14.4.1 1 Uncontrolled RCCA Withdrawal at Power 14.4.2 2 RCCA Misalignment - Statically Misaligned RCCAs RCCA Misalignment - Dropped RCCA 14.4.3 2 Chemical and Volume Control System Malfunction 14.4.4 2 Startup of an Inactive Reactor Coolant Loop 14.4.5 2 Feedwater Malfunction - Feedwater Temperature Reduction 2 Feedwater Malfunction - Feedwater Flow Increase 14.4.6 2 Excessive Load Increase Incident 14.4.7 2 Loss of Reactor Coolant Flow - Flow Coast Down Loss of Reactor Coolant Flow - Locked Rotor 14.4.8 1 Loss of External Electrical Load 14.4.9 2 Loss of Normal Feedwater 14.4.10 2 Loss of All AC Power to the Station Auxiliaries 14.4.11 2 Rupture of a Steam Pipe - Zero-Power Core Response Rupture of a Steam Pipe - Full-Power Core Response 14.5.5 2 RCCA Ejection 14.5.6 1 ATWS 14.8 2 Notes: 1. Complete Analysis 2. Evaluation

5-3 5.1.1 Uncontrolled RCCA Withdrawal from a Subcritical Condition (USAR Section 14.4.1) 5.1.1.1 Accident Description The rod cluster control assembly (RCCA) withdrawal accident is defined as an uncontrolled addition of reactivity to the reactor core caused by withdrawal of RCCA banks resulting in a power excursion. While the occurrence of a transient of this type is unlikely, such a transient could be caused by a malfunction of the reactor control or the control rod drive system. This could occur with the reactor either subcritical, at hot zero power (HZP), or at power. The "at-power" case is discussed in subsection 5.1.2. Withdrawal of an RCCA bank adds reactivity at a prescribed and controlled rate to bring the reactor from a subcritical condition to a low-power level during startup. Although the initial st artup procedure uses the method of boron dilution, the normal startup is with RCCA bank withdrawal. RCCA bank movement can cause much faster changes in reactivity than can be made by changing boron concentration (see subsection 5.1.4, Chemical and Volume Control System Malfunction (that is, Uncontrolled Boron Dilution)). The RCCA drive mechanisms are wired into preselecte d bank configurations that are not altered during core life. These circuits prevent RCCAs from being withdrawn in other than their respective banks. Power supplied to the rod banks is controlled so that no more than two banks can be withdrawn at any time and in their proper withdrawal sequence. The RCCA drive mechanisms are of the magnetic latch type; coil actuation is sequenced to provide variable speed travel. The analysis of the maximum reactivity insertion rate includes the assumption of the simultane ous withdrawal of the two sequential banks having the maximum combined worth at maximum speed. The neutron flux response to a continuous reactivity insertion is characterized by a very fast flux increase terminated by the reactivity feedback effect of the negative Doppler coefficient. This self limitation of the power burst is of primary importance since it limits the power to a tolerable level during the delay time for protective action. Should a continuous control rod assembly withdrawal event occur, the following automatic features of the reactor protection system are available to terminate the transient: The source range high neutron flux reactor trip is actuated when either of two independent source range channels indicates a neutron flux level above a preselected manually adjustable setpoint and provides primary protection below the P6 permissive. This trip function may be manually bypassed when either intermediate range flux channel indicates a flux level above P6. It is automatically reinstated when both intermediate-r ange channels indicate a flux level below P6. The intermediate range high neutron flux reactor trip is actuated when either of two independent intermediate-range channels indicates a flux level above a preselected manually adjustable setpoint. This trip function may be manually bypassed when two of the four power-range channels give readings above the P10 permissive (approximately 10 percent of full-power) and is automatically reinstated when three of the four channels indicate a power below P10. The power range high neutron flux reactor trip (low setting) is actuated when two-out-of-four power-range channels indicate a power level above a preselected manually adjustable setpoint (allowable value 40 percent of full-power). This trip function may be manually bypassed when 5-4 two of the four power-range cha nnels indicate a power level above the P10 permissive and is automatically reinstated when three of the four channels indicate a power level below P10. The power range high neutron flux reactor trip (high setting) is actuated when two-out-of-four power-range channels indicate a power level above a preset setpoint (allowable value 110 percent power). This trip function is always active while the reactor is at power. In addition, control rod stops on high intermediate range flux (one-out-of-two) and high power-range flux (one-out-of-four) serve to cease rod withdrawal and prevent the need to actuate the intermediate-range flux trip and the power-range flux trip, respectively. 5.1.1.2 Method of Analysis The analysis of the uncontrolled RCCA bank withdrawal from subcritical accident is performed in three stages. First, a spatial neutron kinetics computer code, TWINKLE, is used to calculate the core average nuclear power transient, including the various core feedback effects, that is, Doppler and moderator reactivity. FACTRAN uses the average nuclear power calculated by TWINKLE and performs a fuel rod transient heat transfer calculation to determine the average heat flux and temperature transients. Finally, the peak core-average heat flux calculated by FACTRAN is used in VIPRE for transient DNBR calculations. In order to give conservative results for a startup accident, the following assumptions are made: 1. Since the magnitude of the power peak reached during the initial part of the transient for any given rate of reactivity insertion is strongly dependent on the Doppler reactivity coefficient, a conservatively low (absolute magnitude) value for the Doppler power defect is used (1,100 pcm). 2. The contribution of the moderator reactivity coefficient is negligible during the initial part of the transient because the heat transfer time constant between the fuel and the moderator is much longer than the neutron flux response time constant. However, after the initial neutron flux peak, the isothermal temperature coefficient (ITC) can affect the succeeding rate of power increase. The effect of moderator temperature changes on the rate of nuclear power increase is calculated in TWINKLE based on temperature-dependent moderator cross-sections. The ITC value used in the rod withdrawal from subcritical event analysis is + 5 pcm/°F. 3. The analysis assumes the reactor to be at HZP nominal temperature of 547 °F. This assumption is more conservative than that of a lower initial system temperature (that is, shutdown conditions). The higher initial system temperature yields a larger fuel-to-water heat transfer coefficient, a larger specific heat of the water and fuel, and a less-negative (smaller absolute magnitude) Doppler coefficient. The less-negative Doppler coefficient reduces the Doppler feedback effect, thereby increasing the neutron flux peak. The high neutron flux peak combined with a high fuel-specific heat and larger heat transfer coefficient yields a larger peak heat flux. The analysis assumes the initial effective multiplication factor (keff) to be 1.0, since this results in the maximum neutron flux peak.

5-5 4. Reactor trip is assumed to be initiated by power-range high neutron flux (low setting). The most adverse combination of instrumentation and setpoint errors are accounted for by assuming a 10-percent increase in the power range flux trip setpoint (low setting), raising it from the allowable value of 40 percent to a value of 50 percent. Figure 5.1.1-1 shows that the rise in nuclear flux is so rapid that the effect of error in the trip setpoint on the actual time at which the rods are released is negligible. In addition, the total reactor trip reactivity is based on the assumption that the highest worth RCCA is stuck in its fully withdrawn position. Further, the delays for trip signal actuation and control rod assembly release are account ed for in the reactor trip delay time. 5. The maximum positive reactivity insertion rate assumed (93.5 pcm/in) is a plant-specific value confirmed for each reload cycle and is equal to that for the simultaneous withdrawal of the two sequential control banks having the greatest combined worth at a conservative speed (48.125 in/min, which corresponds to 77 steps/min). It should be noted that the assumption of 77 steps/min as the maximum rod withdrawal speed is based upon the recommendations of NSAL-01-001 (Reference 7). 6. The DNB analysis assumes the most-limiting axial and radial power shapes possible during the fuel cycle associated with having the two highest combined worth banks in their high worth position. 7. The analysis assumes the initial power level to be below the power level expected for any HZP just-critical condition (10-9 fraction of nominal power). The combination of highest reactivity insertion rate and low initial power produces the highest peak heat flux. 8. Note that the Technical Spec ifications require that two RCPs be in operation for Mode 2 and Mode 3 when capable of rod withdrawal. The analysis is performed at HZP conditions with one RCP in operation, which is conservative, and bounds this accident in lower modes. This assumption also minimizes the resulting DNBR. 9. The accident analysis employs the Standard Thermal Design Procedure (STDP) methodology. Use of the STDP stipulates that the RCS flow rate will be based on the thermal design flow and that the RCS pressure is the nominal pressure minus the uncertainty. Since the event is analyzed from HZP, the steady-state STDP uncertainties on core power and RCS average temperature are not considered in defining the initial conditions. 10. A core flow reduction corresponding to the maximum potential reactor coolant loop flow asymmetry of 5 percent, associated with a maximum loop-to-loop SGTP imbalance of 10 percent, has been applied. 11. The fuel rod heat transfer calculations performed to determine temperature transients during this event assume a total peaking factor or hot channel factor, FQ, that is a function of the axial and radial power distributions. The conservatively high value used in this analysis is presented in Table 5.1.1-1.

5-6 12. The transition to Westinghouse 422V+ fuel with up to 8 w/o gadolinium content was considered in this analysis and the most bounding results are reported here. 13. Note that different DNBR correlations with different DNBR limits were used for the regions above and below the first mixing vane grid. Both sets of results are listed in Table 5.1.1-1. 5.1.1.3 Results Figures 5.1.1-1 through 5.1.1-5 show the transient behavior for a reactivity insertion rate of 75 pcm/sec.

The rate is greater than that calculated for the two highest worth sequential control banks, with both assumed to be in their highest incremental worth region. Figure 5.1.1-1 shows the neutron flux transient. The neutron flux overshoots the full-power nominal value for a very short period of time; therefore, the energy release and fuel temperature increase are relatively small. The heat flux response of interest for the DNB considerations is shown in Figure 5.1.1-2. The beneficial effect of the inherent thermal lag in the fuel is evidenced by a peak heat flux of much less than the nominal full-power value. Figures 5.1.1-3 through 5.1.1-5 show the transient response of the hot spot fuel centerline, fuel average, and cladding temperatures, respectively. DNBR calculations indicate that the minimum DNBR remains above the safety analysis limit value at all times. Table 5.1.1-1 presents the assumptions and results of the analysis. Table 5.1.1-2 presents the calculated sequence of events. After reactor trip, the plant returns to a stable condition. The plant may subsequently be cooled down further by following normal shutdown procedures. 5.1.1.4 Conclusions In the event of an RCCA withdrawal accident from the subcritical conditi on, the core and the RCS are not adversely affected since the combination of thermal power and coolant temperature result in a DNBR greater than the limit value. Thus, no fuel or cladding damage is predicted as a result of this transient.

5.1.2 Uncontrolled

RCCA Withdrawal at Power (USAR Section 14.4.2) The uncontrolled RCCA bank withdrawal at power event is defined as the inadvertent addition of reactivity to the core caused by the withdrawal of RCCA banks when the core is above the power defined by the P-10 setpoint. The reactivity insertion resulting from the bank (or banks) withdrawal will cause an increase in the core nuclear power and subsequent increase in the core heat flux. An RCCA bank withdrawal can occur with the reactor subcritical, at HZP, or at power. The uncontrolled RCCA bank at power event is analyzed for Mode 1 (power operation). The uncontrolled RCCA bank withdrawal from a subcritical or low-power condition is considered as an independent event in subsection 5.1.1. To address the 422V+ Fuel Transition Program, an evaluation was performed to disposition the changes associated with the new fuel. The results of the evaluation demonstrate that the event is insignificantly impacted by the transition to 422V+ fuel. It is therefore concluded that the high neutron flux, the positive flux rate trip, and overtemperature delta-T (OTT) trip functions continue to provide adequate protection over the entire range of possible reactivity insertion rates such that the minimum calculated DNBR is always greater than the safety analysis limit value. In addition, it is concluded that the peak system pressures in the 5-7 RCS and MSS do not exceed 110 percent of their respective design pressures. For all cases, operation with either the Westinghouse OSGs or the Framatome ANP RSGs was addressed. Thus, all pertinent criteria are met for the uncontrolled RCCA bank withdrawal at power transient in support of the Prairie Island 422V+ Fuel Transition Program.

5.1.3 RCCA Misalignment (USAR Section 14.4.3) An RCCA Misalignment event is a Condition II event that is assumed to be initiated by a single electrical or mechanical failure. The resulting negative reactivity insertion causes nuclear power to rapidly decrease. An increase in the hot channel factor may occur due to skewed power distribution representative of a dropped or misaligned RCCA configuration. The RCCA misalignment accidents include: Dropped full-length RCCAs Dropped full-length RCCA banks Statically misaligned full-length RCCAs To address the dropped rod event, the LOFTRAN computer code (Reference 3) is used to generate dropped RCCA(s) reactor system statepoints for bounding ranges of dropped rod and control bank reactivity worths. For each case, the statepoint represents the transient system conditions (temperature, pressure, and power) at the limiting point in the transient. With respect to the 422V+ Fuel Transition Program, the statepoints remain valid as the fuel type is not explicitly modeled. As such, a detailed DNB analysis was performed using the existing statepoints. The results of the detailed DNB analysis demonstrate that the minimum calculated DNBR is always greater than the safety analysis limit value. Therefore, no fuel or cladding damage is predicted and all applicable acceptance criteria are met in support of the Prairie Island 422V+ Fuel Transition Program. In order to address the misaligned RCCA event, detailed DNB analysis was performed to calculate an FH limit that corresponds to the safety analysis DNBR limit. The core design conceptual models were then used to show that the misaligned rod FH limit is not exceeded for the limiting misaligned rod events. Therefore, no fuel or cladding damage is predicted and all applicable acceptance criteria are met in support of the Prairie Island 422 V+ Fuel Transition Program.

5.1.4 Chemical

and Volume Control System Malfunction (USAR Section 14.4.4) The accident considered here is the malfunction of the chemical and volume control system (CVCS) resulting in the injection of non-borated water at the maximum possible flow rate to the RCS under at-power conditions. With the reactor in automatic mode, the decrease in the boron concentration will cause the power and temperature to increase, resulting in the insertion of the RCCAs and a decrease in shutdown margin. With the reactor in manual mode, the decrease in the boron concentration will cause the power and temperature to increase. This will eventually result in the OTT or overpower delta-T (OPT) reactor trip if the operator does not intervene.

5-8 To address the 422V+ Fuel Transition Program, an evaluation was performed to disposition the changes associated with the new fuel. The results of the evaluation demonstrate that the event is insignificantly impacted by the transition to 422V+ fuel. It is, therefore, concluded that the operator has sufficient time to terminate the RCS dilution before a complete loss of shutdown margin occurs. With respect to a CVCS malfunction occurring while subcritical, minimum shutdown boron concentrations are calculated for each cycle to ensure the acceptance criteria described in USAR Section 14.4.4.2.4 are met. Since this is a cycle-specific calculation using the specific fuel types installed, the use of 422V+ fuel will be addressed and will also not impact any existing analyses. This process will continue to be followed on a reload-specific basis consistent with the current approach. Thus, all pertinent criteria are met for the CVCS malfunction transient in support of the Prairie Island 422V+ Fuel Transition Program.

5.1.5 Startup

of an Inactive Reactor Coolant Loop (USAR Section 14.4.5) Following the startup of an inactive RCP, flow in the inactive reactor coolant loop will accelerate to full flow in the forward direction over a period of several seconds. However, the Prairie Island Technical Specifications require that both RCPs be operating when the reactor is in Mode 1 or Mode 2. Therefore, the maximum initial core power level for the startup of an inactive coolant loop is near 0 MWt. Under these conditions, there can be no significant reactivity insertion because the reactor control system is initially a nearly uniform temperature. Furthermore, the reactor will initially be subcritical by the Technical Specification requirement. Thus, there will be no increase in core power, and no automatic or manual protective action is required.

The startup of an inactive reactor coolant loop event results in an increase in reactor vessel flow while the reactor remains in a subcritical condition. No analysis is required to show that the DNB limit is satisfied for this event.

5.1.6 Feedwater

Malfunction (USAR Section 14.4.6) A change in steam generator feedwater conditions that results in an increase in feedwater flow or a decrease in feedwater temperature could result in excessive heat removal from the plant primary coolant system. Such changes in feedwater flow or feedwater temperature are a result of a failure of a feedwater control valve or feedwater bypass valve, failure in the feedwater control system, or operator error. The occurrences of these failures that result in an excessive heat removal from the plant primary coolant system cause the primary side temperature and pressure to decrease significantly. The existence of a negative moderator and fuel temperature reactivity coefficients, and the actions initiated by the reactor rod control system can cause core reactivity to rise, as the primary side temperature decreases. In the absence of the reactor protection system (RPS) reactor trip or other protective action, this increase in core power, coupled with the decrease in primary side pressure, can challenge the core thermal limits. To address the 422V+ Fuel Transition Program, an evaluation was performed to disposition the changes associated with the new fuel. The results of the evaluation demonstrate that the event, as analyzed, is insignificantly impacted by the transition to 422V+ fuel. It is, therefore, concluded that the minimum 5-9 calculated DNBR is always greater than the safety analysis limit value. It is also concluded that the feedwater malfunction event analyzed for an increase in feedwater flow from zero-power initial conditions was determined to be less severe than that of the steam line break - core response event discussed in subsection 5.1.13. Additionally, it is concluded that the decreased enthalpy caused by the feedwater temperature reduction event continues to be bounded by an equivalent enthalpy reduction that results from an excessive load increase incident. Thus, all pertinent criteria are met for the feedwater malfunction transient in support of the Prairie Island 422V+ Fuel Transition Program.

5.1.7 Excessive

Load Increase Incident (USAR Section 14.4.7) An Excessive Load Increase (ELI) incident is defined as an event resulting in a rapid increase in the steam generator steam flow that causes a power mismatch between the reactor core power and the steam generator load demand. The reactor control system is typically designed to accommodate a 10-percent step load increase or a 5-percent per minute ramp load increase (without a reactor trip) in the range of 15 to 95 percent of full power. Any loading rate in excess of these values could cause a reactor trip actuated by the RPS. This accident could result from either an administrative violation, such as excessive loading by the operator, or an equipment malfunction in either the steam dump control or the turbine speed control. During power operation, steam dump to the condenser is controlled by the reactor coolant condition signals; that is, a high reactor coolant temperature indicates a need for steam dump. A single controller malfunction does not cause steam dump; an interlock is provided that blocks the opening of the valves unless a large turbine load decrease or turbine trip has occurred.

The possible consequence of this accident (assuming no protective functions) is a DNB with subsequent fuel damage. To address the 422V+ Fuel Transition Program, an evaluation was performed to disposition the changes associated with the new fuel. The results of this evaluation demonstrate that the event is insignificantly impacted by the transition to 422V+ fuel. It is, therefore, concluded that the minimum calculated DNBR is always greater than the safety analysis limit value consistent with the previous analysis. The results of the analysis do not challenge the RCS overpressure limit since this is a cooldown event. Moreover, as the event is initiated by an increase in the MSS flow rate, which results in overcooling the RCS and a decrease in the MSS pressure, the MSS overpressure limit is not challenged during the event. Thus, all pertinent criteria are met for the ELI transient in support of the Prairie Island 422V+ Fuel Transition Program.

5.1.8 Loss of Reactor Coolant Flow - Flow Coastdown (USAR Section 14.4.8)

The loss of reactor coolant flow events are categorized as follows in the Prairie Island Units 1&2 USAR: Flow coastdown accidents Locked-rotor accident 5-10 The first category includes the partial and complete loss of reactor coolant flow events. The second category includes the hypothetical event that addresses an instantaneous seizure of an RCP rotor. 5.1.8.1 Partial Loss of Reactor Coolant Flow Accident Description The Partial Loss-of-Coolant-Flow accident can result from a mechanical or electrical failure in an RCP, or from a fault in the power supply to the RCP. If the reactor is at power at the time of the accident, the immediate effect of loss-of-coolant flow is a rapid increase in the coolant temperature. This increase could result in DNB subsequent fuel damage if the reactor is not tripped promptly. Normal power for the pumps is available through individual buses connected to the generator and the offsite power system. When a generator trip occurs, the buses continue to be supplied from external power lines, and the pumps continue to supply coolant to the core. The necessary protection against a partial loss-of-coolant-flow accident is provided by the low primary coolant flow reactor trip signal, which is actuated in any reactor coolant loop by two-out-of-three low flow signals. Above the Permissive 8 setpoint, low flow in either loop will actuate a reactor trip. Above the Permissive 7 setpoint, low flow in both loops will actuate a reactor trip. Method of Analysis The loss of an RCP with both loops in operation event is analyzed to show that: (1) the integrity of the core is maintained as the DNBR remains above the safety analysis limit value, and (2) the peak RCS and secondary system pressures remain below the design limits. Of these, the primary concerns are DNB and assuring that the DNBR limit is met. The loss of an RCP event is analyzed with two computer codes. First, the RETRAN computer code (Reference 2) is used to calculate the loop and core flow during the transient, the time of reactor trip based on the calculated flows, the nuclear power transient, and the primary system pressure and temperature transients. The VIPRE computer code is then used to calculate the hot channel heat flux transient and DNBR, based on the nuclear power and RCS temperature (enthalpy), pressure, and flow from RETRAN. The DNBR transients presented represent the minimum of the typical or thimble cell. This event is analyzed with the Revised Thermal Design Procedure (RTDP). Initial reactor power, pressurizer pressure, and RCS temperature are assumed to be at their nominal values. Minimum measured flow is also assumed. A conservatively large absolute value of the Doppler-only power coefficient is used, along with the most-positive ITC limit for full-power operation (0 pcm/°F). These assumptions maximize the core power during the initial part of the transient when the minimum DNBR is reached. A limiting end-of-cycle (EOC) DNB axial power shape is assumed in VIPRE for the calculation of DNBR. This shape provides the most limiting minimum DNBR for the loss-of-flow events. A conservatively low trip reactivity value (4.0-percent ) is used to minimize the effect of rod insertion following reactor trip and maximize the heat flux state point used in the DNBR evaluation for this event.

5-11 This value is based on the assumption that the highest worth RCCA is stuck in its fully withdrawn position. A conservative trip reactivity worth versus rod position was modeled in addition to a conservative rod drop time (2.4 seconds to dashpot for flow equal to mechanical design flow). The trip reactivity versus rod position curve is confirmed to be valid as part of the Reload Safety Analysis Checklist (RSAC) verification process. The flow coastdown analysis is based on a momentum balance around each reactor coolant loop and across the reactor core. This momentum balance is combined with the continuity equation, a pump momentum balance, and the pump characteristics. Also, it is based on conservative estimates of system pressure losses. The Framatome RSGs were modeled. However, the analysis applies to both the Westinghouse OSGs and Framatome RSGs since this event is not sensitive to the secondary side modeling. A maximum, uniform, SGTP level was assumed in the RETRAN analysis. Additionally, a core flow reduction of 1.1 percent, which addresses the potential reactor coolant flow asymmetry associated with a maximum loop-to-loop SGTP imbalance of 10 percent, was applied.

Results Figures 5.1.8-1 through 5.1.8-8 illustrate the transient response for the loss of an RCP with both loops initially in operation. The minimum DNBR is 1.798/1.833 (thimble/typical), which occurred at 4.20 seconds (DNBR limit: 1.415/1.415 (thimble/typical)).

The calculated sequence of events table is shown in Table 5.1.8-1. This event trips on a low primary reactor coolant flow trip setpoint, which is assumed to be 87 percent of loop flow. Following reactor trip, the affected RCP will continue to coast down, and the core flow will reach a new equilibrium value corresponding to the remaining pump still in operation. With the reactor tripped, a stable plant condition will eventually be attained. Normal plant shutdown may then proceed. Conclusions The analysis performed has demonstrated that, for the partial loss-of-coolant event, the DNBR does not decrease below the limit value at any time during the transient. Therefore, no fuel or cladding damage is predicted and all applicable acceptance criteria are met. 5.1.8.2 Complete Loss of Forced Reactor Coolant Flow Accident Description A complete loss of forced reactor coolant flow may result from a simultaneous loss of electrical supplies to all RCPs. If the reactor is at power at the time of the accident, the immediate effect of the loss-of-coolant flow is a rapid increase in the coolant temperature. This increase could result in DNB with subsequent fuel damage if the reactor were not tripped promptly. Normal power for the RCPs is available through buses from a transformer connected to the generator and the offsite power system. Each pump is on a separate bus. When a generator trip occurs, the buses 5-12 continue to be supplied from external power lines and the pumps continue to supply coolant flow to the core. The following signals provide the necessary protection against a complete loss-of-flow accident: RCP power supply undervoltage reactor trip Low reactor coolant loop flow reactor trip Pump circuit breaker opening (RCP supply underfrequency opens pump circuit breaker, which trips the reactor) The reactor trip on RCP undervoltage is provided to protect against conditions that can cause a loss of voltage to all RCPs; that is, station blackout. This function is blocked below approximately 10-percent nuclear instrumentation system and 10-percent turbine power (Permissive 7). The reactor trip on low primary coolant flow is provided to protect against loss-of-flow conditions that affect one or both reactor coolant loops. This function is generated by two-out-of-three low flow signals per reactor coolant loop. Above the Permissive 8 setpoint, low flow in either loop will actuate a reactor trip. Above the Permissive 7 setpoint, low flow in both loops will actuate a reactor trip. The reactor trip on RCP underfrequency (pump circuit brea ker opening) is available to trip the reactor for an underfrequency condition, resulting from frequency disturbances on the power grid. However, the analysis conservatively assumes that this function is not available to provide a reactor trip. Therefore, the low primary coolant flow reactor trip function is assumed to provide primary protection against an underfrequency event. This event is conservatively analyzed to the following acceptance criteria: Pressure in the RCS and MSS should be maintained below 110 percent of the design values. Fuel cladding integrity shall be maintained by ensuring that the minimum DNBR remains above the limit value. An incident of moderate frequency should not generate a more serious plant condition without other faults occurring independently. Method of Analysis The complete loss-of-flow transient is analyzed as a loss of both RCPs with both loops in operation. The event is analyzed to show that the integrity of the core is maintained as the DNBR remains above the safety analysis limit value. The loss-of-flow events do result in an increase in RCS and MSS pressures, but these pressure increases are generally not severe enough to challenge the integrity of the RCS and MSS. Since the maximum RCS and MSS pressures do not exceed 110 percent of their respective design pressures for the loss-of-load event, it is concluded that the maximum RCS and MSS pressures will also remain below 110 percent of their respective design pressures for the loss-of-flow events.

5-13 The transients are analyzed with two computer codes. First, the RETRAN computer code (Reference 2) is used to calculate the loop and core flow during the transient, the time of reactor trip based on the calculated flows, the nuclear power transient, and the primary system pressure and temperature transients. The VIPRE computer code is then used to calculate the heat flux and DNBR transients based on the nuclear power and RCS temperature (enthalpy), pressure, and flow from RETRAN. The DNBR transients presented represent the minimum of the typical or thimble cell for the fuel. This event is analyzed with RTDP. Initial reactor power, pressurizer pressure, and RCS temperature are assumed to be at their nominal values. Minimum measured flow is also assumed. A conservatively large absolute value of the Doppler-only power coefficient is used, along with the most-positive ITC limit for full-power operation (0 pcm/°F). These assumptions maximize the core power during the initial part of the transient when the minimum DNBR is reached. A limiting EOC DNB axial power shape is assumed in VIPRE for the calculation of DNBR. This shape provides the most limiting minimum DNBR for the loss-of-flow events. A conservatively low trip reactivity value (4.0-percent ) is used to minimize the effect of rod insertion following reactor trip and maximize the heat flux state point used in the DNBR evaluation for this event. This value is based on the assumption that the highest worth RCCA is stuck in its fully withdrawn position. A conservative trip reactivity worth versus rod position was modeled in addition to a conservative rod drop time (2.4 seconds to dashpot at mechanical design flow). The trip reactivity versus rod position curve is confirmed to be valid as part of the RSAC verification process. The flow coastdown analysis is based on a momentum balance around each reactor coolant loop and across the reactor core. This momentum balance is combined with the continuity equation, a pump momentum balance, and the pump characteristics. Also, it is based on conservative estimates of system pressure losses.

The Framatome RSGs were modeled. However, the analysis applies to both the Westinghouse OSGs and Framatome RSGs since this event is not sensitive to the secondary side modeling. A maximum, uniform, SGTP level was assumed in the RETRAN analysis. Reactor coolant system loop flow asymmetry due to a loop-to-loop SGTP imbalance does not need to be considered for transients in which both RCPs experience a coastdown.

Results Figures 5.1.8-9 through 5.1.8-16 illustrate the transient response for the complete loss of flow associated with a loss of power to both RCPs with both loops in operation. The minimum DNBR is 1.466/1.470 (thimble/typical), which occurred at 4.65 seconds (DNBR limit: 1.415/1.415 (thimble/typical)). The calculated sequence of events for the complete loss-of-flow case is shown on Table 5.1.8-2. Following reactor trip, the RCPs will continue to coast down, and natural circulation flow will eventually be established. With the reactor tripped, a stable plant condition will eventually be attained. Normal plant shutdown may then proceed.

5-14 Conclusions The analysis performed has demonstrated that for the complete loss-of-flow event, the DNBR does not decrease below the limit value at any time during the transient. Therefore, no fuel or cladding damage is predicted and all applicable acceptance criteria are met.

5.1.9 Loss of Reactor Coolant Flow - Locked Rotor (USAR Section 14.4.8) 5.1.9.1 Accident Description The postulated locked-rotor accident is an instantaneous seizure of an RCP rotor. Flow through the affected reactor coolant loop is rapidly reduced, leading to an initiation of a re actor trip on a low-flow signal. The consequences of a postulated pump shaft break accident are similar to the locked-rotor event. With a broken shaft, the impeller is free to spin, as opposed to it being fixed in position during the locked-rotor event. Therefore, the initial rate of reduction in core flow is greater during a locked-rotor event than in a pump shaft break event because the fixed shaft causes greater resistance than a free-spinning impeller early in the transient, when flow through the affected loop is in the positive direction. As the transient continues, the flow direction through the affected loop is reversed. If the impeller is able to spin free, the flow to the core w ill be less than that available with a fixed shaft during periods of reverse flow in the affected loop. Because peak pressure, cladding temperature, and DNB occur very early in the transient, the reduction in core flow during the period of forward flow in the affected loop dominates the severity of the results. Consequently, the bounding results for the locked-rotor transients also are applicable to the RCP shaft break. After the locked rotor, reactor trip is initiated on an RCS low-flow signal. The analysis assumes that the unaffected RCP continues to operate throughout the event. Following initiation of the reactor trip, heat stored in the fuel rods continues to be transferred to the core coolant causing the coolant to expand. At the same time, heat transfer to the shell side of the steam generators is reduced. This is because, first, the reduced flow results in a decreased tube side film coefficient and then because the reactor coolant in the tubes cools down while the shell side temperature increases (turbine steam flow is reduced to zero upon plant trip). The rapid expansion of the coolant in the reactor core, combined with reduced heat transfer in the steam generators, causes an insurge into the pressurizer and a pressure increase throughout the RCS. The insurge into the pressurizer compresses the steam volume, actuates the automatic spray system, opens the power-operated relief valves (PORVs), and opens the pressurizer safety valves, in that sequence. The two PORVs are designed for reliable operation and would be expected to function properly during the accident. However, for conservatism in the peak-pressure evaluation, their pressure-reducing effect and the pressure-reducing effect of the pressurizer sprays are not included in the analysis.

5-15 The locked-rotor event is analyzed to the following criteria: Pressure in the RCS should be maintained below the designated limit (see below). Coolable core geometry is ensured by showing that the peak cladding temperature and maximum oxidation level for the hot spot are below 2,700°F and 16.0 percent by weight, respectively. Activity release is such that the calculated doses meet 10 CFR Part 100 guidelines. For Prairie Island Units 1&2, the locked-rotor RCS pressure limit is equal to 110 percent of the design value, or 2,750 psia. For the secondary side, the locked-rotor pressure limit is also assumed to be equal to 110 percent of design pressure, or 1,210 psia. Since the loss-of-load analysis bounds the locked rotor, a specific MSS overpressurization analysis is not performed. A hot spot evaluation is performed to calculate the peak cladding temperature and maximum oxidation level. Finally, a calculation of the "rods-in-DNB" is performed for use in the radiological dose analysis. 5.1.9.2 Method of Analysis The locked-rotor transient is analyzed with two primary computer codes. First, the RETRAN computer code (Reference 2) is used to calculate the loop and core flow during the transient, the time of reactor trip based on the calculated flows, the nuclear power transient, and the primary system pressure and temperature transients. The VIPRE code is then us ed to calculate the rods-in-DNB and peak cladding temperature using the nuclear power and RCS temperature (enthalpy), pressure, and flow from RETRAN. For the case analyzed to determine the maximum RCS pressure and peak cladding temperature, the plant is assumed to be in operation under the most adverse steady-state operating conditions; that is, a maximum steady-state thermal power, maximum steady-state pressure, and maximum steady-state coolant average temperature. The case analyzed to determine the rods-in-DNB utilizes the RTDP methodology. Initial reactor power, pressurizer pressure, and RCS temperature are assumed to be at their nominal values. Minimum measured flow is also assumed. The Framatome ANP Model 56/19 RSGs were modeled. However, the analysis applies to both the Westinghouse OSGs and Framatome RSGs since this event is not sensitive to the secondary side modeling. A maximum, uniform, SGTP level was assumed in the RETRAN analysis. However, a core flow reduction of 1.1 percent, which addresses the potential reactor coolant flow asymmetry associated with a maximum loop-to-loop SGTP imbalance of 10 percent, was applied. A conservatively large absolute value of the Doppler-only power coefficient is used, along with the most-positive ITC limit for full-power operation (0 pcm/°F). These assumptions maximize the core power during the initial part of the transient when the peak RCS pressures and hot spot results are reached. A conservatively low trip reactivity value (4.0-percent ) is used to minimize the effect of rod insertion following reactor trip and maximize the heat flux state point used in the DNBR evaluation for this event. This value is based on the assumption that the highest worth RCCA is stuck in its fully withdrawn 5-16 position. A conservative trip reactivity worth versus rod position was modeled in addition to a conservative rod drop time (2.4 seconds from dashpot at mechancial design flow). The trip reactivity versus rod position curve is confirmed to be valid as part of the RSAC verification process. For the peak RCS pressure evaluation, the initial pre ssure is conservatively estim ated as 60 psi above the nominal pressure (2,250 psia) to allow for errors in the pressurizer pressure measurement and control channels. This is done to obtain the highest possible rise in the coolant pressure during the transient. The peak RCS pressure occurs in the lower plenum of the vessel. The pressure transient in the lower plenum is shown in Figure 5.1.9-6.

For this accident, an evaluation of the consequences with respect to the fuel rod thermal transient is performed. The evaluation incorporates the assumption of rods going into DNB as a conservative initial condition to determine the cladding temperature and zirconium water reaction resulting from the locked rotor. Results obtained from the analysis of this hot spot condition represent the upper limit with respect to cladding temperature and zirconium water reaction. In the evaluation, the rod power at the hot spot is assumed to be 2.74 times the average rod power (that is, FQ = 2.74) at the initial core power level.

Film Boiling Coefficient The film boiling coefficient is calculated in the VIPRE code using the Bishop-Sandberg-Tong film boiling correlation. The fluid properties are evaluated at film temperature. The program calculates the film coefficient at every time step based upon the actual heat transfer conditions at the time. The nuclear power, system pressure, bulk density, and RCS flow rate as a function of time are based on the RETRAN results. Fuel Cladding Gap Coefficient The magnitude and time dependence of the heat transfer coefficient between fuel and cladding (gap coefficient) has a pronounced influence on the thermal results. The larger the value of the gap coefficient, the more heat is transferred between th e pellet and cladding. Based on investigations on the effect of the gap coefficient upon the maximum cladding temperature during the transient, the gap coefficient was assumed to increase from a steady-state value consistent with initial fuel temperature to approximately 10,000 Btu/hr-ft 2-°F at the initiation of the transient. Therefore, the large amount of energy stored in the fuel because of the small initial value is released to the cladding at the initiation of the transient. Zirconium-Steam Reaction The zirconium-steam reaction can become significant above 1,800°F (cladding temperature). The Baker-Just parabolic rate equation is used to define the rate of zirconium-steam reaction. The effect of the zirconium-steam reaction is included in the calculation of the hot spot cladding temperature transient. 5.1.9.3 Results Figures 5.1.9-1 through 5.1.9-8 illustrate the transient response for the locked-rotor event (peak RCS pressure/peak cladding temperature case). The peak RCS pressure is 2,574 psia and is less than the 5-17 acceptance criterion of 2,750 psia. Also, the peak cladding temperature is 1,926°F, which is considerably less than the limit of 2,700°F. The zirconium-steam reaction at the hot spot is 0.44 percent by weight, which meets the criterion of less than 16-percent zirconium-steam water reaction. Also, the percentage of fuel rods calculated to experience DNB limit (20.0 percent) is met (rods-in DNB case). The sequence of events for the peak RCS pressure/peak cladding temperature case is given in Table 5.1.9-1. This transient trips on a low primary reactor coolant flow trip setpoint, which is assumed to be 87 percent of the normalized flow value. 5.1.9.4 Conclusions The analysis performed has demonstrated that for the locked-rotor event, the RCS pressure remains below 110 percent of the design pressure and the hot spot cladding temperature and oxidation levels remain below the limit values. Therefore, all applicable acceptance criteria are met. In addition, the percentage of fuel rods calculated to experience DNB limit is met.

5.1.10 Loss of External Electrical Load (USAR Section 14.4.9) The Loss-of-External-Electrical-Load event is defined as a complete loss of steam load or a turbine trip from full power without a direct reactor trip. This anticipated transient is analyzed as a turbine trip from full power because it bounds both the Loss-of-External-Electrical-Load event and the Turbine Trip event. The Turbine Trip event is more severe than the total Loss-of-External-Electrical-Load event since it results in a more rapid reduction in steam flow. To address the 422V+ Fuel Transition Program, an evaluation was performed to disposition the changes associated with the new fuel. The results of the evaluation demonstrate that the event is insignificantly impacted by the transition to 422V+ fuel. It is, therefore, concluded that the reactor protection system continues to provide adequate protection such that the minimum calculated DNBR is always greater than the safety analysis limit value and the maximum RCS and MSS pressures are always lower than the respective safety analysis limit values. Thus, all pertinent criteria are met for the loss of external electrical load transient in support of the Prairie Island 422V+ Fuel Transition Program.

5.1.11 Loss of Normal Feedwater (USAR Section 14.4.10) A loss of normal feedwater (from a pipe break, pump failure, or valve malfunction) results in a reduction of the ability of the secondary system to remove the heat generated in the reactor core. If the reactor were not tripped during this accident, core damage could possibly occur from a sudden loss of heat sink. If an alternate supply of feedwater were not supplied to the steam generators, residual heat following reactor trip and RCP heat would cause the primary system water to expand to the point where water relief from the pressurizer would occur. A significant loss of water from the RCS could conceivably lead to core damage. Since the reactor is tripped well before the steam generator heat transfer capability is reduced, the primary system never approaches a condition where the DNBR limit may be violated. To address the 422V+ Fuel Transition Program, an evaluation was performed to disposition the changes associated with the new fuel. The results of the evaluation demonstrate that the event is insignificantly 5-18 impacted by the transition to 422V+ fuel. It is, therefore, concluded that the auxiliary feedwater (AFW) system is capable of removing the stored energy, residual decay heat, and RCP heat such that a more serious plant condition will not occur. Thus, all pertinent criteria are met for the loss of normal feedwater transient in support of the Prairie Island 422V+ Fuel Transition Program.

5.1.12 Loss of All AC Power to the Station Auxiliaries (USAR Section 14.4.11) A complete loss of non-emergency AC power results in the loss of all power to the plant auxiliaries; such as the RCPs, or main feedwater and condensate pumps. The loss of power may be caused by a complete loss of the offsite grid accompanied by a turbine generator trip at the station, or by a loss of the onsite AC distribution system. To address the 422V+ Fuel Transition Program, an evaluation was performed to disposition the changes associated with the new fuel. The results of the evaluation demonstrate that the event is insignificantly impacted by the transition to 422V+ fuel. It is, therefore, concluded that the AFW system is capable of removing the stored energy and residual decay heat such that a more serious plant condition will not occur. Thus, all pertinent criteria are met for the loss of normal feedwater transient in support of the Prairie Island 422V+ Fuel Transition Program.

5.1.13 Rupture of a Steam Pipe (USAR Section 14.5.5) 5.1.13.1 Full-Power Core Response Increased steam flow from the steam generators causes an increase in the heat extraction rate from the RCS, resulting in a reduction of primary coolant temperature and pressure. Because negative moderator temperature and Doppler fuel temperature reactivity coefficients are a characteristic of Westinghouse core designs, the core power will inherently seek a level bounded by the steam load demand, assuming no intervention of control, protection, or engineered safeguards systems. The rate at which the pressusrized water reactor (PWR) approaches equilibrium power with the secondary load is greatest when the reactivity coefficients are the most negative, which corresponds to end-of-life in a fuel cycle. Thus, in the absence of any protective actions, a reactor power level dictated by steam flow rate could be established; however, for analytical purposes, the RPS is credited to terminate the transient consistent with the existing analyses. The analysis examines a range of break sizes to conservatively determine the limiting break size with respect to DNB concerns. To address the 422V+ Fuel Transition Program, an evaluation was performed to disposition the changes associated with the new fuel. The results of the evaluation concluded that the minimum DNBR remains above the safety analysis DNBR limit throughout the course of the transient and the peak linear heat generation rate does not exceed a value that would cause fuel centerline melt. Therefore, it is concluded that fuel and cladding damage will not occur. Thus, all pertinent criteria are met for the steam line break - full-power core response transient in support of the Prairie Island 422V+ Fuel Transition Program.

5-19 5.1.13.2 Zero-Power Core Response A steam line break transient while at hot zero-power conditions would result in an uncontrolled increase in steam flow release from the steam generators, with the flow decreasing as the steam pressure drops. This steam flow release increases the heat removal from the RCS, which decreases the RCS temperature and pressure. With the existence of a negative moderator temperature coefficient (MTC), the RCS cooldown results in a positive reactivity insertion, and consequently a reduction of the core shutdown margin. If the most reactive RCCA is assumed stuck in its fully withdrawn position after reactor trip, the possibility is increased that the core will become critical and return to power. A return to power following a steam line break is a concern with the high-power peaking factors that may exist when the most reactive RCCA is stuck in its fully withdrawn position. Following a steam line break, the core is ultimately shut down by the boric acid, which is injected into the RCS by the emergency core cooling system (safety injection). The analysis examines a double-ended rupture (DER) of a steam line while the reactor is at hot zero-power conditions. To address the 422V+ Fuel Transition Program, an evaluation was performed to disposition the changes associated with the new fuel. The results of the evaluation concluded that the minimum DNBR remains above the safety analysis DNBR limit remains below the applicable safety analysis limit throughout the course of the transient and the peak linear heat generation rate does not exceed a value that would cause centerline melt. Therefore, it is concluded that fuel and cladding damage will not occur. Thus, all pertinent criteria are met for the steam line break - zero-power core response transient in support of the Prairie Island 422V+ Fuel Transition Program.

5.1.14 RCCA Ejection (USAR Section 14.5.6) 5.1.14.1 Accident Description This accident is the result of the extremely unlikely mechanical failure of a control rod drive mechanism pressure housing such that the RCS pressure would eject the RCCA and drive shaft. The consequences of this mechanical failure, in addition to being a minor LOCA, may also be a rapid reactivity insertion together with an adverse core power distribution, possibly leading to localized fuel rod damage. Certain features in Westinghouse PWRs are intended to preclude the possibility of a rod ejection accident, or to limit the consequences if the accident were to occur. These include a sound, conservative mechanical design of the rod housings, along with a thorough quality control (testing) program during assembly, and a nuclear design that lessens the potential ejection worth of control rod assemblies and minimizes the number of assemblies inserted at high power levels. Even if a rupture of the control rod mechanism housing is postulated, the operation of a chemical shim plant is such that the severity of an ejected rod is inherently limited. In general, the reactor is operated with control rods inserted only far enough to permit load follow. Reactivity changes caused by core depletion and xenon transients are compensated by boron changes. Further, the location and groupings of control rod banks are selected during the core nuclear design to lessen the severity of an ejected control rod assembly. Therefore, should an RCCA be ejected from the reactor vessel during normal operation, there would probably be no reactivity excursion since most of the control rods are fully withdrawn from the core, or a minor reactivity excursion if an inserted RCCA is ejected from its normal position.

5-20 However, it may occasionally be desirable to operate with larger control rod insertions. For this reason, rod insertion limits are defined in the Technical Specifications as a function of power level. Operation with the RCCAs above this limit guarantees adequate shutdown capability and acceptable power distribution. The position of all RCCAs is continuously indicated in the control room. An alarm will occur if a bank of RCCAs approaches its insertion limit or if one RCCA deviates from its bank. There are low and low-low level insertion monitors with visual and audio signals. Operating instructions require boration when receiving either alarm. If an RCCA ejection accident were to occur, a fuel rod thermal transient that could cause DNB may occur together with limited fuel damage. The amount of fuel damage that can result from such an accident will be governed mainly by the worth of the ejected RCCA and the power distribution attained with the remaining control rod pattern. The transient is limited by the Doppler reactivity effects of the increase in fuel temperature and is terminated by a reactor trip that is actuated by neutron flux signals. The reactor trip will occur before conditions are reached that can result in damage to the reactor coolant pressure boundary or significant disturbances in the core, its support structures, or other reactor pressure vessel internals that would impair the capability to cool the core.

The neutron flux response to a continuous reactivity insertion is characterized by a very fast flux increase terminated by the reactivity feedback effect of the negative Doppler coefficient. This self-limitation of the power burst is of primary importance since it limits the power to a tolerable level during the delay time for protective action. Should an RCCA ejection accident occur, the following automatic features of the reactor protection system are available to terminate the transient: The source-range high neutron flux reactor trip is actuated when either of two independent source-range channels indicates a neutron flux level above a preselected manually adjustable setpoint. This trip function may be manually bypassed when either intermediate-range flux channel indicates a flux level above a specified level. It is automatically reinstated when both intermediate-range channe ls indicate a flux level below a specified level. The intermediate-range high neutron flux reactor trip is actuated when either of two independent intermediate-range channels indicates a flux level above a preselected manually adjustable setpoint. This trip function may be manually bypassed when two-out-of-four power-range channels give readings above approximately 10 percent of full-power and is automatically reinstated when three-out-of-four channels indicate a power below this value.

5-21 The power-range high neutron flux reactor trip (low setting) is actuated when two-out-of-four power-range channels indicate a power level above a preselected manually adjustable setpoint (allowable value, 40 percent power). This trip function may be manually bypassed when two-out-of-four power-range channels indicate a power level above approximately 10 percent of full-power and is automatically reinstated when three-out-of-four channels indicate a power level below this value.

- The power-range high neutron flux reactor trip (high setting) is actuated when two-out-of-four power-range cha nnels indicate a power level above a preset setpoint (allowable value, 110 percent power). This trip function is always active when the reactor is at power.

- The high nuclear flux rate reactor trip is actuated when the positive rate of change of neutron flux on two-out-of-four nuclear power-range channels indicates a rate above the preset setpoint. This trip function is always active. The ultimate acceptance criteria for this event is that any consequential damage to either the core or the RCS must not prevent long-term core cooling, and that any offsite dose consequences must be within the guidelines of 10 CFR 100. To demonstrate compliance with these requirements, it is sufficient to show that the RCS pressure boundary remains intact, and that no fuel dispersal in the coolant, gross lattice distortions, or severe shock waves will occur in the core. Therefore, the following acceptance criteria are applied to the RCCA ejection accident: 1. Maximum average fuel pellet enthalpy at the hot spot must remain below 200 cal/g (360 Btu/lbm). 2. Peak RCS pressure must remain below that which would cause the stresses in the RCS to exceed the faulted condition stress limits. 3. Maximum fuel melting must be limited to the innermost 10 percent of the fuel pellet at the hot spot, independent of the above pellet enthalpy limit.

Since this event is sensitive to the fuel characteristics, the event has been completely re-analyzed in support of the transition to Westinghouse 422V+ fuel. 5.1.14.2 Method of Analysis The calculation of the RCCA ejection transient is performed in two stages: a neutron kinetic analysis and a hot spot fuel heat transfer analysis. The spatial neutron kinetics code TWINKLE (Reference 5) is used in a one-dimensional (1-D) axial kinetics model to calculate the core nuclear power including the various total core feedback effects, that is, Doppler reactivity and moderator reactivity. The average core nuclear power is multiplied by the post-ejection hot channel factor, and the fuel enthalpy and temperature transients at the hot spot are calculated with the detailed fuel and cladding transient heat transfer computer code, FACTRAN (Reference 1). The power distributi on calculated without feedback is pessimistically assumed to persist throughout the transient. Additional details of the methodology are provided in WCAP-7588 (Reference 8).

5-22 The overpressurization of the RCS and number of rods in DNB, as a result of a postulated ejected rod, have both been analyzed on a generic basis for Westinghouse PWRs as detailed in Reference 8. If the safety limits for fuel damage are not exceeded, there is little likelihood of fuel dispersal into the coolant or a sudden pressure increase from thermal-to-kinetic energy conversion. The pressure surge for this analysis can, therefore, be calculated on the basis of conventional heat transfer from the fuel and prompt heat generation in the coolant. A detailed calculation of the pressure surge for an ejection worth of one dollar at beginning of life (BOL), hot full-power, indicates that the peak pressure does not exceed that which would cause stresses in the RCS to exceed their faulted condition stress limits. Since the severity of the Prairie Island analysis does not exceed this worst-case analysis, the RCCA ejecti on accident will not result in an excessive pressure rise or further damage to the RCS. Reference 8 also documents a detailed multi-channel thermal-hydraulics code calculation, which demonstrates an upper limit to the number of rods in DNB for the RCCA ejection accident as 10 percent. Since the severity of the Prairie Island analysis does not exceed this worst-case analysis, the maximum number of rods in DNB following an RCCA ejection will be less than 10 percent, which is well within the value used in the radiological dose evaluation. The most limiting break size resulting from an RCCA ejection will not be sufficient to uncover the core or cause DNB at any later time. Since the maximum number of fuel rods experiencing DNB is limited to 10 percent, the fission product release will not exceed that associated with the guidelines of 10 CFR 100. In calculating the nuclear power and hot spot fuel rod transients following RCCA ejection, the following conservative assumptions are made:

1. The STDP (maximum uncertainties in initial conditions) is used for the RCCA ejection analysis. The analysis assumes uncertainties of +4.0F in nominal vessel T avg, and -60 psi in nominal pressurizer pressure. A reactor power level of 1,683 MWt was modeled, consistent with the maximum reactor power including all applicable uncertainties. 2. A minimum value for the delayed neutron fraction for beginning of cycle (BOC) and EOC conditions is assumed which increases the rate at which the nuclear power increases following an RCCA ejection accident. 3. A minimum value of the Doppler power defect, which conservatively results in the maximum amount of energy deposited in the fuel following an RCCA ejection accident, is assumed. A minimum value of the moderator feedback is also assumed. A positive moderator temperature coefficient is assumed for the BOC, zero-power case. 4. Maximum values of ejected RCCA worth and post-ejection total hot channel factors are assumed for all cases considered. These parameters are ca lculated using standard nuclear design codes for the maximum allowed bank insertion at a given power level, as determined by the rod insertion limits. No credit is taken for the flux flattening effects of reactivity feedback. 5. The start of rod motion occurs 0.45 seconds after the high neutron flux trip point is reached.

5-23 The analysis is performed to address the transition to Westinghouse 422V+ fuel (UO 2 and up to 8 w/o gadolinium-doped UO

2) and a maximum loop-to-loop steam generator tube plugging imbalance of 10 percent. 5.1.14.3 Results Figures 5.1.14-1 through 5.1.14-12 are representative nuclear power and hot spot fuel rod thermal transients following an RCCA ejection accident. The transient results of the analysis are summarized in Tables 5.1.14-1 and 5.1.14-2. A time sequence of events is provided in Table 5.1.14-3. For all cases, the maximum fuel pellet enthalpy remained below 200 cal/g. For the hot full-power cases with gadolinium fuel, the peak hot spot fuel centerline temperature reached the fuel melting temperature; however, melting was restricted to less than 10 percent of the pellet. For the hot zero-power cases, no fuel melting was predicted. The UO 2 cases are bounding for all fuel types, including gadolinium-doped fuel. 5.1.14.4 Conclusions Even on the most pessimistic basis, the analyses indicate that the fuel and cladding limits are not exceeded. It is concluded that there is no danger of sudden fuel dispersal into the coolant. Since the pressure does not exceed that which would cause stresses to exceed the faulted condition stress limits, it is concluded that there is no danger of further consequential damage to the primary coolant system. The amount of fission products released as a result of the assumed failure of fuel rods entering into DNB will not exceed the guidelines of 10 CFR 100.

5.1.15 ATWS (USAR Section 14.8) 10 CFR 50.62(b) requires that Westinghouse PWRs implement the anticipated transient without scram (ATWS) mitigating system actuation circuitry (AMSAC). AMSAC is installed at Prairie Island, so the requirements of 10 CFR 50.62(b) are satisfied. A diverse scram system (DSS) is installed at Prairie Island as a supplement to AMSAC. The Nuclear Regulatory Commission (NRC) has approved the implementation of AMSAC and the DSS at Prairie Island as documented in Section 14.8 of the Prairie Island USAR. The AMSAC and DSS will be maintained and operated following implementation of the 422V+ Fuel Transition Program consistent with their design bases and as approved by the NRC. The analysis of the DSS transients are performed to demonstrate that the DNBR remains greater than the analytical limit of 1.17 and that the RCS pressure remains greater than the analytical limit of 3,200 psig. To address the 422V+ Fuel Transition Program, an evaluation was performed to disposition the changes associated with the new fuel. With respect to DNB concerns, it is concluded that the increase in the nominal DNBR associated with the fuel transition outweighs the relatively small changes associated with the remainder of the fuel-related parameters. In addition, key core physics parameters assumed in the analyses will be checked for each specific cycle as part of the normal RSAC. With respect to RCS pressure concerns, the relatively small changes associated with the 422V+ fuel do not have a significant impact on the pressure transient associated with each of the DSS events. Thus, all pertinent criteria are met for the DSS events in support of the Prairie Island 422V+ Fuel Transition Program.

5-24 Table 5.1.1-1 Assumptions and Results - Uncontrolled RCCA Withdrawal from a Subcritical Condition Initial Power Level, % 0 Reactivity Insertion Rate, pcm/sec 75 Delayed Neutron Fraction 0.0072 Doppler Power Defect, pcm 1,100 Trip Reactivity, % k 1.0 Hot Channel Factor 6.64 Number of RCPs Operating 1 Results Calculated Value Limit Peak Fuel Centerline Temperature, °F 2,402 4,746 (1) Peak Fuel Average Temperature, °F 1,930 4,746 Below the First Mixing Vane Grid (W-3 Correlation): Minimum DNBR (Thimble cell) 2.081 1.428 Minimum DNBR (Typical cell) 1.866 1.428 Above the First Mixing Vane Grid (WRB-1 Correlation): Minimum DNBR (Thimble Cell) 2.206 1.285 Minimum DNBR (Typical Cell) 2.202 1.285 Note: 1. Limit is for fuel with 8 w/o Gadolinia 5-25 Table 5.1.1-2 Sequence of Events - Uncontrolled RCCA Withdrawal from a Subcritical Condition Event Time (seconds)

(1) Initiation of Uncontrolled RCCA Bank Withdrawal 0 Power-Range High Neutron Flux Low Setpoint Reached 10.0 Peak Nuclear Power Occurs 10.1 Rod Motion Begins 10.45 Peak Heat Flux Occurs 12.3 Minimum DNBR Occurs 12.3 Peak Cladding Temperature Occurs 12.7 Peak Fuel Average Temperature Occurs 12.9 Peak Fuel Centerline Temperature Occurs 14.1 Note: 1. The times reported are for the without Gadolinia case.

Table 5.1.8-1 Sequence of Events - Partial Loss of Reactor Coolant Flow Event Time (seconds) One Operating RCP Loses Power and Begins Coasting Down 0.0 Low Flow Reactor Trip Setpoint is Reached 1.58 Rods Begin to Drop 2.78 Minimum DNBR Occurs 4.2 Table 5.1.8-2 Sequence of Events - Complete Loss of Reactor Coolant Flow Event Time (seconds) All Operating RCPs Lose Power and Coastdown Begins 0.0 Low Flow Reactor Trip Setpoint is Reached 1.78 Rods Begin to Drop 2.98 Minimum DNBR Occurs 4.65 5-26 Table 5.1.9-1 Sequence of Events - Reactor Coolant Pump Locked Rotor Event Time (seconds) Rotor on One Pump Locks 0.00 Low Flow Reactor Trip Setpoint Reached 0.07 Rods Begin to Drop 1.27 Maximum Cladding Temperature Occurs 3.42 Maximum RCS Pressure Occurs 5.00 Table 5.1.14-1 Assumptions and Results-RCCA Ejection Beginning of Cycle Full-Power Zero-Power Initial Power Level, % 100 (1) 0 Ejected RCCA Worth, % k 0.380 0.770 Delayed Neutron Fraction 0.0049 0.0049 Doppler Power Defect, % k 1.000 1.000 Feedback Reactivity Weighting 1.139 2.008 Trip Reactivity, % k 4.0 1.0 F Q Before Ejection 2.5 N/A F Q After Ejection 4.2 11.0 Number of Reactor Coolant Pumps (RCPs) Operating 2 1 End of Cycle Full-Power Zero-Power Initial Power Level, % 100 (1) 0 Ejected RCCA Worth, % k 0.30 0.954 Delayed Neutron Fraction 0.0047 0.0047 Doppler Power Defect, % k 0.980 0.980 Feedback Reactivity Weighting 1.316 2.755 Trip Reactivity, % k 4.0 1.0 F Q Before Ejection 2.5 N/A F Q After Ejection 5.69 18.42 Number of RCPs Operating 2 1 Note: 1. The full-power cases considered a reactor power of 1,683 MWt, which includes all applicable uncertainties.

5-27 Table 5.1.14-2 Prairie Island Rod Ejection Results with and Without Gadolinium BOL-HZP BOL-HFP EOL-HZP EOL-HFP Parameter UO2 8 w/o Gad (1) UO 2 8 w/o Gad (1) UO2 8 w/o Gad (1) UO 2 8 w/o Gad (1) Max. Fuel Centerline Temp. ( F) 3,790 3,829 4,882 4,929 3,861 3,934 4,738 4,826 Max. Fuel Average Temp. (F) 3,299 3,339 3,646 3,556 3,429 3,483 3,501 3,433 Max. Fuel Stored Energy (Btu/lb) 249.6 253.1 280.9 272.6 261.1 265.9 267.5 261.5 Fuel Melt (%) 0.0 0.0 0.0 1.97 0.0 0.0 0.00 1.65 Max. Cladding Average Temp. (F) 2,452 2,509 1,974 1,723 2,629 2,710 1,893 1,669 Reacted zirc (%) 1.79 1.91 0.37 0.16 2.56 2.82 0.29 0.14 Note: 1. It should be noted that the HZP gadolinium-doped UO 2 cases did not credit the available power suppression since the reduction was not needed to provide acceptable results. The BOL and EOL cases credit 20% power suppression (0.80 power suppression factor).

5-28 Table 5.1.14-3 Sequence of Events - RCCA Ejection Beginning of Cycle - Hot Zero-Power Time (seconds) RCCA Ejection Occurs 0.000 High Neutron Flux Setpoint (Low Setting) is Reached 0.211 Peak Nuclear Power Occurs 0.252 Rods Begin to Fall into the Core 0.661 Peak Cladding Average Temperature Occurs 2.191 (1) Peak Fuel Average Temperature Occurs 2.432 (1) Beginning of Cycle - Hot Full-Power Time (seconds) RCCA Ejection Occurs 0.000 High Neutron Flux Setpoint (High Setting) is Reached 0.030 Peak Nuclear Power Occurs 0.135 Rods Begin to Fall into the Core 0.480 Peak Fuel Average Temperature Occurs 1.974 (1) Peak Cladding Average Temperature Occurs 2.085 (1) End of Cycle - Hot Zero-Power Time (seconds) RCCA Ejection Occurs 0.000 High Neutron Flux Setpoint (Low Setting) is Reached 0.156 Peak Nuclear Power Occurs 0.183 Rods Begin to Fall into the Core 0.606 Peak Cladding Average Temperature Occurs 1.574 (1) Peak Fuel Average Temperature Occurs 1.865 (1) End of Cycle - Hot Full-Power Time (seconds) RCCA Ejection Occurs 0.000 High Neutron Flux Setpoint (High Setting) is Reached 0.035 Peak Nuclear Power Occurs 0.130 Rods Begin to Fall into the Core 0.485 Peak Fuel Average Temperature Occurs 2.025 (1) Peak Cladding Average Temperature Occurs 2.154 (1) Note: 1. The times presented for peak fuel average temperature and peak cladding average temperature are from the UO 2 fuel case. All other times presented are the same for the UO 2 and Gad cases.

5-29 Figure 5.1.1-1. Uncontrolled RCCA Withdrawal from a Subcritical Condition - Reactor Power Versus Time 5-30 Figure 5.1.1-2. Uncontrolled RCCA Withdrawal from a Subcritical Condition - Heat Flux Versus Time 5-31 Figure 5.1.1-3. Uncontrolled RCCA Withdrawal from a Subcritical Condition - Hot Spot Fuel Centerline Temperature Versus Time 5-32 Figure 5.1.1-4. Uncontrolled RCCA Withdrawal from a Subcritical Condition - Hot Spot Fuel Average Temperature Versus Time 5-33 Figure 5.1.1-5. Uncontrolled RCCA Withdrawal from a Subcritical Condition - Hot Spot Cladding Temperature Versus Time 5-34 Figure 5.1.8-1. Total Core Inlet Flow Versus Time - Partial Loss of Flow (PLOF), One Pump Coasting Down 5-35 Figure 5.1.8-2. RCS Faulted Loop Flow Versus Time - PLOF, One Pump Coasting Down 5-36 Figure 5.1.8-3. Nuclear Power Versus Time - PLOF, One Pump Coasting Down 5-37 Figure 5.1.8-4. Core Average Heat Flux Versus Time - PLOF, One Pump Coasting Down 5-38 Figure 5.1.8-5. Pressurizer Pressure Versus Time - PLOF, One Pump Coasting Down laap 1 bat Leq Loop 1 C~ld leg Figure 5.1.8-6. RCS Faulted Loop Temperature Versus Time - PLOF, One Pump Coasting Down 5-40 Figure 5.1.8-7. Hot Channel Heat Flux Versus Time - PLOF, One Pump Coasting Down 5-41 Figure 5.1.8-8. DNBR Versus Time - PLOF, One Pump Coasting Down 5-42 Figure 5.1.8-9. Total Core Inlet Flow Versus Time - Complete Loss of Flow (CLOF) - Two Pumps Coasting Down 5-43 Figure 5.1.8-10. RCS Loop Flow Versus Time - CLOF - Two Pumps Coasting Down 5-44 Figure 5.1.8-11. Nuclear Power Versus Time - CLOF - Two Pumps Coasting Down 5-45 Figure 5.1.8-12. Core Average Heat Flux Versus Time - CLOF - Two Pumps Coasting Down 5-46 Figure 5.1.8-13. Pressurizer Pressure Versus Time - CLOF - Two Pumps Coasting Down Loop 2 Ho'i ie g ----- Loop I Ccid leg 5 e 12 Time (sec) Figure 5.1.8-14. RCS Faulted Loop Temperature Versus Time - CLOF - Two Pumps Coasting Down 5-48 Figure 5.1.8-15. Hot Channel Heat Flux Versus Time - CLOF - Two Pumps Coasting Down 5-49 Figure 5.1.8-16. DNBR Versus Time - CLOF - Two Pumps Coasting Down 5-50 Figure 5.1.9-1. Total Core Inlet Flow Versus Time - Locked Rotor/Shaft Break - RCS Pressure/Peak Cladding Temperature Case 5-51 Figure 5.1.9-2 RCS Loop Flow Versus Time - Locked Rotor/Shaft Break - RCS Pressure/Peak Cladding Temperature Case 5-52 Figure 5.1.9-3. Nuclear Power Versus Time - Locked Rotor/Shaft Break - RCS Pressure/Peak Cladding Temperature Case 5-53 Figure 5.1.9-4. Core Average Heat Flux Versus Time - Locked Rotor/Shaft Break - RCS Pressure/Peak Cladding Temperature Case 5-54 Figure 5.1.9-5. Pressurizer Pressure Versus Time - Locked Rotor/Shaft Break - RCS Pressure/Peak Cladding Temperature Case 5-55 Figure 5.1.9-6. Vessel Lower Plenum Pressure Versus Time - Locked Rotor/Shaft Break - RCS Pressure/Peak Cladding Temperature Case 5-56 Figure 5.1.9-7. RCS Loop Temperature Versus Time - Locked Rotor/Shaft Break - RCS Pressure/Peak Cladding Temperature Case 5-57 Figure 5.1.9-8. Hot Spot Cladding Inner Temperature Versus Time - Locked Rotor/Shaft Break - RCS Pressure/Peak Cladding Temperature Case 5-58 Figure 5.1.14-1. RCCA Ejection - BOC Full-Power Reactor Power Versus Time 5-59 Figure 5.1.14-2. RCCA Ejection - BOC Full-Power Fuel and Cladding Temperatures Versus Time for U0 2 Fuel Case 5-60 Figure 5.1.14-3. RCCA Ejection - BOC Full-Power Fuel and Cladding Temperatures Versus Time for Gadolinium Fuel Case 5-61 Figure 5.1.14-4. RCCA Ejection - BOC Zero-Power Reactor Power Versus Time 5-62 Figure 5.1.14-5. RCCA Ejection - BOC Zero-Power Fuel and Cladding Temperatures Versus Time for U0 2 Fuel Case 5-63 Figure 5.1.14-6. RCCA Ejection - BOC Zero-Power Fuel and Cladding Temperatures Versus Time for Gadolinium Fuel Case 5-64 Figure 5.1.14-7. RCCA Ejection - EOC Full-Power Reactor Power Versus Time 5-65 Figure 5.1.14-8. RCCA Ejection - EOC Full-Power Fuel and Cladding Temperatures Versus Time for U0 2 Fuel Case 5-66 Figure 5.1.14-9. RCCA Ejection - EOC Full-Power Fuel and Cladding Temperatures Versus Time for Gadolinium Fuel Case 5-67 Figure 5.1.14-10. RCCA Ejection - EOC Zero-Power Reactor Power Versus Time 5-68 Figure 5.1.14-11. RCCA Ejection - EOC Zero-Power Fuel and Cladding Temperatures Versus Time 5-69 Figure 5.1.14-12. RCCA Ejection - EOC Zero-Power Fuel and Cladding Temperatures Versus Time for Gadolinium Fuel Case 5-70 5.2 LOSS-OF-COOLANT ACCIDENTS 5.2.1 Large-Break Best-Estimate LOCA Prairie Island recently implemented the best-estimate large-break (LBLOCA) methodology using the automated statistical treatment of uncertainty method (ASTRUM) statistical applications of uncertainties for both Units 1&2. This was approved by the NRC as documented in Reference 9. This LBLOCA analysis was completely re-analyzed for the transition to 422 Vantage+ (V+) fuel due to the impact of the larger fuel rods on the core water volume and resultant fuel temperatures. All outstanding 50.46 assessments were addressed in the re-analysis. No other significant input changes were incorporated into the re-analysis. The same methodology was used as described in Reference 9. The transition from Westinghouse 400V+ to Westinghouse 422V+ fuel has been evaluated for the effects of hydraulic mismatch and differences in fuel designs for both Prairie Island Units 1&2. Transition core peak cladding temperature (PCT) results for both units showed that the ASTRUM analysis for 422V+ bounds the transition core cycles and the full core 400V+ analysis, which is the current analysis of record. Table 5.2.1-1 and Table 5.2.1-2 present the LBLOCA analysis results and the conditions analyzed for Unit 1, and Table 5.2.1-3 and Table 5.2.1-4 present the LBLOCA analysis results and the conditions analyzed for Unit 2. These tables show that Prairie Island Units 1&2 meet the 10 CFR 50.46 acceptance criteria and the best-estimate upper plenum injection (UPI) Safety Evaluation Report (SER) requirements. Prevention of Return to Recriticality During Reflood As part of the transition from Westinghouse 400V+ to Westinghouse 422V+, the Safety Analysis Checklist (SAC) limits were reconfirmed. As such, a check was performed to show that the core remains subcritical in the short term following an LBLOCA. Depending on the boron concentration of the water reflooding the core, the reactivity in this period coul d increase the core power so as to challenge the 10 CFR 50.46 limits. The core engineering calculations indicated that the current minimum accumulator boron concentration Technical Specification Limit of 1,900 ppm is not sufficient to assure a prevention of a recriticality. Further calculations, using representative loading patterns, showed an accumulator boron concentration of 2,300 ppm was sufficient to prevent the return to criticality. These calculations provide reasonable assurance that this accumulator minimum boron concentration will prevent a recriticality event as described above. Therefore, a change to the minimum accumulator boron concentration Technical Specification limit is required to ensure the plant operates in a manner that provides acceptable levels of protection for health and safety of the public. This parameter will be checked for each cycle specific reload as part of the normal core design and Reload Safety Evaluation processes.

5.2.2 Small-Break LOCA 5.2.2.1 Introduction The small-break LOCA (SBLOCA) analysis of record for Prairie Island Units 1&2 was completed using the 1985 Westinghouse SBLOCA Evaluation Model (EM) with NOTRUMP (NOTRUMP-EM; References 10 through 12). The SBLOCA analysis was completed to address the upgrade to 422V+ fuel and incorporated all prior 10 CFR 50.46 assessments, except the planned plant modification evaluation of 5-71 two reconstituted rods for Unit 1, primarily through the use of the latest code ve rsions and selection of input values. Also addressed in the analysis are loop seal restriction assumptions consistent with Westinghouse methodology (Reference 13) and other minor items identified through discussions between Nuclear Management Company (NMC) and Westinghouse.

The following sections provide an overview of the SBLOCA assumptions and initial conditions, analysis methodology, acceptance criteria, results, and conclusions. 5.2.2.2 Assumptions and Initial Conditions The SBLOCA methodology using NOTRUMP-EM was developed in accordance with the requirements of 10 CFR 50 Appendix K (Reference 14). This regulation was designed to produce a conservative prediction of the analysis results and includes various conservative modeling requirements such as the decay heat model (1971 ANS Infinite + 20%) and the zirconium-water reaction model (Baker-Just). For the SBLOCA analysis, loss of offsite power (LOOP) is assumed, which results in the limiting single failure assumption of the loss of one emergency diesel generator (EDG) and subsequent loss of one train of pumped emergency core cooling system (ECCS) capacity. The SBLOCA analysis assumes that reactor trip occurs coincident with the LOOP, which results in the following: (1) RCP trip and coastdown and (2) steam dump system being inoperable. Additional input assumptions and initial conditions for the SBLOCA analysis are found in Tables 5.2.2-1 through 5.2.2-5 for Units 1&2. 5.2.2.3 Description of Analysis The requirements for an acceptable ECCS evaluation model are presented in Appendix K of 10 CFR 50 (Reference 14). The NOTRUMP-EM (References 10 through 12) is used to determine the RCS response to design basis SBLOCAs. The NOTRUMP-EM consists of the NOTRUMP and LOCTA-IV computer codes. The NOTRUMP code is employed to calculate the transient depressurization of the RCS as well as the mass and energy release of the fluid flow through the break. Among the features of the NOTRUMP code are: calculation of the thermal non-equilibrium in all fluid volumes, flow regime-dependent drift flux calculations with counter-current flooding limitations, mixture level tracking logic in multiple-stacked fluid nodes, regime-dependent drift flux calculations in multiple-stacked fluid nodes, and regime-dependent heat transfer correlations. These features provide NOTRUMP with the capability to accurately calculate the mass and energy distribution throughout the RCS during the course of a SBLOCA. The RCS model is nodalized into volumes interconnected by flow paths. The broken loop and intact loop are modeled explicitly for both Units 1&2 because of their two-loop configuration. Transient behavior of the system is determined from governing conservation equations of mass, energy, and momentum. The multi-node capability of the program enables explicit, detailed spatial representation of various system components which, among other capabilities, enables a calculation of the behavior of the loop seals during an SBLOCA. The reactor core is represented as heated control volumes with associated phase separation models to permit transient mixture height calculations.

5-72 Fuel cladding thermal analyses are performed with a version of the LOCTA-IV (Reference 11) using the NOTRUMP calculated core pressure, fuel rod power history, core flow, and mixture heights as boundary conditions. The LOCTA-IV code models the hot rod and average hot assembly rod, assuming a conservative power distribution that is skewed to the top of the core. Figure 5.2.2-1 illustrates the core interface for the NOTRUMP-EM. This analysis was performed assuming a full core of 14x14 422V+ fuel assemblies; however, it is applicable for mixed cores consisting of both 422V+ and 400V+. The only mechanism available to cause significant differences in overall RCS thermal-hydraulic response would be flow redistribution due to fuel assembly hydraulic resistance mismatch between the 400V+ and 422V+ assemblies. Even for larger SBLOCAs, the thermal-hydraulic response is quasi-one dimensional, with st ratified flow conditions eventually existing throughout the core and the RCS. Under these conditions, core flow rate is relatively low which provides enough time to maintain flow equilibrium between fuel assemblies (that is, crossflow is not a factor). Therefore, this SBLOCA analysis is applicable for the core configuration with 400V+ fuel and the mixed core at Prairie Island Units 1&2. The loop seal restriction is removed consistent with standard Westinghouse NOTRUMP-EM analyses (discussed in NSBU-NRC-00-5972 (Reference 13)). In prior analyses, Prairie Island Units 1&2 kept the loop seal restriction on for all break sizes. This is considered to be an overly conservative, unrealistic application of the EM and is not consistent with the standard approved methodology. The standard methodology is to remove the non-faulted loop restriction for larger cold leg break sizes as discussed in detail in Reference 11.

For smaller break sizes, only the faulted loop seal is permitted to clear. This has always been the standard Westinghouse practice with the NOTRUMP-EM and has been accepted many times, more recently for the Beaver Valley and Ginna Extended Power Uprate Programs (References 15 and 16, respectively). The Prairie Island Units 1&2 SBLOCA analysis evaluates cold leg breaks. The effects of break location on SBLOCA analyses have been generically evaluated as part of the application of the NOTRUMP-EM (Reference 17), where it was concluded from the break location study that a break in the RCS cold leg was limiting. Additionally, the effects of break orientation were considered during the evaluation of safety injection in the broken loop and application of the COSI Condensation Model (Reference 12). This work concluded that the standard break orientation was limiting with respect to PCT. There have been past concerns cited by the NRC staff that top oriented breaks (independent of size) would lead to longer term core uncovery should the RCP suction piping refill and upset the vapor vent path from the core. The Westinghouse position on this, which was first established in 1997, is that this will not occur. In a simplified sense, the inventory in the core and vessel is acting as a manometer at this point. However, in order to establish this hypothesized scenario in a long-term sense, a delicate set of conditions with regard to vent paths in the RCS which rely on lower steaming rates would need to occur. In the short term (that is, prior to reaching UPI cut-in pressure), this would be extremely unlikely because of the high boil-off rates resulting from near-term decay heat. As such, the loop seal plugging and purging would be an ongoing process that would lead to many mixture level oscillations, but no long-term uncovery periods. This has been shown to be the case in facility tests such as ROSA as well as NOTRUMP simulations.

5-73 In addition, it should be noted that the NOTRUMP-EM methodology does not consider the presence of the gaps between the core barrel upper plenum nozzles and the vessel. Although small, when considered around the entire circumference of the hot leg nozzles, the resulting flow area is not trivial. These gaps will be present since the hypothesized scenario dictat es a condition where this area of the vessel will be filled with a saturated mixture and/or subcooled liquid such that temperature expansion concerns do not exist. In the extended time frame assumed for the scenario to occur, it is quite likely that this flow area, when coupled with the vessel upper bypass flow path, will provide enough vapor relief path to bypass the loop seal(s) entirely and eliminate the effects of loop seal plugging and re-plugging on core uncovery. Also, while not modeled in the analysis, all plants use a cooldown depressurization method using the steam generators to get to cold shutdown conditions. In this process, excess vapor in the RCS is condensed in the steam generator tubes. Thus, the potential for a differential pressure between the core and downcomer is relieved. 5.2.2.4 Acceptance Criteria The acceptance criteria for the SBLOCA analysis are specified in 10 CFR 50.46 (Reference 18) and are summarized as follows: 1. The calculated maximum fuel element cladding temperature shall not exceed 2,200F. 2. The calculated total oxidation of the cladding shall nowhere exceed 0.17 times the total cladding thickness before oxidation. 3. The calculated total amount of hydrogen generated from the chemical reaction of the cladding with water or steam shall not exceed 0.01 times the hypothetical amount that would be generated if all of the metal in the cladding cylinders surrounding the fuel, excluding the cladding surrounding the plenum volume, were to react. 4. Calculated changes in core geometry shall be such that the core remains amenable to cooling. 5. After any calculated successful initial operation of the ECCS, the calculated core temperature shall be maintained at an acceptably low value and decay heat shall be removed for the extended period of time required by the long-lived radioactivity remaining in the core. Criteria 1 through 3 are addressed explicitly by the NOTRUMP-EM. Criterion 4 is addressed implicitly via the blockage model used when LOCTA-IV predicts rod burst and Criterion 5 is addressed in part and implicitly by the termination criteria for the NOTRUMP calculation to account for the switchover from injection to cold leg recirculation (as discussed in subsection 5.2.4). 5.2.2.5 Results and Discussion The sequence of events and beginning-of-life fuel rod heatup results for the SBLOCA analysis is presented in Tables 5.2.2-6 and 5.2.2-7 for Unit 1 and Tables 5.2.2-8 and 5.2.2-9 for Unit 2. A break spectrum consisting of 1.5-inch, 2-inch, 3-inch, 4-inch, 6-inch, and 8-inch cold leg breaks along with a 10.126-inch accumulator line break was performed for both Units 1&2 to determine the limiting PCT small-break transient. Per Reference 19, this break spectrum is considered acceptable because the 5-74 SBLOCA PCTs are significantly less than 1700°F and significantly less than the LBLOCA PCTs for both Units 1&2. The LBLOCA analysis remains the limiting LOCA transient. The limiting case for Unit 1 is the 3-inch break with a PCT of 959F and the limiting case for Unit 2 is the 2-inch break with a PCT of 965F. The maximum local oxidation (MLO) is 0.01 percent for both the Unit 1 and Unit 2 limiting cases. The total local oxidation (pre-transient plus transient oxidation) will remain below the 10 CFR 50.46 limit of 17 percent at all times in the life of the fuel. The core-wide hydrogen generation remains well below the 10 CFR 50.46 acceptance limit of 1 percent, and the core geometry remains amenable to cooling. The following parameters have been presented for both the 3-inch break (Unit 1) and 2-inch break (Unit 2) as Figures 5.2.2-3 through 5.2.2-8 and Figures 5.2.2-9 through 5.2.2-14 respectively: RCS pressure Core mixture level Core exit vapor temperature Broken loop and intact loop pumped safety injection (SI) flow rates Total break flow and total pumped SI flow rates Cladding temperature at PCT elevation and local oxidation at MLO elevation The non-limiting transient parameters (summarized below) are presented in Figures 5.2.2-15 through 5.2.2-28 for Unit 1 and Figures 5.2.2-29 through 5.2.2-42 for Unit 2. The 1.5-inch, 6-inch, 8-inch, and 10.126-inch break cases for both Unit 1 and Unit 2 resulted in either minimal or no core uncovery; therefore, fuel rod heatup calculations were not performed for these break sizes. RCS pressure Core mixture level Cladding temperature at PCT elevation and local oxidation at MLO elevation (2-inch (Unit 1), 3-inch (Unit 2), and 4-inch (Units 1 & 2) only) 5.2.2.6 Conclusions The SBLOCA analysis results meet the pertinent acceptance criteria of 10 CFR 50.46 (Reference 18). The PCT is less than 2,200 F; the MLO is less than 17 percent; the core-wide hydrogen generation is less than 1 percent; the core geometry remains amenable to cooling; and the core temperature is maintained at an acceptably low value by the time the transient is terminated for all cases examined herein.

5.2.3 Post-LOCA Long-Term Subcritica lity Cooling Evaluation The Westinghouse licensing position for satisfying the requirements of 10 CFR Part 50.46 (b)(5) "Long Term Cooling" is defined in WCAP-8339-NP-A (Reference 20), WCAP-8472-NP-A (Reference 21), and Technical Bulletin NSID-TB-86-08 (Reference 22). Analyses demonstrate that the reactor will remain shut down by borated ECCS water alone after a LOCA.

5-75 Since credit for the control rods is not taken for an LBLOCA, the RCS inventory, spilled ECCS water, and other water that ends up in the containment sump must have a sufficient boron concentration such that the reactor core remains subcritical assuming all control rods out. The confirmation of the adequacy of the sump boron concentration is performed for each reload cycle as part of the Westinghouse Reload Evaluation Process (Reference 23). The introduction of 422V+ fuel features would have a negligible effect on the calculation of the minimum boron concentration in the containment sump after a large-break LOCA. Changes to the RCS initial conditions (pre-LOCA RCS liquid mass) would also have a small effect on the calculated containment sump boron concentration. The post-LOCA long-term containment mixed mean sump boron concentration was recalculated for the 422V+ FU Program. This mixed mean sump boron calculation bounds the recommended increase in minimum accumulator boron concentrations, as described in subsection 5.2.1. A plot of the post-LOCA containment sump boron concentration as a function of the pre-trip RCS boron concentration is included in the RSAC and is used to confirm that each cycle-specific reload will remain subcritical during long-term cooling. As part of the 422V+ FU study, four representative loading patterns were evaluated to provide reasonable assurance that Technical Specifications on minimum refueling water storage tank (RWST) and accumulator boron concentrations would be reasonably adequate to assure post-LOCA long-term subcriticality margin for future core designs. It was determined that sufficient margin existed for future core designs with consideration of the normal iterative core design process, which can utilize additional burnable absorbers to reduce the core Keff. Subcriticallity will be confirmed for each specific core reload as part of the RSAC.

5.2.4 Post-LOCA Boron Buildup Analysis and Long-Term Post-LOCA Cooling For Prairie Island Units 1&2, a post-LOCA core boric acid precipitation analysis was most recently performed as part of the Safety Analysis Transition Program/RSG Program (Reference 24). This analysis was redone for the 422V+ Fuel Transition Program because of the effects of the slightly larger fuel rods on the core region volume. The boric acid precipitation re-analysis for Prairie Island Units 1&2 for the 422V+ Fuel Transition Program demonstrates the acceptability of a 7.5-hour criteria for initiating low-head safety injection (LHSI) after a LOCA. Th is is consistent with the existing USAR Section 6.2.3.12 timing criteria.

In order to demonstrate effective decay heat removal for long-term cooling, the capability of the safety injection system after switchover to sump recirculation mode cooling must be confirmed. For LBLOCA, core cooling is ensured by confirming that there is sufficient ECCS flow to offset core boiloff and boiling in the downcomer and lower plenum (Reference 25). This analysis was determined not to be adversely impacted by the transition to 422V+ fuel at FU operating conditions. For an SBLOCA, any potential effects due to ECCS flow interruptions, reductions, and/or enthalpy changes at switchover to recirculation mode are considered as part of the SBLOCA analysis. For the 422V+ Fuel Transition Program, it was determined that recirculation ECCS flows are sufficient to maintain post-LOCA long-term core cooling.

5-76 5.2.5 LOCA Hydraulic Forces The LOCA hydraulic forces created by a hypothetical break in the RCS piping are principally caused by the motion of the decompression wave through the RCS. The strength of the decompression wave is primarily a function of the assumed break opening time; break area; and RCS operating conditions of power, temperature, and pressure. Thermal-hydraulic data that bounds the 422V+ fuel was used in generating vessel LOCA forcing functions for Prairie Island Units 1&2. The LOCA hydraulic forces analysis for the reactor vessel and internals were generated using the advanced beam model version of MULTIFLEX (3.0) (Reference 26), assuming a conservative break opening time of 1 msec; in accordance with the methodology accepted by the NRC in WCAP-15029 (Reference 27). This version of the MULTIFLEX code shares a common hydraulic modeling scheme with the NRC-approved MULTIFLEX (1.0) computer code (Reference 28), with the differences being confined to a more realistic downcomer hydraulic network and a more realistic core barrel structural model that accounts for nonlinear boundary conditions and vessel motion. Generally, this improved modeling results in more realistic, but still conservative, hydraulic forces on the core barrel.

Post-processing codes were then used to generate the LOCA forces on the component of interest from the thermal-hydraulic data calculated from the MULTIFLEX code. The LATFORC and FORCE2 post-processing codes (Reference 28) were then used to generate the LOCA forces on the component of interest from the thermal-hydraulic data calculated from the MULTIFLEX code. Forces acting on the RCS loop piping and steam generator as a result of a hypothesized LOCA are not significantly influenced by changes in fuel assembly design.

5-77 Table 5.2.1-1 Prairie Island Unit 1 Best-Estimate UPI Large-Break LOCA Results Value Criteria 95th Percentile PCT (°F) 1,765 <2,200 Maximum Cladding Oxidation (%) 0.62 <17 Maximum Hydrogen Generation (%) 0.014 <1 Coolable Geometry Core remains coolable Core remains coolable Long-Term Cooling Core remains cool in long term Core remains cool in long term Table 5.2.1-2 Prairie Island Unit 1 Conditions Analyzed with WCOBRA/TRAC Compared to Best-Estimate UPI Test Conditions Condition BE UPI Test Prairie Island Unit 1 Core Power, MWt 1,980 1,683 Low Power Region Average Linear Heat Rate (kW/ft) 6.9 4.56 Peak Linear Heat Rate (kW/ft) 17.0 14.1 Table 5.2.1-3 Prairie Island Unit 2 Best-Estimate UPI Large-Break LOCA Results Value Criteria 95th Percentile PCT (°F) 1,623 <2,200 Maximum Cladding Oxidation (%) 0.56 <17 Maximum Hydrogen Generation (%) 0.002 <1 Coolable Geometry Core remains coolable Core remains coolable Long-Term Cooling Core remains cool in long term Core remains cool in long term Table 5.2.1-4 Prairie Island Unit 2 Conditions Analyzed with WCOBRA/TRAC Compared to Best-Estimate UPI Test Conditions Condition BE UPI Test Prairie Island Unit 2 Core Power, MWt 1,980 1,683 Low Power Region Average Linear Heat Rate (kW/ft) 6.9 4.56 Peak Linear Heat Rate (kW/ft) 17.0 14.1 5-78 Table 5.2.2-1 Input Assumptions and Initial Conditions for Units 1&2 Unit 1 Unit 2 A. Core Parameters Analyzed Core Power 1,683 MWt Total Core Peaking Factor, F Q 2.5 Channel Enthalpy Rise Factor, FH 1.77 Axial Power Shape Figure 5.2.2-2 Axial Offset 13% K(z) Limit Single line segment B. Reactor Coolant System Thermal Design Flow 89,000 gpm/loop Nominal Vessel Average Temperature (T AVG) 560.0 °F Vessel Average Temperature Uncertainty +/- 4 °F Pressurizer Pressure 2,250 psia Pressurizer Pressure Uncertainty +/- 60 psi C. Reactor Protection System Low Pressurizer Pressure Reactor Trip Setpoint 1,700 psia Reactor Trip Signal Processing Time (Includes Rod Drop Time) 4.4 seconds D. Auxiliary Feedwater System Maximum AFW Temperature 100 °F Minimum AFW Flow Rate 90 gpm/SG Initiation Signal Low Pressurizer Pressure Reactor Trip with LOOP AFW Delivery Delay Time 60 seconds E. Steam Generators Steam Generator Tube Plugging 10% 25% MFW Isolation Signal Low Pressurizer Pressure SI Signal MFW Isolation Delay Time 1.5 seconds MFW Flow Coastdown Time 4 seconds Feedwater Temperature 437.5 °F Steam Generator Safety Valve Flow Rates Table 5.2.2-2 5-79 Table 5.2.2-1 Input Assumptions and Initial Conditions for Units 1&2 (cont.) Unit 1 Unit 2 F. Safety Injection (SI) Limiting Single Failure 1 EDG Maximum SI Water Temperature 120°F Low Pressurizer Pressure Signal Setpoint 1,700 psia SI Delay Time 27 seconds Safety Injection Flow Rates Tables 5.2.2-3 through 5.2.2-5 G. Accumulators Water/Gas Temperature 120°F Nominal Accumulator Water Volume 1,270 ft 3 Minimum Cover Gas Pressure (including uncertainty) 699.7 psia H. RWST Draindown Input Maximum Containment Spray Flow 1,600 gpm/train Minimum Usable RWST Volume (1) 117,400 gal (RHR) 140,880 gal (HHSI) Maximum Interruption of High-Head Safety Injection (HHSI)

Flow During Switchover to Cold Leg Recirculation 504 seconds Safety Injection Flow Rate vs. Pressure During HHSI Switchover to Cold Leg Recirculation No flow Safety Injection Flow Rate vs. Pressure During RHR Switchover to Cold Leg Recirculation Table 5.2.2-3 Maximum SI Water Temperature After Switchover to Recirculation 212°F Note: 1. The minimum RWST volume represents the smallest volume of liquid that would be drained from the RWST and initiate switchover to cold leg recirculation for the indicated pump. These values were used to conservatively model when switchover to cold leg recirculation would begin during each transient.

5-80 Table 5.2.2-2 Steam Generator Safety Valve Flows per Steam Generator MSSV - Set Pressure psig Setpoint Uncertainty % of Nominal Steam Accumulation % of Nominal Minimum Full Flow at Steam Accumulation lbm/hr 1 1,077 3 3 695,728 2 1,093 3 3 706,602 3 1,110 3 3 718,252 4 1,120 3 3 725,146 5 1,131 3 3 732,816 Table 5.2.2-3 HHSI Flow for One HHSI Pump with Faulted Loop Spilling to RCS Pressure (Break sizes less than 5.187 inches) RCS Pressure psia Spilled Flow gpm Injected Flow gpm 14.7 292.3 276.4 114.7 285.0 269.5 214.7 277.5 262.5 314.7 270.0 255.4 414.7 262.4 248.2 514.7 254.6 240.9 614.7 246.7 233.4 714.7 238.2 225.3 814.7 229.5 217.1 914.7 220.6 208.7 1,014.7 211.6 200.2 1,114.7 202.5 191.5 1,214.7 192.1 181.6 1,314.7 181.1 171.2 1,414.7 169.7 160.6 1,514.7 158.1 149.5 1,614.7 144.0 136.2 1,714.7 127.6 120.7 1,814.7 110.0 104.0 1,914.7 86.2 81.6 2,014.7 55.9 52.9 2,114.7 0.0 0.0 5-81 Table 5.2.2-4 HHSI Flows for One HHSI Pump with Faulted Loop Spilling to Containment Pressure (0 psig) (Break sizes greater than 5.187 inches) RCS Pressure psia Spilled Flow gpm Injected Flow gpm 14.7 292.3 276.5 114.7 380.1 259.6 214.7 418.0 213.7 314.7 457.5 165.1 414.7 498.9 113.1 514.7 542.9 56.8 614.7 586.5 0.0 2,314.7 586.5 0.0 Table 5.2.2-5 RHR Flows for One RHR Pump with No Spilling During Injection from RWST RCS Pressure psia Injected Flow gpm 14.7 1,605.4 34.7 1,473.1 54.7 1,330.0 74.7 1,165.1 94.7 972.7 114.7 741.2 134.7 404.9 145.6 0.0

5-82 Table 5.2.2-6 Time Sequence of Events for Unit 1 Event (sec) 1.5-Inch 2-Inch 3-Inch 4-Inch 6-Inch 8-Inch 10.126-Inch Transient Initiated 0.0 0.0 0.0 0.0 0.0 0.0 0.0 Reactor Trip Signal 57.6 29.8 13.2 8.1 5.7 5.2 5.0 Safety Injection Signal 57.6 29.8 13.2 8.1 5.7 5.2 5.0 Safety Injection Begins (1) 84.6 56.8 40.2 35.1 32.7 32.2 32.0 Loop Seal Clearing Occurs (2) 1,191 653 265 158 27 15 10 Top of Core Uncovered N/A (3) 1,401 574 393 N/A (3) N/A (3) N/A (3) Accumulator Injection Begins 6,704 2,301 705 362 164 94 53 Top of Core Recovered N/A (3) 2,548 889 429 N/A (3) N/A (3) N/A (3) RWST Low Level N/A (4) 5,288 2,304.3 2,283.8 1,668.7 1,523.4 1,448.9 Notes: 1. SI begins 27.0 seconds (SI delay time) after the SI signal is generated. 2. Loop seal clearing is considered to occur when the broken loop (BL) loop seal vapor flow rate is sustained above 1 lbm/s. 3. There is no core uncovery for the 1.5-inch break case and only brief core uncovery for the 6-inch, 8-inch and 10.126-inch break cases. 4. The RWST low level is not reached for this break size.

Table 5.2.2-7 Beginning-of-Life Fuel Rod Heatup Results for Unit 1 Results 1.5-Inch 2-Inch 3-Inch 4-Inch 6-Inch 8-Inch 10.126-Inch PCT, °F 954.2 958.9 603.7 PCT Time, sec 1,922.9 787.1 414.4 PCT Elevation, ft 11.25 10.75 11.00 Burst Time (1), sec N/A N/A N/A Burst Elevation (1), ft N/A N/A N/A Maximum ZrO 2, % 0.01 0.01 0.00 Maximum ZrO 2 Elevation, ft 11.25 11.00 12.00 Average ZrO 2 , % N/A (2) 0.0 0.0 0.0 N/A (2) N/A (2) N/A (2) Notes: 1. Neither the hot rod nor the hot assembly average rod burst during the LOCTA-IV calculations. 2. The core either does not uncover or only uncovers for a very short time; therefore, LOCTA-IV calculations are not warranted for these break sizes.

5-83 Table 5.2.2-8 Time Sequence of Events for Unit 2 Event (sec) 1.5-Inch 2-Inch 3-Inch 4-Inch 6-Inch 8-Inch 10.126-Inch Transient Initiated 0.0 0.0 0.0 0.0 0.0 0.0 0.0 Reactor Trip Signal 57.5 29.8 13.4 8.5 6.4 5.9 5.7 Safety Injection Signal 57.5 29.8 13.4 8.5 6.4 5.9 5.7 Safety Injection Begins (1) 84.5 56.8 40.4 35.5 33.4 32.9 32.7 Loop Seal Clearing Occurs (2) 1,089 589 236 138 27 14 19 Top of Core Uncovered N/A (3) 1,332 603 340 N/A (3) N/A (3) N/A (3) Accumulator Injection Begins 6,113 2,172 677 332 153 81 46 Top of Core Recovered N/A (3) 2,527 839 403 N/A (3) N/A (3) N/A (3) RWST Low Level N/A (4) 5,467.2 2,303.2 2,283.1 1,654.8 1,513.5 1,460.2 Notes: 1. SI begins 27.0 seconds (SI delay time) after the SI signal is generated. 2. Loop seal clearing is considered to occur when the BL loop seal vapor flow rate is sustained above 1 lbm/s. 3. There is no core uncovery for the 1.5-inch and 6-inch break cases and only brief core uncovery for the 8-inch and 10.126-inch break cases. 4. The RWST low level is not reached for this break size.

Table 5.2.2-9 Beginning-of-Life Fuel Rod Heatup Results for Unit 2 Results 1.5-Inch 2-Inch 3-Inch 4-Inch 6-Inch 8-Inch 10.126-Inch PCT, °F 964.4 905.1 668.5 PCT Time, sec 1,835.9 760.1 390.8 PCT Elevation, ft 11.25 10.75 11.25 Burst Time (1), sec N/A N/A N/A Burst Elevation (1), ft N/A N/A N/A Maximum ZrO 2, % 0.01 0.00 0.00 Maximum ZrO 2 Elevation, ft 11.25 11.00 11.25 Average ZrO 2 , % N/A (2) 0.00 0.00 0.00 N/A (2) N/A (2) N/A (2) Notes: 1. Neither the hot rod nor the hot assembly average rod burst during the LOCTA-IV calculations. 2. The core either does not uncover or only uncovers for a very short time; therefore, LOCTA-IV calculations are not warranted for these break sizes.

Figure 5.2.2-1. Code Interface Description for Small Break Model N 0 T R U M P CORE PRESSURE, CORE FLOW, MIXTURE LEVEL, AND FUEL ROD POWER HISTORY O<TIME<CORE COVERED v b "I 5-85 Figure 5.2.2-2. Hot Rod Axial Power Shape 5-86 Figure 5.2.2-3. Reactor Coolant System Pressure Inch Break (Unit 1) 5-87 Figure 5.2.2-4. Core Mixture Level Inch Break (Unit 1) 5-88 Figure 5.2.2-5. Core Exit Vapor Temperature Inch Break (Unit 1) 5-89 Figure 5.2.2-6. Broken Loop and Intact Loop Pumped SI Flow Rates Inch Break (Unit 1) 5-90 Figure 5.2.2-7. Total Break Flow and Total Pumped SI Flow Rates Inch Break (Unit 1) 5-91 Figure 5.2.2-8. Cladding Temperature at PCT Elevation and Local Oxidation at MLO Elevation Inch Break (Unit 1) 5-92 Figure 5.2.2-9. Reactor Coolant System Pressure Inch Break (Unit 2) 5-93 Figure 5.2.2-10. Core Mixture Level Inch Break (Unit 2) 5-94 Figure 5.2.2-11. Core Exit Vapor Temperature Inch Break (Unit 2) 5-95 Figure 5.2.2-12. Broken Loop and Intact Loop Pumped SI Flow Rates Inch Break (Unit 2) 5-96 Figure 5.2.2-13. Total Break Flow and Total Pumped SI Flow Rates Inch Break (Unit 2) 5-97 Figure 5.2.2-14. Cladding Temperature at PCT Elevation and Local Oxidation at MLO Elevation Inch Break (Unit 2) 5-98 Figure 5.2.2-15. Reactor Coolant System Pressure - 1.5-Inch Break (Unit 1) 5-99 Figure 5.2.2-16. Core Mixture Level - 1.5-Inch Break (Unit 1) 5-100 Figure 5.2.2-17. Reactor Coolant System Pressure Inch Break (Unit 1) 5-101 Figure 5.2.2-18. Core Mixture Level Inch Break (Unit 1) 5-102 Figure 5.2.2-19. Cladding Temperature at PCT Elevation and Local Oxidation at MLO Elevation Inch Break (Unit 1) 5-103 Figure 5.2.2-20. Reactor Coolant System Pressure Inch Break (Unit 1) 5-104 Figure 5.2.2-21. Core Mixture Level Inch Break (Unit 1) 5-105 Figure 5.2.2-22. Cladding Temperature at PCT Elevation and Local Oxidation at MLO Elevation Inch Break (Unit 1) 5-106 Figure 5.2.2-23. Reactor Coolant System Pressure Inch Break (Unit 1) 5-107 Figure 5.2.2-24. Core Mixture Level Inch Break (Unit 1) 5-108 Figure 5.2.2-25. Reactor Coolant System Pressure Inch Break (Unit 1) 5-109 Figure 5.2.2-26. Core Mixture Level Inch Break (Unit 1) 5-110 Figure 5.2.2-27. Reactor Coolant System Pressure - 10.126-Inch Break (Unit 1) 5-111 Figure 5.2.2-28. Core Mixture Level - 10.126-Inch Break (Unit 1) 5-112 Figure 5.2.2-29. Reactor Coolant System Pressure - 1.5-Inch Break (Unit 2) 5-113 Figure 5.2.2-30. Core Mixture Level - 1.5-Inch Break (Unit 2) 5-114 Figure 5.2.2-31. Reactor Coolant System Pressure Inch Break (Unit 2) 5-115 Figure 5.2.2-32. Core Mixture Level Inch Break (Unit 2) 5-116 Figure 5.2.2-33. Cladding Temperature at PCT Elevation and Local Oxidation at MLO Elevation Inch Break (Unit 2) 5-117 Figure 5.2.2-34. Reactor Coolant System Pressure Inch Break (Unit 2) 5-118 Figure 5.2.2-35. Core Mixture Level Inch Break (Unit 2) 5-119 Figure 5.2.2-36. Cladding Temperature at PCT Elevation and Local Oxidation at MLO Elevation Inch Break (Unit 2) 5-120 Figure 5.2.2-37. Reactor Coolant System Pressure Inch Break (Unit 2) 5-121 Figure 5.2.2-38. Core Mixture Level Inch Break (Unit 2) 5-122 Figure 5.2.2-39. Reactor Coolant System Pressure Inch Break (Unit 2) 5-123 Figure 5.2.2-40. Core Mixture Level Inch Break (Unit 2) 5-124 Figure 5.2.2-41. Reactor Coolant System Pressure - 10.126-Inch Break (Unit 2) 5-125 Figure 5.2.2-42. Core Mixture Level - 10.126-Inch Break (Unit 2) 5-126 5.3 RUPTURES IN PRIMARY SIDE PIPING The uncontrolled release of pressurized high-temperature reactor coolant from the primary system piping, known as a LOCA, will inject high-energy steam and water into the containment. This release of mass and energy to the containment will increase containment pressure and temperature. This section discusses the design issues related to a postulated LOCA and containment pressurization that must be considered for the Prairie Island Units 1&2 with 422V+ fuel at FU analyzed operating conditions. The long-term LOCA mass and energy releases are analyzed and are utilized as input to the containment integrity analysis, which demonstrates the acceptability of the containment safeguards systems to mitigate the consequences of a hypothetical LBLOCA. The containment safeguards systems must be capable of limiting the peak containment pressure to less than the design pressure of the containment shell and to limit the temperature excursion to less than the acceptance limits. Subsection 5.3.1 discusses the long-term LOCA mass and energy releases generated for this program. The mass and energy releases from this analysis were provided for use in the containment integrity analysis (subsection 5.3.2). An evaluation of the short-term containment pressurization was performed (subsection 5.3.4), and it was determined that application of leak-before-break (LBB) technology to Prairie Island has reduced the mass and energy releases to well below the original licensing basis.

5.3.1 Long-Term LOCA Mass and Energy Releases Inside Containment The mass and energy release rates described in this section form the input to further computations to evaluate the containment conditions following the postulated accident. The long-term LOCA mass and energy releases for the hypothetical double-ended pump suction (DEPS) break with minimum safeguards, the DEPS break with maximum safeguards and double-ended hot-leg (DEHL) rupture break cases are discussed in this section. These breaks represent the bounding containment pressurization accident sequences for the LBLOCA events. The mass and energy releases from these three LOCA cases are used for the long-term containment integrity analyses described in subsection 5.3.2. For this program, Westinghouse generated the mass and energy releases using the March 1979 model, described in WCAP-10325-P-A (Reference 29). The review and approval letter for the methodology is included within WCAP-10325-P-A (Reference 29). 5.3.1.1 Input Parameters and Assumptions The mass and energy release analysis is sensitive to the assumed characteristics of various plant systems, in addition to other key modeling assumptions. Where appropriate, bounding input values are utilized and instrumentation uncertainties are included. All input parameters are chosen consistent with accepted analysis methodology (Reference 29). Tables 5.3-1 through 5.3-3 present key data assumed in the analysis.

5-127 The following assumptions were employed to ensure that the mass and energy releases are conservatively calculated, thereby maximizing the energy release to containment: Maximum expected operating temperature of the RCS (at full-power conditions) Allowance for RCS temperature uncertainty (+4.0°F) Margin in RCS volume of 3 percent (which is composed of 1.6-percent allowance for thermal expansion, and 1.4-percent allowance for uncertainty) Core rated power plus uncertainties of 1,683 MWt Conservative heat transfer coefficients (that is, steam generator primary/secondary heat transfer, and RCS metal heat transfer) Allowance in core-stored energy for effect of fuel densification A margin in core-stored energy (+15 percent to account for manufacturing tolerances) for the bounding fuel product An allowance for RCS initial pressure uncertainty (+40 psi) A maximum containment backpressure equal to design pressure (46.0 psig) Steam generator tube plugging leveling (0-percent uniform)

- Maximizes reactor coolant volume and fluid release

- Maximizes heat transfer area across the steam generator tubes

- Reduces coolant loop resistance, which reduces the P upstream of the break for the pump suction breaks and increases break flow 5.3.1.2 Description of Analyses The Westinghouse evaluation model used for the long-term LOCA mass and energy release calculations is the March 1979 model described in WCAP-10325-P-A (Reference 29). Assumptions in the Westinghouse mass and energy release methodology identified in Reference 30 to potentially create nonconservatism in containment pressure calculations were explicitly addressed in this analysis. These mass and energy releases are then subsequently used in the containment integrity analysis (subsection 5.3.2).. One mass and energy release model was created that bounds both Prairie Island units. The reactor coolant systems of the two units are very similar, but where differences exist that could impact the mass and energy releases (that is, volume of the accumulator piping), the conservative value was chosen for the parameter. Framatome RSGs were modeled to produce bounding mass and energy releases since more energy is available to be released from the RSGs than from the Westinghouse OSGs.

5-128 LOCA Mass and Energy Release Phases The containment system receives mass and energy releases following a postulated rupture in the RCS. These releases continue over a time period, which for the LOCA mass and energy analysis, is typically divided into four phases: 1. Blowdown - the period of time from accident initiation (when the reactor is at full-power, steady-state operation) to the time that the RCS and containment reach an equilibrium state. 2. Refill - the period of time when the lower plenum is filled by accumulator and ECCS water. At the end of blowdown, water remains in the cold legs, downcomer, and lower plenum. To conservatively consider the refill period for the purpose of containment mass and energy releases, it is assumed that this water is instantaneously transferred to the lower plenum along with sufficient accumulator water to completely fill the lower plenum. This allows an uninterrupted release of mass and energy to containment. Thus, the refill period is conservatively neglected in the mass and energy release calculation. 3. Reflood - begins when the water from the lower plenum enters the core and ends when the core is completely quenched. 4. Post-reflood (FROTH) - describes the period following the reflood phase. For the pump suction break, a two-phase mixture exits the core, passes through the hot legs, and is superheated in the steam generators prior to exiting the break as steam. After the broken loop steam generator cools, the break flow becomes two phase. Computer Codes The WCAP-10325-P-A (Reference 29) mass and energy release evaluation model is comprised of mass and energy release versions of the following codes: SATAN-VI, WREFLOOD, FROTH, and EPITOME. These codes were used to calculate the long-term LOCA mass and energy releases for Prairie Island. SATAN-VI calculates the blowdown conditions during the first phase of the thermal-hydraulic transient following break initiation, including pressure, enthalpy, density, mass and energy flow rates, and energy transfer between primary and secondary systems as a function of time. The WREFLOOD code addresses the core reflood phase of the transient after the primary coolant system has depressurized (blowdown). The core reflooding phase occurs when water supplied by the ECCS refills the reactor vessel and provides cooling to the overheated core. The most important feature of WREFLOOD is the steam/water mixing model (see Section "Reflood Mass and Energy Release Data" in subsection 5.3.1.4.) The FROTH code models the post-reflood phase of the accident sequence during the secondary system heat addition from the broken and intact loop steam generators to the steam and water passing through the tubes.

5-129 EPITOME continues the FROTH post-reflood portion of the transient from the time at which the secondary system equilibrates to containment design pressure to the end of the transient. Break Size and Location Generic studies have been performed with respect to the effect of postulated break size on the LOCA mass and energy releases. The double-ended guillotine break has been found to be limiting due to larger mass flow rates during the blowdown phase of the transient. During the reflood and froth phases, the break size has little effect on the releases. Three distinct locations in the RCS loop can be postulated for a pipe rupture for mass and energy release purposes: 1. Hot leg (between vessel and steam generator) 2. Cold leg (between pump and vessel) 3. Pump suction (between steam generator and pump) The break locations analyzed for this program are the DEPS rupture (10.46 ft

2) and the DEHL rupture (9.154 ft 2). Break mass and energy releases are calculated for the blowdown, reflood, and post-reflood phases of the LOCA for the DEPS cases. For the DEHL case, the releases are calculated only for the blowdown. The DEHL and DEPS cases are used to analyze long-term LOCA containment integrity. Application of Single-Failure Criterion An analysis of the effects of the single-failure criterion has been performed on the mass and energy release rates for each break analyzed. An inherent assumption in the generation of the mass and energy release is that offsite power is lost. Loss of offsite power results in the actuati on of the EDGs, required to power the safety injection system. This is not an issue for the blowdown period, which is limited by the DEHL break. Two cases have been analyzed to determine bounding mass and energy releases while addressing a single failure. The first case assumes minimum safeguards SI flow based on the postulated single failure of one EDG, which results in the loss of one train of safeguards equipment. The second case assumes maximum safeguards SI flow consistent with the approved methodology described in WCAP-10325-P-A (Reference 29). In the maximum safeguards case, a single failure of the containment heat removal equipment is assumed in the containment response analysis documented in subsection 5.3.2. The analysis of the minimum and maximum safeguards cases for the DEPS break provides confidence that a bounding peak containment pressure is identified with consideration of credible single failures. 5.3.1.3 Acceptance Criteria for Analysis An LBLOCA accident is classified as an American Nuclear Society (ANS) Condition IV event, an infrequent fault. To satisfy the NRC acceptance criteria presented in the Standard Review Plan (SRP), Section 6.2.1.3, the relevant requirements are the following: 10 CFR 50, Appendix A, "General Design Criteria for Nuclear Power Plants."

5-130 10 CFR 50, Appendix K, "ECCS Evaluation Models," paragraph I.A, "Required and Acceptable Features of the ECCS Evaluation Models - Sources of Heat during a LOCA." To meet these requirements, the mass and energy releases must address: Sources of energy Break size and location Calculation of each phase of the accident Each of these items is considered in the Westinghouse mass and energy release methodology (Reference 29). These mass and energy releases are then imposed upon a model of the containment to confirm that the containment conditions remain within the design limits of the containment pressure boundary (See subsection 5.3.2.4) 5.3.1.4 Mass and Energy Release Results Blowdown Mass and Energy Release Data The SATAN-VI code is used for computing the blowdown transient. The code utilizes the control volume (element) approach with the capability for modeling a large variety of thermal fluid system configurations. The methodology for the use of this model is described in WCAP-10325-P-A (Reference 29). Reflood Mass and Energy Release Data for the DEPS Breaks The WREFLOOD code is used for computing the refl ood transient considering transient phenomena such as pumped SI and accumulators, RCP performance, and steam generator energy release. A complete thermal equilibrium mixing condition for the steam and ECCS injection water during the reflood phase has been assumed for each loop receiving ECCS water. This is consistent with the usage and application of the WCAP-10325-P-A (Reference 29) mass and energy release evaluation model in approved analyses, for example, D. C. Cook Docket (Reference 31). This assumption is supported by test data (Reference 32). Post-Reflood Mass and Energy Release Data The FROTH code (Reference 33) computes the post-reflood transient. The code calculates the heat release rates resulting from a two-phase mixture present in the steam generator tubes. The mass and energy releases that occur during this phase are typically superheated (Reference 34) due to the depressurization and equilibration of the broken loop and intact loop steam generators. The calculation stops when the secondary side equilibrates to the saturation temperature (Tsat) at the containment design pressure, after this point the EPITOME code completes the steam generator depressurization (see "Steam Generator Equilibration and Depressurization" in this section for additional information). The methodology for the use of this model is described in WCAP-10325-P-A (Reference 29).

5-131 Decay Heat Model for Mass and Energy Release Calculation The ANS approved ANS Standard 5.1 (Reference 35) for the determination of decay heat. This standard was used in the mass and energy release model for Prairie Island. Significant assumptions in the generation of the decay heat curve for use in the LOCA mass and energy releases analysis include the following: The decay heat sources considered are fission product decay and heavy element decay of U-239 and Np-239. The decay heat power from the fission of isotopes other than U-235 is assumed to be identical to that of U-235. The fission rate is constant over the operating history of maximum power level. The factor accounting for neutron capture in fission products has been taken from Reference 35. The fuel has been assumed to be at full-power for 10 8 seconds. The total recoverable energy associated with one fission has been assumed to be 200 MeV/fission. Two sigma uncertainty (two times the standard deviation) has been applied to the fission product decay. Based upon NRC staff review, [SER of the March 1979 evaluation model (Reference 29)], use of the ANS Standard-5.1, November 1979 decay heat model was approved for the calculation of mass and energy releases to the containment following a LOCA. Steam Generator Equilibration and Depressurization Steam generator equilibration and depressurization is the process by which secondary side energy is removed from the steam generators in stages. The FROTH code calculates the heat removal from the secondary mass until the secondary temperature is the saturation temperature at the containment design pressure. After the FROTH calculations, the EPITOME code continues the calculation for steam generator cooldown until the secondary system reaches the saturation temperature at 14.7 psia, or 212°F. The heat removal of the broken loop and intact loop steam generators are calculated separately. Sources of Mass and Energy The sources of mass considered in the LOCA mass and energy release analysis are the RCS, accumulators, and pumped SI. The energy inventories considered in the LOCA mass and energy release analysis are the following:

5-132 RCS water Accumulator water (both inject) Pumped SI water Decay heat Core-stored energy RCS metal (includes steam generator tubes) Steam generator metal (includes transition cone, shell, wrapper, and other internals) Steam generator secondary energy (includes fluid mass and steam mass) Secondary transfer of energy (feedwater into and steam out of the steam generator secondary) The analysis used the following energy reference points: Available energy: 212°F; 14.7 psia (energy available that could be released) Total energy content: 32°F; 14.7 psia (total internal energy of the RCS) In the mass and energy release data presented, no zirconium-water reaction heat was considered because the fuel cladding temperature does not rise high enough for the rate of the zirconium-water reaction to proceed. 5.3.1.5 Evaluation of the Long-Term Containment Mass and Energy Releases for 422V+ Fuel at FU Analyzed Conditions The mass and energy releases were generated for Prairie Island Units 1&2 with 400 V+ fuel. These releases calculated for Prairie Island are bounding and conservative for the plants with 422V+ fuel at FU analyzed conditions. The parameters pertaining to the WCAP-10325-P-A (Reference 29) methodology were reviewed for differences between the analyzed conditions and the FU conditions. The description of the overall impact of the fuel upgrade on the RCS and secondary system energy available for release to the containment is discussed in this section.

The reactor power including uncertainties, 1,683 MWt, used in the calculation of the mass and energy releases is the same as the reactor power including uncertainties for the FU analyzed conditions. The use of plant-specific core power uncertainty in the WCAP-10325-P-A methodology was approved by the NRC in Reference 34. The bounding core-stored energy value of 5.77 full-power seconds (FPSs) used in the mass and energy release calculation is a conservative value that bounds the core stored energy of the 422V+ fuel product (4.60 FPS) at the FU analyzed conditions. The reactor vessel average temperature and the thermal design flow have not changed. Therefore, the energy that can be released from the RCS at FU analyzed conditions is bounded by the mass and energy releases calculated for the Prairie Island. The enthalpy of the steam generator secondary system steam/water is essentially unchanged at the FU analyzed conditions and the mass of the secondary system steam/water is slightly decreased at the higher output power level. Therefore, the sensible energy of the steam generator secondary system that can be released to the containment is essentially unchanged due to the FU analyzed conditions. The LOCA containment mass and energy releases calcul ated for Prairie Island Units 1&2 are conservative and applicable to the 422V+ fuel product at FU analyzed conditions.

5-133 5.3.1.6 Conclusions of the Mass and Energy Release Calculations The use of the WCAP-10325-P-A (Reference 29) methodology (consideration of the various break sizes, energy sources and accident phases in the long-term mass and energy release analysis) provides assurance that all required elements of the LOCA mass and energy release have been included in this analysis. Thus, the review guidelines presented in SRP Section 6.2.1.3 have been satisfied. The mass and energy release results from this analysis for the DEHL break, the DEPS break with minimum safeguards, and the DEPS break with maximum safeguards were provided for use in the containment peak pressure integrity analysis presented in subsection 5.3.2 where a limiting case is determined.

5.3.2 LOCA Containment Peak Pressure Response Analysis The Prairie Island containment system is designed so that for all LOCA break sizes, up to and including the double-ended severance of a reactor coolant pipe, the containment peak pressure remains below the design pressure. This section details the containment response subsequent to a hypothetical LOCA. The containment response analysis uses the long-term LOCA mass and energy release data described in subsection 5.3.1. The containment response analysis demonstrates the acceptability of the containment safeguards systems to mitigate the consequences of a LOCA inside containment. 5.3.2.1 Accident Description A break in the primary RCS piping causes a loss of coolant, which results in a rapid release of mass and energy to the containment atmosphere. The blowdown phase causes a rapid increase in the containment pressure, which results in the actuation of the emergency fan cooler and containment spray systems. The RCS accumulators begin to refill the lower plenum and downcomer of the re actor vessel with water after the end of blowdown. The reflood phase begins after the vessel fluid level reaches the bottom of the fuel. During this phase, the core is quenched with water from both the accumulators and pumped SI. The quenching process creates a large amount of steam and entrained water that is released to containment through the break. If the break is located in the cold leg or pump suction leg piping, the two-phase mixture from the core would pass through the steam generators and absorb energy from the secondary-side of the steam generators. The LOCA mass and energy release decreases with time as the system cools. Core decay heat is removed by nucleate boiling after the reflood and froth phases are complete. The long term core fluid level is maintained by pumping water back into the vessel from the sump recirculation system. The containment heat removal systems continue to condense steam and reduce the containment pressure and temperature over time. 5.3.2.2 Input Parameters and Assumptions A series of analyses, using bounding break sizes and locations, was performed for the LOCA containment response. Subsection 5.3.1 documents the mass and energy releases for the DEPS and DEHL breaks. The DEPS break cases were run with both minimum and maximum safeguards. The minimum safeguards 5-134 case assumes a diesel train failure. This assumpti on leaves one of two containment spray pumps and two of four containment fan coil units (CFCUs) available as active containment heat removal systems. One residual heat removal (RHR) pump and heat exchanger is available in the minimum SI case. For the maximum safeguards DEPS case, one train of sprays and one train of fan coolers were also modeled along with two RHR pumps and heat exchangers. The containment parameters assumed for the peak pressure containment response analyses are shown in Table 5.3-4. The values are chosen conservatively to maximize the peak pressure for the containment integrity analysis. The major assumptions made in the containment response analysis are the following: The LOCA mass and energy release input to the containment model is described in subsection 5.3.1. Homogeneous mixing is assumed. The steam-air mixture and the water phases each have uniform properties. More specifically, thermal equilibrium between the air and the steam is assumed. However, this does not imply thermal equilibrium between the steam-air mixture and the water phase. Air is taken as an ideal gas, while compressed water and steam tables are employed for water and steam thermodynamic properties. For the blowdown portion of the LOCA analysis, the discharge flow separates into steam and water phases at the breakpoint. The water phase is saturated at the total containment pressure, while the steam phase is saturated at the partial pressure of the steam in the containment. Steam and water releases are input separately for the post-blowdown portion of the LOCA analysis. The saturation temperature at the partial pressure of the steam is used to calculate condensation heat transfer to the heat sinks and the fan coolers. 5.3.2.3 Description of the Prairie Island GOTHIC Containment Model The containment integrity analysis uses the GOTHIC version 7.1 patch 1 computer code to calculate the peak pressure and temperatures for the mass and energy releases presented in subsection 5.3.1. The Prairie Island GOTHIC containment evaluation model consists of a single lumped-parameter node; the diffusion layer model (DLM) is used for heat transfer to all structures in the containment. The model was described in WCAP-16219 (Reference 36). This evaluation model was approved by the NRC in Reference 37. 5.3.2.4 Acceptance Criteria The containment response for design-basis containment integrity is an ANS Condition IV event, an infrequent fault. The containment analysis methodology satisfies the current NRC acceptance criteria from 10 CFR 50, Appendix A and SRP 6.2.1.1.A. The Prairie Island licensing basis for the containment is described in the USAR Sections 5.2.1.1 and 14.9.3. The SRP criteria provide equivalent acceptance 5-135 criteria as the Prairie Island licensing basis. The relevant General Design Criteria (GDC) requirements that are met are: GDC 16 and GDC 50: To satisfy the requirements of GDC 16 and GDC 50, the peak calculated containment pressure should be less than the containment design pressure of 46 psig, considering the most severe single failure GDC 38 and GDC 50: To satisfy the requirements of GDC 38 and GDC 50, the calculated pressure at 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> should be less than 50 percent of the peak calculated value. (This is related to the criteria for doses at 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />.) 5.3.2.5 Analysis Results The sequences of events for the DEPS with minimum safeguard, DEPS with maximum safeguards, and DEHL cases are shown in Tables 5.3-5 through 5.3-7. The containment response calculations for the DEPS cases were performed for 172,800 seconds (2 days). Since the steam generator secondary side energy is effectively isolated for hot leg breaks, the containment response calculation for the DEHL case was performed for the blowdown phase only (approximately 20 seconds). The containment pressure, gas temperature, liner temperature, and water (sump) temperature transients from each of the LOCA cases are shown in Figure 5.3-1 through Figure 5.3-4. Table 5.3-8 summarizes the peak LOCA containment response results for the three cases studied. The following is a summary of the three analyzed cases: Double-Ended Hot Leg Break This analysis assumes a LOOP coincident with a double-ended rupture of the RCS piping between the reactor vessel outlet nozzle and the steam generator inlet (the RCS hot leg). The associated single-failure assumption is the failure of a diesel to start, resulting in one train of safety injection and containment safeguards equipment being available. Figures 5.3-1 through 5.3-4 show the containment pressure, gas temperature, containment liner temperature and sump temperature transients. The peak pressure for this case was 42.9 psig at 18.0 seconds. The sequence of events for this case is shown in Table 5.3-5. Double-Ended Pump Suction Break with Minimum Safeguards This analysis assumes a LOOP coincidence with a double-ended rupture of the RCS piping between the steam generator outlet and the RCS pump inlet. The associated single-failure assumption is the failure of a diesel to start, resulting in only one train of SI pumps and containment safeguards equipment being available. This combination results in a minimum set of safeguards being available. Furthermore, a LOOP delays the actuation times of the safeguards equipment due to the required diesel startup time after receipt of the SI signal. Figures 5.3-1 through 5.3-4 show the containment pressure, atmosphere temperature, containment liner temperature, and sump temperature transients. Table 5.3-6 shows the detailed sequence of events.

5-136 The containment pressure reaches its peak of 43.5 psig at 3,601 seconds (Table 5.3-8). After the steam generator heat release is finished, the containment begins to cool and the pressure and temperature drop. This trend continues to the end of the transient at 172,800 seconds. Double-Ended Pump Suction Break with Maximum Safeguards This analysis assumes a LOOP coincidence with a double-ended rupture of the RCS piping between the steam generator outlet and the RCS pump inlet. The associated single-failure assumption is the failure of one train of containment sprays and fan cooler being available. Furthermore, a LOOP delays the actuation times of the safeguards equipment due to the required diesel startup time after receipt of the SI signal. However, for these cases, it is assumed both trains of ECCS pumps are available consistent with the approved methodology. Figures 5.3-1 through 5.3-4 provide the containment pressure, gas temperature, containment liner temperature and sump temperature transient response. Table 5.3-7 provides the key sequence of events, and Table 5.3-8 shows that a peak pressure of 40.6 psig at 14 seconds was calculated. 5.3.2.6 Conclusions Bounding LOCA containment response analyses have been performed for Prairie Island with 422V+ at FU analyzed operating conditions. The analyses included long-term pressure and temperature profiles for each case. The peak containment pressure was less than 46 psig for all cases. The containment pressure was less than 50 percent of the peak value within 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />. The containment liner temperature was less than 268F in all cases. Based on the results, all applicable containment integrity acceptance criteria for Prairie Island have been met.

5.3.3 Short-Term Mass and Energy Release and Containment Subcompartment Pressurization Evaluation The licensing basis analysis for the Prairie Island Units containment subcompartments was performed by the architect-engineer. Subsequent to the original licensing activities, Prairie Island has applied LBB to the main RCS piping. Application of LBB to Prairie Island has reduced the mass and energy releases to values well below the original licensing basis analysis. In addition, the minor changes to the RCS energy and mass would have minimal impact on these releases. Thus, LOCA short-term mass and energy releases and subcompartment pressurization do not require re-analysis for the 422V+ fuel and design margin remains adequate.

5-137 Table 5.3-1 System Parameters Initial Conditions Parameters Value Core Thermal Power (MWt)

(1) 1,683 RCS Total Flow Rate (lbm/sec)

(1) 19,111.1 Core Outlet Temperature (°F)

(1) 599.8 Core Inlet Temperature (°F)(1) 531.9 Initial Steam Generator Steam Pressure (psia) 784 Steam Generator Design Framatome 56/19 Steam Generator Tube Plugging (%) 0 Initial Steam Generator Secondary-Side Mass (lbm)

(1) 130,711 Assumed Maximum Containment Backpressure (psia) 60.7 Accumulator Water volume (ft

3) per accumulator N 2 cover gas pressure (psig) Temperature (°F) 1,270 685 120 SI Delay, Total (sec) (from beginning of event) 20.1 Note: 1. Core thermal power, RCS total flow rate, RCS coolant temperatures, and steam generator secondary side mass include appropriate uncertainty and/or allowance.

5-138 Table 5.3-2 SI Flow Minimum Safeguards RCS Pressure (psig) Total Flow (lbm/sec) Injection Mode (Reflood Phase) 0 308.80 20 290.20 40 270.00 60 246.90 80 220.00 100 187.70 120 141.00 140 84.80 160 84.40 180 83.90 Recirculation Sequence Vessel Injection (lbm/sec)

Time (sec) From RWST From Sump 1,200 87.9 0 1,440 87.9 207.3 2,520 0 207.3 10,000 0 207.3 50,000 0 207.3 82,500 0 207.3 118,500 0 207.3 1,000,000 0 207.3

5-139 Table 5.3-3 SI Flow Maximum Safeguards RCS Pressure (psig) Total Flow (lbm/sec) Injection Mode (Reflood Phase) 0 753.80 20 713.40 40 669.70 60 625.30 80 577.30 100 522.60 120 460.60 140 376.80 160 225.30 180 161.60 Recirculation Sequence Vessel Injection (lbm/sec)

Time (sec) From RWST From Sump 600 376.9 0 840 376.9 277.9 1,440 0 277.9 1,680 0 555.8 10,000 0 555.8 100,000 0 555.8 1,000,000 0 555.8

5-140 Table 5.3-4 LOCA Containment Response Analysis Parameters Service Water Temperature ( F) 95 RWST Water Temperature ( F) 120 Initial Containment Temperature ( F) 120 Initial Containment Pressure (psia) 16.7 Initial Relative Humidity (%) 30 Net Free Volume (ft

3) 1.32 x 10 6 CFCU Total 4 Analysis Maximum 2 Analysis Minimum 2 Containment High Setpoint (psig) 5.00 Delay Time (sec) 60 Heat Removal as a Function of Steam Temperature Vap Sat Temp ( F) Heat Removal (Btu/s) 0.0 100.0 120.0 140.0 160.0 180.0 200.0 220.0 240.0 240.1 260.0 265.0 270.0 270.1 300.0 0.0 500 2,000 4,500 6,250 8,250 11,000 14,000 17,750 9,861 12,500 13,194 13,611 0 0 Containment Spray Pumps Total 2 Analysis Maximum 1 Analysis Minimum 1 Flow Rate (gpm) Injection Phase (per pump)

Recirculation Phase 1,200 0 Containment High-High Setpoint (psig) 24.0 Delay time (sec) 72. Recirculation Switchover, Full Flow Established, (sec) Minimum Safeguards Maximum Safeguards 1,440 1,680 CS Termination Time, (sec) 1,440

5-141 Table 5.3-4 LOCA Containment Response Analysis Parameters (cont.) RHR System RHR Heat Exchangers Modeled in Analysis Minimum SI Maximum SI 1

2 RHR Flows through RHR Heat Exchangers Minimum Safeguards Time (sec) 0.0 1,439 1,440 172,800 Flow (lbm/s) 0.0 0.0 207.3 207.3 Maximum Safeguards Time (sec) 0.0 839 840 1,679 1,680 172,800 Flow (lbm/s) 0.0 0.0 277.9 277.9 555.8 555.8 CCW Flow per RHR Heat Exchanger - lbm/s 305.5 CCW Heat Exchangers Modeled in Analysis Minimum SI Maximum SI 1

2 CCW Flow (lbm/s) 305.5 Service Water Flow per CCW Heat Exchanger (lbm/s) 275 Additional Heat Loads per CCW Heat Exchanger, Btu/hr 500,000 5-142 Table 5.3-5 DEHL Break Sequence of Events Time (sec) Event Description 0.0 Break Occurs, Reactor Trip and Loss of Offsite Power Are Assumed 0.22 Containment High Pressure Setpoint of 5.0 psig Reached 0.38 Compensated Pressurizer Pressure for Turbine Trip - 1,850 psia Reached 3.6 Low-Pressurizer Pressure SI Setpoint - 1,700 psia Reached 3.7 Containment High-High Pressure Setpoint of 23.0 psig Reached 7.05 Broken Loop Accumulator Begins Injecting Water 7.09 Intact Loop Accumulator Begins Injecting Water 18.0 Peak Containment Steam Temperature Occurs 18.0 Peak Containment Pressure Occurs 20.2 End of Blowdown Phase 20.2 Accumulator Mass Adjustment for Refill Period 20.2 Transient Modeling Terminated

5-143 Table 5.3-6 DEPS Break Sequence of Events (Minimum Safeguards) Time (sec) Event Description 0 Break Occurs, Reactor Trip and Loss-of-Offsite Power Are Assumed 0.3 Containment High Pressure Setpoint of 5.0 psig Reached 0.4 Compensated Pressurizer Pressure Turbine Trip - 1850 psia Reached 3.6 Containment High-High Pressure Setpoint of 24.0 psig Reached 3.6 Low-Pressurizer Pressure SI Setpoint - 1700 psia Reached 7.2 Broken Loop Accumulator Begins Injecting Water 7.3 Intact Loop Accumulator Begins Injecting Water 14.8 End of Blowdown Phase 14.8 Accumulator Mass Adjustment for Refill Period 20.1 Pumped SI Begins after 20-Second Diesel Delay 40.1 Broken Loop Accumulator Water Injection Ends 42.7 Intact Loop Accumulator Water Injection Ends 60.3 Emergency CFCUs Heat Removal Begins 75.6 CS Pump (from RWST) Begins 215.3 End of Reflood Phase 1,200 Recirculation Sequence Begins - Low-Head Pump Stopped and HHSI continues to pump from RWST 1,255 Mass & Energy Release Assumption: Broken Loop Steam Generator Equilibration 1,440 Containment Spray Terminated 1,440 LH Flow from Sump Begins 1,611 Mass & Energy Release Assumption: Intact Loop Steam Generator Equilibration 2,520 HHSI Pump Stopped 3,600 Mass & Energy Release Assumption: Both Steam Generators Equilibrate to 14.7 psia 3,601 Peak Containment Pressure Occurs 86,400 Containment Steam Temperature at 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> = 196°F 172,800 Transient Modeling Terminated

5-144 Table 5.3-7 DEPS Break Sequence of Events (Maximum Safeguards) Time (sec) Event Description 0 Break Occurs, Reactor Trip and Loss-of-Offsite Power Are Assumed 0.3 Containment High Pressure Setpoint of 5.0 psig Reached 0.4 Compensated Pressurizer Pressure Turbine Trip - 1,850 psia Reached 3.6 Containment High-High Pressure Setpoint of 24.0 psig Reached 3.6 Low-Pressurizer Pressure SI Setpoint - 1,700 psia Reached 7.2 Broken Loop Accumulator Begins Injecting Water 7.3 Intact Loop Accumulator Begins Injecting Water 14.1 Peak Containment Pressure Occurs 14.8 End of Blowdown Phase 14.8 Accumulator Mass Adjustment for Refill Period 20.1 Pumped SI Begins after 20-Second Diesel Delay 40.4 Broken Loop Accumulator Water Injection Ends 43.0 Intact Loop Accumulator Water Injection Ends 60.4 Emergency CFCUs Heat Removal Begins 75.6 CS Pump (from RWST) Begins 171.5 End of Reflood Phase 600 Recirculation Sequence Begins - Train 1 of SI Pumps Stopped - Train 2 continues pumping from RWST 840 Recirc Flow from Sump Begins from Train 1 941.8 Mass & Energy Release Assumption: Broken Loop Steam Generator Equilibration 1,440 Containment Spray (from RWST) Terminated 1,440 2nd Train of SI Pumps from RWST Stopped 1,463.2 Mass & Energy Release Assumption: Intact Loop Steam Generator Equilibration 1,680 Recirc Flow from Sump Begins from Train 2 3,600 Mass & Energy Release Assumption: Both Steam Generators Equilibrate to 14.7 psia 86,400 Containment Gas Temperature at 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> = 189.6°F 172,800 Transient Modeling Terminated

5-145 Table 5.3-8 LOCA Containment Integrity Response Results (Loss-of-Offsite Power Assumed)

Case Peak Press. (psig) Peak Gas Temp. (°F) Peak Liner Temp. (°F) Pressure (psig) @ 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> Temperature (°F) @ 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> DEHL 42.9 @ 18.0 sec 265.0 @ 18.0 sec Not applicable Not applicable Not applicable DEPS Min SI Framatome RSGs 43.5 @ 3,601 sec 265.1 @ 1,611 sec 262.2 @ 3,601 sec 12.2 @ 86,400 sec 195.9 @ 86,400 sec DEPS Max SI Framatome RSGs 40.8 @ 14.1 sec 261.9 @ 730.2 sec 255.7 @ 3,602 sec 11.2 @ 86,400 sec 189.6 @ 86,400 sec

5-146 Prairie Island Containment Integrity AnalysisContainment Pressure 0 5 10 15 20 25 30 35 40 451101001000100001000001000000Time (sec)Pressure (psig)PSL minSIPSL maxSIDEHL Figure 5.3-1. LOCA Containment Integrity Analysis - Containment Pressure Response Prairie Island Containment Integrity AnalysisContainment Gas Temperature1701801902002102202302402502602701101001000100001000001000000Time (sec)Temperature (°F)PSL minSIPSL maxSIDEHL Figure 5.3-2. LOCA Containment Integrity Analysis - Containment Gas Temperature Response 5-147 Prairie Island Containment Integrity AnalysisContainment Sump Temperature120140160180 2002202402601101001000100001000001000000Time (sec)Temperature (°F)PSL minSIPSL maxSIDEHL Figure 5.3-3. LOCA Containment Integrity Analysis - Containment Sump Temperature Prairie Island Containment Integrity AnalysisContainment Liner Temperature1701801902002102202302402502602701101001000100001000001000000Time (sec)Temperature (°F)PSL minSI ShellPSL minSI DomePSL maxSI ShellPSL maxSI Dome Figure 5.3-4. LOCA Containment Integrity Analysis: - Containment Liner Temperature 5-148 5.4 RUPTURES IN SECONDARY SIDE PIPING This section addresses analyses that determine the mass and energy releases from secondary side pipe ruptures, such as steam line break. Subsection 5.4.1 addresses breaks inside containment while subsection 5.4.2 addresses breaks outside containment. Also addressed within subsection 5.4.1 is the containment response to the steam line break mass and energy release.

5.4.1 Steam

Line Break Inside Containment The steam line break mass and energy release inside containment and containment integrity analyses were analyzed by Westinghouse for the 422V+ Fuel Transition Program. 5.4.1.1 Introduction and Background Steam line ruptures occurring inside a reactor containment structure may result in significant releases of high-energy fluid to the containment environment that could produce high-pressure conditions for extended periods of time. The magnitude of the releases following a steam line rupture is dependent upon the plant initial operating conditions and the size of the rupture as well as the configuration of the plant steam system and the containment design. There are competing effects for variations in the postulated accident scenario, which makes it difficult to determine the worst cases for containment pressure following a steam line break. Therefore, the steam line break and containment response analysis considers a spectrum of cases that vary the break size, initial power condition, and the postulated single failure. The following subsections identify the major plant assumptions, the analysis methodology, the selection of cases, and the results of the analysis. 5.4.1.2 Input Parameters and Assumptions This subsection summarizes key input assumptions that affect the plant response to the postulated steam line breaks. The areas discussed below are the plant protection logic, the RCS, the secondary side system, and the containment. 5.4.1.2.1 Protection Logic and Setpoints The pertinent signals and setpoints that are actuated in these analyses are summarized below. The first SI signal is generated by either a low steam line pressure signal or a high (Hi) containment pressure signal. The low steam line pressure is the first signal that would occur for large-break cases while high containment pressure would be the first signal for smaller breaks. The low steam line pressure safety analysis limit setpoint is 500 psia, with a dynamic compensation lead/lag of 12/2. The high-1 (Hi) containment pressure safety analysis limit setpoint is 19.7 psia. After appropriate electronic and mechanical delays, the SI signal is credited to cause: Start of SI pumps Reactor trip 5-149 Start of AFW pumps Feedwater isolation (including closure of valves and tripping of pumps) Start of containment fan coolers The containment spray pumps are started due to the high-2 (Hi-Hi) containment pressure signal. The safety analysis limit for the setpoint is 38.7 psia. 5.4.1.2.2 RCS Assumptions The RCS determines the amount of energy that gets transferred to the secondary side which is an important element in determining the rate of the steam line break mass and energy release. The major features of the primary side analysis model are summarized below: The assumed NSSS power is 1,690 MWt. RCS average temperature is the full-power nominal value of 560.0°F plus an uncertainty of 4.0°F. Continued operation of the RCPs maintains a high heat transfer rate to the steam generators. (No LOOP) The model includes consideration of the heat that is stored in the RCS metal. Reverse heat transfer from the intact steam generator to the RCS coolant is modeled as the temperature in the RCS falls below the steam generator fluid temperature. Core residual heat generation is assumed based on the 1979 ANS decay heat plus 2 model (Reference 35). Conservative core reactivity coefficients corresponding to end-of-cycle conditions with the most reactive rod stuck out of the core are assumed. This maximizes the reactivity feedback effects as the RCS cools down as a result of the steam line break. All cases have credited a minimum shutdown margin of 1.7-percent k, assuming the most reactive rod does not insert. Minimum flow rates are modeled from ECCS injection, to conservatively minimize the amount of boron that provides negative reactivity feedback. The boron concentration upstream of the check valve in the SI line is credited at a conservatively low boron concentration based on the Technical Specification minimum RWST concentration (2,600 ppm). No SGTP is assumed to maximize the primary-to-secondary heat transfer rate. 5.4.1.2.3 Secondary Side Assumptions This subsection summarizes the major input assumptions associated with the steam generator, the main feedwater system, the auxiliary feedwater system, and the steam line.

5-150 Initial Steam Generator Inventory The steam generator level initial condition is based on the assumption that the plant is operating at nominal programmed level when the accident is initiated. The nominal level is 44-percent narrow range span (NRS) between full-power and 20-percent power, ramping down to 33-percent NRS at zero-power. The control systems are assumed to function properly to maintain plant parameters at the programmed values, with any deviations being temporary in nature or caused by a plant transient, which need not be considered as an initial condition for normal operation. When operated in a manual mode, it is assumed that the operators control to the same desired programs, as supported by indications and alarms that alert them to deviations. Because a high initial steam generator inventory is conservative for this analysis, instrument and process level uncertainties and biases are added to the programmed level to account for the potential that the actual level may be higher than the indicated level. In addition to uncertainties and biases, it is assumed that the initial steam generator water level is at the nominal high level deviation alarm, which is 5-percent NRS above the nominal program. Main Feedwater System The rapid depressurization that occurs following a steam line break typically results in large amounts of water being added to the steam generators through the main feedwater system. A rapid-closing main feedwater regulator valve (FRV) near each steam generator will limit the feedwater addition. Prior to isolation, the main feedwater flow rate to the faulted steam generator is maximized to be conservative, since it increases the water mass inventory that will be converted to steam and released from the break. Following the initiation of the steam line break, the FRV on the faulted loop is assumed to quickly open in response to the increased steam flow. For small breaks , the feedwater flow rate increases as it is assumed that the feedwater control system instantly responds to the increased steam flow rate. For large breaks, it would be unrealistic for the feedwater flow rate to equal the steam flow rate because this would exceed the capacity of the main feedwater system. The double-ended rupture cases have more detailed main feedwater modeling based on maximum pumped flow rates, assuming the intact loop FRV is fully closed. The closure of the FRV on the faulted loop terminates the addition of pumped main feedwater to the faulted steam generator. The feedwater isolation valve (FIV) is modeled with a 1.5-second electronic delay after the SI setpoint is reached and a 7.0-second valve stroke time. For additional conservatism, full feedwater flow is assumed until the valve is fully closed. Leakage of 150 gpm across the closed FRV is considered until the FIV is fully closed; the FIV has a 45.0-second stroke time. The feedwater in the unisolable feedline between the FRV and faulted steam generator is also considered in the analysis. The hot main feedwater usually reaches saturated conditions as the steam generator and feedline depressurize. The decrease in density as flashing occurs causes most of the unisolable feedwater to enter the faulted steam generator.

Auxiliary Feedwater The AFW flow rate to the faulted steam generator is conservatively high to maximize the water mass inventory that will be converted to steam and released from the break. In addition, the fluid temperature 5-151 is modeled conservatively high to maximize energy input. The motor-driven pump and turbine-driven pump are conservatively modeled as starting with full flow at the time the SI setpoint is reached, with no electronic or mechanical delay. All of the AFW is conservatively assumed to go to the faulted steam generator due to the lower pressure of the faulted steam generator compared to the intact steam generator. The flow rate to the faulted steam generator is modeled as varying depending on the pressure of the faulted steam generator. The AFW flow rate is assumed to continue until the pumps are tripped due to a low discharge pressure. Steam Line The steam line check valve in the faulted steam line is credited to prevent a prolonged reverse blowdown from the steam line header and intact steam generator. The steam remaining in the unisolable section of the steam line between the steam generator and the check valve is included in the mass and energy released from the break. Quality of the Break Effluent The break effluent is assumed to be dry, saturated steam throughout most of the transient. However, when a large double-ended break first occurs, it is expected that there will be a significant quantity of liquid in the break effluent. Entrainment of water in the blowdown discharge is a result of the swell of the steam generator two-phase mixture and flow reversal through the steam separator drains of the steam generator due to the sudden depressurization. The entrainment characteristics for large steam line breaks are not sensitive to the steam generator design. The NRC staff agreed to this position in Reference 39 for a Point Beach license amendment request. The break quality is input as a function of time, and varies depending on the init ial power level and break size. The break quality characteristics in WCAP-8822 (Reference 40) were calculated to include Model 51 steam generators, and the bounding break quality characteristics from WCAP-8822 are used for Unit 2. For the Unit 1 ANP 56/19 steam generators, conservatively drier break quality is assumed compared to the WCAP-8822 calculations used for Unit 2. 5.4.1.2.4 Containment Assumptions The GOTHIC containment model and input assumptions for Prairie Island are documented in WCAP-16219-P (Reference 36). Initial conditions that are used in the containment analysis are summarized in Table 5.4-1. 5.4.1.3 Description of Analysis This subsection identifies the methods and computer codes used to calculate the steam line break mass and energy releases and the containment pressure response, along with the basis for the spectrum of cases analyzed. Separate analyses have been done for Unit 1 and Unit 2 due to the different steam generator models.

5-152 5.4.1.3.1 Mass and Energy Release Methodology and Computer Code The Westinghouse steam line break mass and energy release methodology is documented in WCAP-8822, "Mass and Energy Releases Following a Steam Line Rupture" (Reference 40) and was approved by the NRC (Reference 41). WCAP-8822 forms the basis for the assumptions used in the calculation of the mass and energy releases resulting from a steam line rupture. WCAP-8822 used MARVEL as the mass/energy release system code. This was subsequently replaced by LOFTRAN (References 42 and 43), which has been used in this analysis. 5.4.1.3.2 Containment Response Methodology and Computer Code The containment integrity analysis uses the GOTHIC computer code. The Prairie Island GOTHIC containment evaluation model consists of a single lumped-parameter node; the DLM is used for heat transfer to all structures in the containment. The model was described in WCAP-16219-P (Reference 36). This evaluation model was approved by the NRC in Reference 37. The model and sample analyses in WCAP-16219-P used Version 7.1 patch 1 of GOTHIC 7 (QA). The current analysis has used the most recent release of GOTHIC, Version 7.2a. Changes in the GOTHIC code versions are detailed in Appendix A of the GOTHIC User Manual (Reference 38). The most recently released version of the code is used to ensure that any corrections to coding errors are addressed; no new user options are implemented and the input model is consistent with that approved in Reference 37. 5.4.1.3.3 Case Definitions and Single Failures There are many factors that influence the quantity and rate of the mass and energy release from the steam line. To encompass these factors, a spectrum of cases varies the break size, the initial power level, and the single failure. This subsection summarizes the basis of the cases that have been defined for Prairie Island.

The power level at which the plant is operating when the steam line break is postulated can cause different competing effects that make it difficult to pre-determine a single limiting case. For example, at higher power levels there is less initial water/steam in the steam generator, which is a benefit. However, at a higher power level there is a higher initial feedwater flow rate, higher feedwater temperature, higher decay heat, and there is a higher rate of heat transfer from the primary side, which are all penalties. Therefore, cases consider initial power levels varying from full power to zero power. The specific initial power levels that are analyzed are 100 percent, 70 percent, 30 percent, and 0 percent as presented in WCAP-8822. Variations in the break size are also analyzed. The largest possible break is a DER immediately downstream of the steam generator outlet. For Unit 1, the effective forward break area is limited by the 1.4 ft 2 cross-sectional area of the integral flow restrictor in the ANP 56/19 steam generator. For Unit 2, the flow restrictor is in the steam line, creating an effective break area of 1.4 ft 2 only if the break occurs downstream of the flow restrictor. Upstream of the flow restrictor the effective break area is 4.6 ft

2. The DER cases credit liquid entrainmen t in the initial blowdown phase.

5-153 A smaller break size without crediting any liquid entrainment is also analyzed. This break size relies on the high containment pressure signal as the first SI signal, encompassing the purpose of the split breaks defined in WCAP-8822. Several single failures can be postulated that would impair the performance of various steam line break protection systems. The single failures either reduce the heat removal capacity of the containment safeguards systems or increase the energy release from the steam line break. The single failures that have been postulated for Prairie Island are summarized below. The analysis cases separately consider each single failure at each initial power level.

1) Containment Safeguards Failure The containment safeguards failure is the loss of one safeguards train. The main impact is that the active containment heat removal is reduced with the loss of one train of fan coolers and one containment spray pump. The failure also causes the loss of one train of safety injection that is assumed in the mass and energy release analysis.

2) AFW Runout Protection Failure Prairie Island has AFW runout protection that trips the AFW pumps due to a low discharge pressure. The failure of the runout protection causes the continuation of the AFW pumps when they would have otherwise tripped. This failure does not impact the AFW flow rate, only the duration that the AFW is delivered to the faulted steam generator. Operator action is credited after 10 minutes to re-align the AFW system to terminate the addition of AFW to the faulted steam generator.

3) FRV Failure The Prairie Island main feedwater system contains tw o feedwater valves in seri es on each loop-specific feedline section near each steam generator. The main FRVs are the fast-closing, primary method credited for terminating feedwater addition to the steam generators during a steam line break. One of the postulated single failures is for the FRV in the faulted loop to fail open. If this happens, the FIV is credited to close. Additional feedwater enters the faulted steam generator because the closure time of the FIV is slower than the FRV stroke time. The slower closure of the FIV allows more feedwater to flash and enter the faulted steam generator. The trip of the main feedwater pumps and condensate pumps is credited when the FRV fails open. Thus, the main mechanism for additional main feedwater to the faulted steam generator is the flashing that may occur upstream of the FRV, prior to the closure of the FIV. 5.4.1.4 Acceptance Criteria The main steam line break is classified as an ANS Condition IV event, an infrequent fault. The containment response to a steam line break is analyzed to ensure that the containment pressure remains below the containment design pressure of 46.0 psig.

5-154 5.4.1.5 Results and Conclusions The limiting containment pressure case is the largest DER initiated from 30-percent power with a containment safeguards failure. For Unit 1, the limiting case is a 1.4 ft 2 break, resulting in a peak containment pressure of 44.2 psig. For Unit 2, the limiting case is a 4.6 ft 2 break, resulting in a peak containment pressure of 45.9 psig. The peak containment temperature case for both units is a 0.7 ft 2 break initiated from full power with a containment safeguards failure postulated. Seque nces of key actuations are listed in Table 5.4-2 and Table 5.4-3 for the peak containment pressure and the peak containment temperature cases, respectively. Plots of the mass release rate, the enthalpy of the break, and the peak containment pressure, and peak containment temperature transients are contained in Figures 5.4-1 through 5.4-6. This analysis has demonstrated that the containment pressure remains below the containment design pressure throughout the transient for a postulated secondary system pipe rupture. Thus, the containment integrity has been demonstrated and all applicable acceptance criteria are met.

5.4.2 Mass and Energy Releases Outside Containment High-energy line breaks outside containment are analyzed to support compartment pressurization calculations, compartment and/or equipment temperature calculations, and compartment flood levels. Separate mass and energy release calculations are done to support the short-term pressurization analysis and the long-term temperature analysis. The characteristics of the fuel are not modeled in the analyses of the short-term secondary side mass and energy release. Therefore the short-term mass and energy releases are not impacted by the 422V+ fuel transition. For the long-term steam line break mass and energy release outside containment analysis, a sensitivity case was done to determine the impact of the 422V+ fuel. The sensitivity case addressed reactivity modeling, fuel geometry, pressure drop, heat capacity and fuel mass. Based on this sensitivity study, it was determined that the 422V+ fuel does not adversely impact the mass and energy releases outside containment.

5-155 Table 5.4-1 Initial Containment Conditions Parameter Value Containment Net Free Volume (ft

3) 1,320,000 Initial Containment Temperature (°F) 120.0 Initial Containment Pressure (psia) 16.7 (1) Initial Relative Humidity (%) 30 (1) RWST/Containment Spray Water Temperature (°F) 120 Note: 1. Peak containment vapor temperature cases assume an initial containment pressure of 14.2 psia and 100% humidity since this results in higher peak temperatures.

Table 5.4-2 Sequence of Events for Steam Line Break 30% Initial Power, Containment Safeguards Failure (Peak Containment Pressure Cases)

Time (sec) Event Unit 1 1.4 ft 2 DER Unit 2 4.6 ft 2 DER Break Occurs 0.0 0.0 Low Steam Line Pressure SI Setpoint Reached 0.2 0.2 Auxiliary Feedwater Starts 0.2 0.2 Hi Containment Pressure Setpoint Reached 0.6 0.6 Reactor Trip (rod motion starts) 2.2 2.2 FRV on Faulted Loop Closes (with leakage) 8.7 8.7 Hi-Hi Containment Pressure Setpoint Reached 32.7 12.3 FIV on Faulted Loop Fully Closes 46.7 46.7 Containment Fan Coolers Start 61.1 61.1 Containment Spray Pump Starts 85.7 72.7 AFW Terminated to Faulted SG 98.9 51.0 SG Tubes Start to Uncover 155.9 66.0 Peak Containment Pressure Occurs 211.0 91.0 Break Releases Stop 227.8 95.6

5-156 Table 5.4-3 Sequence of Events for Steam Line Break Full Initial Power, Containment Safeguards Failure (Peak Containment Vapor Temperature Cases)

Time (sec) Event Unit 1 0.7 ft 2 Break Unit 2 0.7 ft 2 Break Break Occurs 0.0 0.0 Hi Containment Pressure Setpoint Reached 8.3 11.2 Hi Containment Pressure Setpoint Credited as First SI Setpoint in Mass/Energy Release Analysis 9.0 11.6 Auxiliary Feedwater Starts 9.0 11.6 Reactor Trip (rod motion starts) 11.0 13.1 FRV on Faulted Loop Closes (with leakage) 17.5 19.6 FIV on Faulted Loop Fully Closes 55.5 57.6 Containment Fan Coolers Start 68.8 71.7 Hi-Hi Containment Pressure Setpoint Reached 74.8 104.3 Peak Containment Temperature Occurs 86.0 88.0 Containment Spray Pump Starts 127.8 157.3 AFW Terminated to Faulted SG 233.0 210.0 SG Tubes start to Uncover 316.0 331.4 Break Releases Stop 397.6 425.0

Unit 1, 1.4 ft2 DER Unit 2, 4.6 ft2 DER 100 1 20 202 300 Time (s) Figure 5.4-1.

Steam Line Break Flow Rate, 30% Power, Containment Safeguards Failure (Peak Containment Pressure Cases) --a - Unit 1, 1.4 ft2 DER Unit 2, 4.6 FtZ DER o 50 1 ao 150 207 2511 300 Time (s) Figure 5.4-2.

Steam Line Break Enthalpy, 30% Power, Containment Safeguards Failure (Peak Containment Pressure Cases)

Unit 1, 0.7 ft2 hrcok Unit 2, fl.7 Ft2 hrcok 0 200 XU 400 Time (s) Figure 5.4-3. Steam Line Break Flow Rate, Full-Power, Containment Safeguards Failure (Peak Containment Temperature Cases) - - - - - Unit 1, 0.7 ft2 break Unit 2, 0.7 ft2 break 0 200 30 400 500 Time (s') Figure 5.4-4. Steam Line Break Enthalpy, Full-Power, Containment Safeguards Failure (Peak Containment Temperature Cases)

---

Unit I, 1.4 Ft2 DEE Unit 2, 4.6 Ft2 DER 50 100 15il 2CO ,700 Time (s) Figure 5.4-5.

Containment Pressure, 30% Power, Containment Safeguards Failure (Peak Containment Pressure Cases)

---

Unit I, 6.7 Ft2 brenk Unit 2, 0.7 Ft2 break Figure 5.4-6.

Containment Temperature, Full-Power, Containment Safeguards Failure (Peak Containment Temperature Cases) 5-160

5.5 REFERENCES

1. WCAP-7908-A, "FACTRAN - A FORTRAN IV Code for Thermal Transients in a UO 2 Fuel Rod," Hargrove, H. G., December 1989. 2. WCAP-14882-P-A, "RETRAN-02 Modeling and Qualification for Westinghouse Pressurized Water Reactor Non-LOCA Safety Analyses," Huegel, D. S., et al., April 1999. 3. WCAP-7907-P-A, "LOFTRAN Code Description," Burnett, T. W. T., et al., April 1984. 4. WCAP-10965-P-A, "ANC: A Westinghouse Advanced Nodal Computer Code," Liu, Y. S., et al., September 1986. 5. WCAP-7979-P-A, "TWINKLE - A Multi-Dimensional Neutron Kinetics Computer Code," Risher, D. H., Jr. and Barry, R. F., June 1975. 6. WCAP-14565-P-A, "VIPRE-01 Modeling and Qualification for Pressurized Water Reactor Non-LOCA Thermal-Hydraulic Safety Analysis," Sung, Y. X., et al., October 1999. 7. NSP-01-009, "Northern States Power Company Prairie Island Units 1&2 NSAL-01-001: Rod Withdrawal Speed," February 22, 2001. 8. WCAP-7588, Revision 1-A, "An Evaluation of the Rod Ejection Accident in Westinghouse Pressurized Water Reactors Using Spatial Kinetics Methods," Risher, D. H., January 1975. 9. Letter, M. L. Chawla (USNRC) to T. J. Palmisano (NMC), "Prairie Island Nuclear Generating Plant, Units 1 and 2 Issuance of Amendments RE: Incorporate Large-Break Loss-of-Coolant Accident Analysis Using ASTRUM,' (TAC NOS. MD2567 and MD2568)," June 28, 2007. 10. WCAP-10079-P-A (Proprietary) and WCAP-10080 -A (Non-Proprietary), "NOTRUMP - A Nodal Transient Small Break and General Network Code," Meyer, P. E., August 1985. 11. WCAP-10054-P-A (Proprietary), and WCAP-10081 -A (Non-Proprietary), "Westinghouse Small Break ECCS Evaluation Model Using the NOTRUMP Code," Lee, N., et al., August 1985. 12. WCAP-10054-P-A, Addendum 2, Revision 1 (Proprietary), "Addendum to the Westinghouse Small Break ECCS Evaluation Model Using the NOTRUMP Code: Safety Injection into the Broken Loop and COSI Condensation Model," Thompson, C. M., et al., July 1997. 13. NSBU-NRC-00-5972, "NRC Report for NOTRUMP Version 38.0 Changes (Proprietary) and NRC Report for NOTRUMP Version 38.0 Changes (Non-Proprietary)," June 2000. 14. Appendix K of 10 CFR 50, "ECCS Evaluation Models," June 2000.

5-161 15. Docket Nos. 50-334 and 50-412, "Safety Evaluation By The Office Of Nuclear Reactor Regulation Related To Amendment Nos. 275 And 156 To Facility Operating License Nos. DPR-66 And NPF-73 FIRSTENERGY Nuclear Operating Company FIRSTENERGY Nuclear Generation Corp. Ohio Edison Company The Toledo Edison Company Beaver Valley Power Station, Unit Nos. 1 And 2 (Bvps-1 And 2)," Nuclear Regulatory Commission ADAMS Accession No. ML061720376, July 19, 2006. 16. Not Used 17. WCAP-11145-P-A (Proprietary) and WCAP-11372-A (Non-Proprietary), "Westinghouse Small Break LOCA ECCS Evaluation Model Generic Study with the NOTRUMP Code," 1986. 18. 10 CFR 50.46, "Acceptance Criteria for Emergency Core Cooling Systems for Light-Water Nuclear Power Reactors," August 2007. 19. LTR-NRC-06-44, "Transmittal of LTR-NRC-06-44 NP-Attachment, 'Response to NRC Request for Additional Information on the Analyzed Break Spectrum for the Small Break Loss of Coolant Accident (SBLOCA) NOTRUMP Evaluation Model (NOTRUMP-EM), Revision 1' (Non-Proprietary)," July 2006. 20. WCAP-8339-NP-A (Non-Proprietary), "Westinghouse ECCS Evaluation Model Summary," Bordelon, F. M., et al., July 1974. 21. WCAP-8471-P-A (Proprietary) and WCAP-8472-A (Non-Proprietary), "Westinghouse ECCS Evaluation Model: Supplementary Information," Bordelon, F. M., et al., April 1975. 22. Westinghouse Technical Bulletin, "Post LOCA Long Term Cooling: Boron Requirements," NSID-TB-86-08, October 31, 1986. 23. WCAP-9273-NP-A, "Westinghouse Reload Safety Evaluation Methodology," Davidson, S. L. (Ed.), et al., July 1985. 24. Letter from S. Swigart (Westinghouse) to D. Vincent (NMC), "Nuclear Management Company, Prairie Island Units 1&2 Transmittal of Proprietary and Non-proprietary version of Safety Analysis Transition Program - Licensing Report," NSP-04-6/LTR-MPG-04-14, January 30, 2004. 25. WCAP-11925, "An Evaluation of Long Term Cooling for Prairie Island," Young, M. Y., et al., September 1988. 26. WCAP-9735, Rev. 2, (Proprietary) and WCAP-9736, Rev. 1, (Non-Proprietary), "MULTIFLEX 3.0 A FORTRAN-IV Computer Program for Analyzing Thermal-Hydraulic-Structural System Dynamics Advanced Beam Model," K. Takeuchi, et al., February 1998. 27. WCAP-15029 and WCAP-15030, "Westinghouse Methodology for Evaluating the Acceptability of Baffle-Former-Barrel Bolting Distributions Under Faulted Load Conditions," January 1999.

5-162 28. WCAP-8708-P-A (Proprietary) and WCAP-8709-A (Non-Proprietary), "MULTIFLEX A FORTRAN-IV Computer Program for Analyzing Thermal-Hydraulic-Structure System Dynamics," K. Takeuchi, et al., September 1977. 29. WCAP-10325-P-A (Proprietary), WCAP-10326-A (Non-Proprietary), "Westinghouse LOCA Mass and Energy Release Model for Containment Design - March 1979 Version," May 1983. 30. Westinghouse Letter NSP-05-211, "Analysis Issues with the Current LOCA Mass and Energy Releases and the Resulting Containment Pressure Impacts for Prairie Island Units 1&2," December 5, 2005. 31. Docket No. 50-315, Amendment No. 126, "Facility Operating License No. DPR-58 (TAC No. 71062), for D. C. Cook Nuclear Plant Unit 1," June 9, 1989. 32. WCAP-8423, EPRI 294-2, "Mixing of Emergency Core Cooling Water with Steam; 1/3-Scale Test and Summary," Final Report, June 1975. 33. WCAP-8264-P-A, Revision 1, August 1975 (Proprietary), WCAP-8312-A (Non-Proprietary), "Westinghouse Mass and Energy Release Data for Containment Design." 34. Letter from Herbert N. Berkow, Director (NRC) to James A. Gresham (Westinghouse), "Acceptance of Clarifications of Topical Report WCAP-10325-P-A, 'Westinghouse LOCA Mass and Energy Release Model for Containment Design - March 1979 Version' (TAC No. MC7980)." 35. ANSI/ANS-5.1-1979, "American National Standard for Decay Heat Power in Light Water Reactors," August 1979. 36. WCAP-16219-P, "Development and Qualification of a GOTHIC Containment Evaluation Model for the Prairie Island Nuclear Generating Plants," Ofstun, R., April 2004. 37. Letter from Mahesh L. Chawla (NRC), "Prairie Island Nuclear Generating Plant, Units 1 and 2 - Issuance of Amendments Re: (TAC Nos. MC4245 and MC4246)," Docket Nos. 50-282 and 50-306, Nuclear Regulatory Commission ADAMS Accession No. ML052000046, August 2005. 38. NAI 8907-02 Rev. 17, "GOTHIC Containment Analysis Package User Manual," Version 7.2a(QA), January 2006. 39. Safety Evaluation by the Office of Nuclear Reactor Regulation Related to Amendment No. 206 to Facility Operating License No. DPR-24 and Amendment No. 211 to Facility Operating License No. DPR-27, Nuclear Management Company, LLC, Point Beach Nuclear Plant, Units 1 and 2, Docket Nos. 50-266 and 50-301, November 2002. 40. WCAP-8822 (Proprietary) and WCAP-8860 (Non-Proprietary), "Mass and Energy Releases Following a Steam Line Rupture," Land, R.E., September 1976.

5-163 41. Letter from Cecil O. Thomas (NRC), "Acceptance for Referencing of Licensing Topical Report WCAP-8821(Proprietary)/8859(Non-Proprietary), 'TRANFLO Steam Generator Code Description,' and WCAP-8822 (Proprietary)/8860 (Non-Proprietary), 'Mass and Energy Release Following a Steam Line Rupture,' " August 1983. 42. WCAP-7907-P-A, "LOFTRAN Code Description," Burnett, T.W.T., et al., April 1984. 43. WCAP-8822-S1-P-A (Proprietary) and WCAP-8860-S1-A (Non-Proprietary), "Mass and Energy Releases Following a Steam Line Rupture, Supplement 1 - Calculations of Steam Superheat in Mass/Energy Releases Following a Steamline Rupture," Osborne, M. P. and Love, D. S., September 1986.

6-1 6 MECHANICAL ANALYSIS 6.1 Reactor Internals Structural Analysis Analyses and evaluations were performed to assess the effect of 14x14 422V+ fuel on the reactor internal components at an NSSS power level of 1,690 MWt (core power of 1,683 MWt). 6.1.1 Introduction Operating a plant with fuel other than that considered in the original design requires that the interface between the reactor pressure vessel (RPV) system and the fuel be thoroughl y addressed to ensure compatibility, and to ensure that the structural integrity of the reactor internals and fuel system is not adversely affected. In addition, thermal-hydraulic analyses are required to determine plant-specific core-bypass flows, pressure drops, and upper head temperatures to provide input to the LOCA and non-LOCA safety analyses and to N SSS performance evaluations. The principal structural areas affected by changes in fuel are: Reactor internals system thermal-hydraulic performance Rod control cluster assembly (RCCA) scram performance Mechanical system evaluations 6.1.2 Thermal-Hydraulic System Evaluations Bypass Flow Analysis Bypass flow is the total amount of reactor coolant flow bypassing the core region and is not considered effective in the core heat transfer process. Variations in the size of some of the bypass flow paths, such as gaps at the outlet nozzles and the core cavity, occur during manufacturing or change due to fuel assembly changes. Plant-specific, as-built dimensions were used to demonstrate that the bypass flow limits were not exceeded. Therefore, analyses were performed to estimate core bypass flow values to either show that the design bypass flow limit for the plant will not be exceeded, or to determine a revised design core bypass flow. Fuel assembly hydraulic characteristics for the 14x14 422V+ fuel and system parameters, such as inlet temperature, reactor coolant pressure, and flow were used to determine the total core bypass flow. The calculated core bypass flow value was 4.1 percent with the thimble plugging devices removed at the FU analyzed operating RCS conditions. Therefore, the design core bypass flow value of 6.0 percent with thimble plugging devices removed was confirmed to remain bounding. Hydraulic Lift Forces A calculation was performed to estimate hydraulic lift forces on the various reactor internal components with 14x14 422V+ fuel at the FU analyzed operating parameters. This was done to show that the reactor internals assembly would remain seated and stable for all conditions.

6-2 The evaluation concluded that the Prairie Island Units 1&2 lower internals will remain seated with 14x14 422V+ fuel for the following conditions: Hot full-power normal conditions Cold zero-power normal conditions Hot pump overspeed (HPO) with seismic operating basis earthquake (OBE) upset conditions HPO without seismic OBE upset conditions Seismic OBE with hot full-flow upset conditions (without HPO) In addition, a minimum of 100,000-pound hold-down force is maintained during normal hot full-power operating conditions. These evaluations conservatively assume that no internals hold-down contribution is provided by the fuel assemblies. Momentum Flux and Fuel Rod Stability Baffle jetting is caused by a hydraulically induced instability or vibration of fuel rods induced by a high velocity jet of water. This jet may be created by high-pressure water being forced through gaps between the baffle plates that surround the core. The baffle jetting phenomenon could lead to fuel cladding damage. For Prairie Island Units 1&2 with the 14x14 422V+ fuel design, the evaluation showed that the momentum flux margins were within the design limits and, therefore, baffle jetting is not predicted for Prairie Island Units 1&2. 6.1.3 RCCA Scram Performance Evaluation The RCCAs represent perhaps the most critical interface between the fuel assemblies and the other internal components. It is imperative to show that the 14x14 422V+ fuel will not adversely affect the operation of the RCCAs, either during accident conditions or during normal operation. The analysis determined the effect of the 14x14 422V+ fuel on the limiting RCCA drop time. The maximum estimated RCCA drop time with the seismic allowance was calculated to be 1.7 seconds to the top of dashpot. This value is less than the current Technical Specification 3.1.4 limit of 1.8 seconds. 6.1.4 Mechanical System Evaluations LOCA and Seismic Loads The RPV LOCA system mathematical model of Prairie Island Units 1&2 is a three-dimensional nonlinear finite element model that represents the dynamic charac teristics of the reactor vessel and its internals in the six geometric degrees of freedom. The model consists of three concentric structural submodels connected by nonlinear impact elements and stiffness matrices, which is connected to a submodel of the control rod drive mechanisms (CRDMs) and CRDM seismic platform, tie rods, and lifting legs. The first submodel represents the reactor vessel shell and associated components. The reactor vessel is restrained by six reactor vessel supports, one at each inlet and outlet nozzle and one at each of the two support pads, 6-3 and by the attached primary coolant piping. Each reactor vessel support is modeled by a linear horizontal stiffness and a vertical impact element. The attached piping is represented by a stiffness matrix. The second submodel, represents the reactor core barrel, neutron panels, lower support plate, tie plates, and secondary core support components. This submodel is physically located inside the first, and is connected to it by a stiffness matrix at the internals support ledge. Core barrel to vessel shell impact is represented by nonlinear elements at the core barrel flange, core barrel nozzle, and lower radial support locations. The third, and innermost, submodel represents the upper support plate, guide tubes, support columns, upper and lower core plates, and fuel. The third submodel is connected to the first and second by stiffness matrices and nonlinear elements. The ANSYS computer code, which is used to determine the response of the reactor vessel and its internals, is a general purpose finite element code. In the finite element approach, the structure is divided into a finite number of members or elements. The inertia and stiffness matrices, as well as the force array, are first calculated for each element in the local coordinates. Employing appropriate transformation, the element global matrices and arrays are then computed. Finally, the global element matrices and arrays are assembled into the global structural matrices and arrays, and used for dynamic solution of the differential equation of motion for the structure. The results of reactor vessel displacements and the impact forces calculated at vessel and internals interfaces were used to evaluate the structural integrity of the reactor vessel internals. The core plate motions for two LOCA auxiliary line breaks (accumulator and pressurizer surge line) were used in the fuel grid analysis to confirm the structural integrity of the fuel. Consideration of these two limiting breaks is consistent with the existing Leak Before Break analysis for Prairie Island as summarized in Section 4.6.2.3 of Reference 1 where breaks in the primary loop piping no longer need to be considered in the structural design basis. Seismic Analyses The non-linear time-history seismic analyses of the RPV system included the development of the system finite element model and the synthesized time-history accelerations. Similar to the response during LOCA, the RPV system seismic model included submodels of the reactor vessel, nozzles, internals, fuel, CRDMs, and the head assembly upgrade package recently installed. The finite element model described for a LOCA was modified to include the fluid-structure interaction in the RPV model for the seismic safe shutdown earthquake (SSE) and the OBE time-history evaluations. The results of the system seismic analysis included time-history displacements and impact forces for all the major components. The reactor vessel displacements and the impact forces calculated at vessel and internals interfaces were used to evaluate the structural integrity of the reactor vessel and its internals. The core plate motions were used in the fuel grid analysis to confirm the structural integrity of the fuel.

6-4 RCCA Insertion Evaluation To assess the feasibility of crediting the RCCA insertion during a postulated faulted event, the loads on the guide tubes were calculated. These loads included the dynamic loads derived from the RPV system response, the acoustic loads, and the crossflow loads during postulated combined SSE and LOCA events. The evaluation showed that the maximum combined LOCA and SSE loads were within the allowable loads that were established for 14 x14 type guide tubes to ensure that the RCCA scram time would be acceptable. . Flow-Induced Vibration Flow-induced vibrations (FIVs) and reactor coolant pump (RCP)-induced vibration of pressurized water reactor internals have been studied by Westinghouse for a number of years. The objective of these studies was to show that the structural integrity and reliability of reactor internal components are acceptable for plant operating conditions. These efforts have included in-plant tests, scale-model tests, as well as tests in fabricators' shops and bench tests of components, along with various analytical investigations. The results of these scale-model and in-plant tests indicate that the vibrational behavior of two-, three-, and four-loop plants is essentially similar, and the results obtained from each of the tests complement one another and make possible a better understanding of the FIV phenomena. The evaluation performed for the Prairie Island Units 1&2 reactor internals indicate that acceptable vibration levels will exist during normal operation at the FU analyzed RCS conditions. Reactor Vessel Internals Structural Integrity Structural evaluations/calculations demonstrated that the structural integrity of reactor vessel internal components was not adversely affected either directly by the FU RCS conditions and transients or by secondary effects on reactor thermal-hydraulic or structural performance. Heat generated in reactor internal components, along with the various fluid temperature changes, resulted in thermal gradients within and between components. These thermal gradients resulted in thermal stresses and thermal growth, which must be considered in the design and analysis of the various components. The conclusion of these evaluations/calculations was that the structural integrity of the reactor vessel internals was maintained with acceptable margins of safety and fatigue usage factors, at the FU RCS conditions. 6.1.5 Conclusions Analyses/evaluations have been performed to assess the effect of the 14x14 422V+ fuel change at the FU analyzed RCS conditions. The results of these analyses/evaluations demonstrated:

6-5 The calculated core bypass flow with thimble plugs removed was confirmed to be less than the design value of 6 percent of the total vessel flow rate for the 14x14 422V+ fuel at the FU analyzed RCS conditions. The Prairie Island Units 1&2 reactor lower internals will remain seated and stable with the 14x14 422V+ fuel at the FU analyzed RCS conditions. Baffle plate momentum flux margins of safety with 14x14 422V+ fuel remained acceptable at the FU analyzed RCS conditions. The RCCA performance evaluation indicated that the current 1.8-second RCCA drop time from gripper release of the drive rod to dashpot entry was satisfied at with the 14x14 422V+ fuel at the FU analyzed RCS conditions. Acceptable vibration levels will exist during normal operation at the FU analyzed RCS conditions. The structural integrity of the reactor vessel internals was maintained with acceptable margins of safety and fatigue usage factors at the FU RCS conditions. In addition evaluations indicated that the 14x14 422V+ fuel will not adversely affect the response of reactor internals systems and components due to normal operation, transients and seismic/LOCA excitations.

6-6

6.2 REFERENCES

1. Prairie Island Updated Safety Analysis Report (USAR), Revision 29, May 2007.

7-1 7 NRC CONDITIONAL REQUIREMENTS FOR THE USE OF 422V+ FUEL, AND COMPUTER CODES UTILIZED 7.1 VANTAGE+ Conditional Requirement 1.1 Reference 1 provides acceptance for referencing of topical report WCAP-12610 "VANTAGE + Fuel Assembly Reference Core Report." As stated in Reference 1, it was concluded that WCAP-12610 provides an acceptable basis for the VANTAGE+ fuel assembly mechanical design up to a rod-average burnup level of 60 GWD/MTU. The Nuclear Regulatory Commission (NRC) approval does not include review or approval of higher level burnups as discussed in Appendix B of Reference 2. Similarly, Reference 1 does not address review or approval of the loss-of-coolant-accident (LOCA) analyses methods (Appendix F and G, Reference 3), which are discussed in Reference 4. No other conditional requirements are specified by Reference 1. Prairie Island is presently licensed for Vantage + fuel (Zirlo cladding) per WCAP-12610 as delineated in Section 3 of the USAR (USAR Reference 65). This is not a change with implementation of the 422 V+ fuel product. . References 1. Letter from A. C. Thadani (NRC) to S. R. Tritch (Westinghouse), "Acceptance for Referencing of Topical Report WCAP-12610 'VANTAGE + Fuel Assembly Reference Core Report'," July 1, 1991. 2. WCAP-12610 and Appendices A through D, "VANTAGE + Fuel Assembly Reference Core Report," Davidson, S. L., Nuhfer, D. L. (Eds.), June 1990. 3. WCAP-12610 Appendices F and G, "Appendix F - LOCA NOTRUMP Evaluation Model: ZIRLOŽ Modifications; Appendix G - Accident Evaluations LOCA Plant Specific," Kachmar, M. P., Iyengar, J., and Shimeck, D. J., December 1990. 4. Letter from A. C. Thadani (NRC) to S. R. Tritch (Westinghouse), "Acceptance for Referencing of Topical Report WCAP-12610, Appendices F, 'Appendix F - LOCA NOTRUMP Evaluation Model: ZIRLOŽ Modifications', and G, 'Appendix G - Accident Evaluations LOCA Plant Specific'," October 9, 1992.

7-2 7.2 LOCA/NOTRUMP-EM: ZIRLO MODIFICATIONS The NOTRUMP-EM consists of the NOTRUMP and LOCTA-IV computer codes. The NOTRUMP code is employed to calculate the transient depressurization of the reactor coolant system (RCS) as well as the mass and energy release of the fluid flow through the break. Among the features of the NOTRUMP code are: calculation of the thermal non-equilibrium in all fluid volumes, flow regime-dependent drift flux calculations with counter-current flooding limitations, mixture level tracking logic in multiple-stacked fluid nodes, regime-dependent drift flux calculations in multiple-stacked fluid nodes, and regime-dependent heat transfer correlations. Conditional Requirement 2.1 Reference 1 provides acceptance for referencing of licensing topical reports WCAP-12610, Appendices F, "LOCA/NOTRUMP Evaluation Model: ZIRLOŽ Modifications," and G, "LOCA Plant Specific Accident Evaluations." As stated in Reference 1, it was concluded that WCAP-12610 Appendices F and G and as clarified in Addendum 4 (Reference 2) is acceptable for referencing in WCAP-12610 licensing applications to the extent specified and under the limitations delineated in the reports and in Reference 1. Conditional Requirement 2.2 Reference 1 states: "Although ZIRLOŽ is similar to Zircaloy, the criteria of acceptance (10 CFR 50.44, 10 CFR 50.46, and 10 CFR 50, Appendix K) cited in the evaluation are specifically identified as appropriate for Zircaloy-clad fuel. Thus, the staff has concluded that exemptions are needed to allow application of those criteria to ZIRLOŽ clad fuel." Justification This requirement is no longer applicable as a result of recent federal regulation changes (Reference 3), the implementation of ZIRLOŽ clad fuel rods is justifiable under 10 CFR 50.59 and requires no prior NRC approval or exemptions. Prairie Island is presently licensed for Vantage + fuel (Zirlo cladding) per WCAP-12610 as delineated in Section 3 of the USAR (USAR Reference 65). This is not a change with implementation of the 422 V+ fuel product. . Conditional Requirement 2.3 Reference 1 states: "WCAP-12610, Appendix F, identifies the following changes in the use of the NOTRUMP model to account for ZIRLOŽ material properties: clad specific heat, high-temperature creep, rupture temperatures, and circumferential strain following rupture. NOTRUMP/LOCTA-IV retains the methodology given in 10 CFR Part 50, Appendix K, for the treatment of material properties, when prescribed by Appendix K and justified as suitably conservative. The retention of the Baker-Just equation 7-3 for the calculation of metal/water reaction rate specified in Appendix K is such a case. The staff considered each of these effects as a functional input to the analytical model and found them acceptable in the SER of July 1, 1991." (Reference 4). Conditional Requirement 2.4 Reference 1 states: "In WCAP-12610, Appendix F, Westinghouse identified that the gamma energy distribution methodology (generalized energy distribution model [GEDM]) approved for use in the latest Westinghouse version of WCOBRA/TRAC (staff SER of February 8, 1991) has been incorporated in the NOTRUMP/LOCTA-IV small-break analysis methodology for application to the VANTAGE+ fuel with ZIRLOŽ material. The staff concludes that the GEDM methodology is applicable to VANTAGE+ fuel assemblies with ZIRLOŽ material. Therefore, the staff finds that the use of the GEDM in the NOTRUMP/LOCTA-IV small-break methodology to analyze VANTAGE + fuel with ZIRLOŽ material acceptable." References 1. Letter from A. C. Thadani (NRC) to S. R. Tritch (Westinghouse), "Acceptance for Referencing of Topical Report WCAP-12610, Appendices F, 'Appendix F - LOCA NOTRUMP Evaluation Model: ZIRLOŽ Modifications', and G, 'Appendix G - Accident Evaluations LOCA Plant Specific'," October 9, 1991. 2. WCAP-12610, Addendum 4, "Additional Information for Appendices F and G of WCAP-12610 Appendix F - LOCA NOTRUMP Evaluation Model: ZIRLOŽ Modifications; Appendix G - Accident Evaluations LOCA Plant Specific," Kachmar, M. P., Nissley, M., and Tauche, W., May 1991. 3. "Use of Fuel with Zirconium-Based (Other than Zircaloy) Cladding (10 CFR 50.44, 50.46, and Appendix K to Part 50)," Federal Register, Vol. 57, No. 169, Rules and Regulations, pages 39353 and 39355, August 31, 1992. 4. Letter from A. C. Thadani (NRC) to S. R. Tritch (Westinghouse), "Acceptance for Referencing of Topical Report WCAP-12610 'VANTAGE + Fuel Assembly Reference Core Report'," July 1, 1991.

7-4 7.3 VIPRE Conditional Requirement 3.1 Reference 1 states: "Selection of the appropriate CHF correlation, DNBR limit, engineered hot channel factors for enthalpy rise and other fuel-dependent parameters for a specific plant application should be justified with each submittal." Justification The WRB-1 correlation with a 95/95 correlation limit of 1.17 was used in the departure from nucleate boiling (DNB) analyses for the Prairie Island 14x14 422V+ fuel. The use of the WRB-1 DNB correlation is based on the notification change that introduces the 14x14 422V+ mid-grid design (References 2 and 3). The basic change is reverting back to the larger OD fuel rod as in standard fuel but with a new low-pressure drop (LPD) mid-grid design. The applicability of WRB-1 to the LPD mid-grid was justified under the fuel criterion evaluation process (FCEP) (WCAP-12488-A) (Reference 4). The use of the plant-specific hot channel factors and other fuel dependent parameters in the DNB analysis for the Prairie Island 422V+ fuel were justified using the same methodologies as for previously approved safety evaluations of other Westinghouse two-loop plants using the same fuel design. Conditional Requirement 3.2 Reference 1 states: "Reactor core boundary conditions determined using other computer codes are generally input into VIPRE for reactor transient analyses. These inputs include core inlet coolant flow and enthalpy, core average power, power shape and nuclear peaking factors. These inputs should be justified as conservative for each use of VIPRE." Justification The core boundary conditions for the VIPRE calculations for the 422V+ fuel are all generated from NRC-approved codes and analysis methodologies. Conservative reactor core boundary conditions were justified for use as input to VIPRE. Continued applicability of the input assumptions is verified on a cycle-by-cycle basis using the Westinghouse reload methodology described in WCAP-9272-P-A (Reference 5). Conditional Requirement 3.3 Reference 1 states: "The NRC Staff's generic SER for VIPRE set requirements for use of new CHF correlations with VIPRE. Westinghouse has met these requirements for using WRB-1, WRB-2 and WRB-2M correlations. The 7-5 DNBR limit for WRB-1 and WRB-2 is 1.17. The WRB-2M correlation has a DNBR limit of 1.14. Use of other CHF correlations not currently included in VIPRE will require additional justification." Justification As discussed in response to Condition 3.1, the WRB-1 correlation with a limit of 1.17 was used in the DNB analyses of 422V+ fuel for Prairie Island. For conditions where WRB-1 is not applicable, the W-3 DNB correlation was used with a limit of 1.30 (1.45, for pressures between 500 psia and 1,000 psia). Conditional Requirement 3.4 Reference 1 states: "Westinghouse proposes to use the VIPRE code to evaluate fuel performance following postulated design-basis accidents, including beyond-CHF heat transfer conditions. These evaluations are necessary to evaluate the extent of core damage and to ensure that the core maintains a coolable geometry in the evaluation of certain accident scenarios. The NRC Staff's generic review of VIPRE did not extend to post CHF calculations. VIPRE does not model the time-dependent physical changes that may occur within the fuel rods at elevated temperatures. Westinghouse proposes to use conservative input in order to account for these effects. The NRC Staff requires that appropriate justification be submitted with each usage of VIPRE in the post-CHF region to ensure that conservative results are obtained." Justification For application to Prairie Island safety analysis, the usage of VIPRE in the post-critical heat flux region is limited to the peak cladding temperature calculation for the locked rotor transient. The calculation demonstrated that the peak cladding temperature in the reactor core is well below the allowable limit to prevent cladding embrittlement. VIPRE modeling of the fuel rod is consistent with the model described in WCAP-14565-P-A and included the following conservative assumptions: DNB was assumed to occur at the beginning of the transient Film boiling was calculated using the Bishop-Sandberg-Tong correlation The Baker-Just correlation accounted for heat generation in fuel cladding due to zirconium-water reaction Conservative results were further ensured with the following inputs: Fuel rod input based on the maximum fuel temperature at the given power The hot spot power factor was equal to or greater than the design linear heat rate Uncertainties were applied to the initial operating conditions in the limiting direction 7-6 References 1. Letter from T. H. Essig (NRC) to H. Sepp (Westinghouse), "Acceptance for Referencing of Licensing Topical Report WCAP-14565, 'VIPRE-01 Modeling and Qualification for Pressurized Water Reactor Non-LOCA Thermal/Hydraulic Safety Analysis,' (TAC No. M98666)," January 19, 1999. 2. Letter from Baumann, M. F. (WEPCO) to Document Control Desk (NRC), "14x14, 0.422" OD VANTAGE + (422V+) Fuel Design," NPL 97-0538, November 1997. 3. Letter from Maurer, B. F. (Westinghouse) to Wermiel, J. S. (NRC), "Fuel Criterion Evaluation Process (FCEP) Notification of Revision to 14x14 422 VANTAGE+ Design (Proprietary/Non-proprietary)," LTR-NRC-05-34 Rev. 1, October 2005. 4. WCAP-12488-A (Proprietary), "Westinghouse Fuel Criteria Evaluation Process," Davidson, S. L. (Ed.), et al., October 1994. 5. WCAP-9272-P-A, "Westinghouse Reload Safety Evaluation Methodology," S. L. Davidson (Ed.), July 1985.

7-7 7.4 FACTRAN FACTRAN calculates the transient temperature distribution in a cross-section of a metal clad UO2 fuel rod and the transient heat flux at the surface of the cladding, using as input the nuclear power and the time-dependent coolant parameters of pressure, flow, temperature, and density. The code uses a fuel model that simultaneously contains the following features: A sufficiently large number of radial space increments to handle fast transients such as a rod ejection accident Material properties that are functions of temperature and a sophisticated fuel-to-cladding gap heat transfer calculation The necessary calculations to handle post-DNB transients: film boiling heat transfer correlations, Zircaloy-water reaction, and partial melting of the fuel Conditional Requirement 4.1 Reference 1 states: "The fuel volume-averaged temperature or surface temperature can be chosen at a desired value which includes conservatisms reviewed and approved by the NRC." Justification The FACTRAN code was used in the analyses of the following transients for Prairie Island: Uncontrolled Rod Cluster Control Assembly (RCCA) Withdrawal from a Subcritical Condition Updated Safety Analysis Report (USAR) 14.4.1 and RCCA ejection (USAR 14.5.6). Initial fuel temperatures used as FACTRAN input in the RCCA ejection analysis were calculated using the NRC-approved PAD 4.0 computer code, as described in WCAP-15063-P-A (Reference 2). As indicated in WCAP-15063-P-A, the NRC has approved the method of determining uncertainties for PAD 4.0 fuel temperatures. Conditional Requirement 4.2 Reference 1 states: "Table 2 presents the guidelines used to select initial temperatures." Justification In summary, Table 2 of the Safety Evaluation Report (SER) specifies that the initial fuel temperatures assumed in the FACTRAN analyses of the following transients should be "High" and include uncertainties: loss of flow, locked rotor, and rod ejection. As discussed above, fuel temperatures were used as input to the FACTRAN code in the RCCA ejection analysis for Prairie Island. The assumed fuel temperatures, which were calculated using the PAD 4.0 computer code (Reference 2), include 7-8 uncertainties and are conservatively high. FACTRAN was not used in the loss of flow and locked rotor analyses. Conditional Requirement 4.3 Reference 1 states: "The gap heat transfer coefficient may be held at the initial constant value or can be varied as a function of time as specified in the input." Justification The gap heat transfer coefficients applied in the FACTRAN analyses are consistent with SER Table 2. For the RCCA withdrawal from a subcritical condition transient, the gap heat transfer coefficient is kept at a conservative constant value throughout the transient; a high constant value is assumed to maximize the peak heat flux (for DNB concerns) and a low constant value is assumed to maximize fuel temperatures. For the RCCA ejection transient, the initial gap heat transfer coefficient is based on the predicted initial fuel surface temperature, and is ramped rapidly to a very high value at the beginning of the transient to simulate clad collapse onto the fuel pellet. Conditional Requirement 4.4 Reference 1 states: "-the Bishop-Sandberg-Tong correlation is sufficiently conservative and can be used in the FACTRAN code. It should be cautioned that since these correlations are applicable for local conditions only, it is necessary to use input to the FACTRAN code which reflects the local conditions. If the input values reflecting average conditions are used, there must be sufficient conservatism in the input values to make the overall method conservative." Justification Local conditions related to temperature, heat flux, peaking factors, and channel information were input to FACTRAN for each transient analyzed for Prairie Island {RCCA withdrawal from a subcritical condition [Updated Safety Analysis Report (USAR 14.4.1)] and RCCA ejection (USAR 14.5.6)}. Therefore, additional justification is not required. Conditional Requirement 4.5 Reference 1 states: "The fuel rod is divided into a number of concentric rings. The maximum number of rings used to represent the fuel is 10. Based on our audit calculations we require that the minimum of 6 should be used in the analyses."

7-9 Justification At least 6 concentric rings were assumed in FACTRAN for each transient analyzed for Prairie Island (RCCA Withdrawal from a Subcritical Condition (USAR 14.4.1) and RCCA Ejection (USAR 14.5.6)). Conditional Requirement 4.6 Reference 1 states: "Although time-independent mechanical behavior (e.g., thermal expansion, elastic deformation) of the cladding are considered in FACTRAN, time-dependent mechanical behavior (e.g., plastic deformation) is not considered in the code. -for those events in which the FACTRAN code is applied (see Table 1), significant time-dependent deformation of the cladding is not expected to occur due to the short duration of these events or low cladding temperatures involved (where DNBR Limits apply), or the gap heat transfer coefficient is adjusted to a high value to simulate clad collapse onto the fuel pellet." Justification The two transients that were analyzed with FACTRAN for Prairie Island (RCCA Withdrawal from a Subcritical Condition (USAR 14.4.1) and RCCA Ejection (14.5.6)) are included in the list of transients provided in Table 1 of the SER; each of these transients is of short duration. For the RCCA Withdrawal from a Subcritical Condition transient, relatively low cladding temperatures are involved, and the gap heat transfer coefficient is kept constant throughout the transient. For the RCCA Ejection transient, a high gap heat transfer coefficient is applied to simulate cladding collapse onto the fuel pellet. The gap heat transfer coefficients applied in the FACTRAN analyses are consistent with SER Table 2. Conditional Requirement 4.7 Reference 1 states: "The one group diffusion theory model in the FACTRAN code slightly overestimates at beginning of life (BOL) and underestimates at end of life (EOL) the magnitude of flux depression in the fuel when compared to the LASER code predictions for the same fuel enrichment. The LASER code uses transport theory. There is a difference of about 3 percent in the flux depression calculated using these two codes. When [T(centerline) - T(Surface)] is on the order of 3000°F, which can occur at the hot spot, the difference between the two codes will give an error of 100°F. When the fuel surface temperature is fixed, this will result in a 100°F lower prediction of the centerline temperature in FACTRAN. This apparent nonconservatism was indicated to Westinghouse. In the letter NS-TMA-2026, dated January 12, 1979, Westinghouse proposed to incorporate the LASER-calculated power distribution shapes in FACTRAN to eliminate this non-conservatism. The use of the LASER-calculated power distribution in the FACTRAN code was found acceptable." Justification The condition of concern (T(centerline) - T(surface) on the order of 3000°F) is expected for transients that reach, or come close to, the fuel melt temperature. As this applies only to the RCCA Ejection 7-10 transient, the LASER-calculated power distributions were used in the FACTRAN analysis of the RCCA Ejection transient for Prairie Island. References 1. Letter from C. E. Rossi (NRC) to E. P. Rahe (Westinghouse), "Acceptance for Referencing of Licensing Topical Report WCAP-7908, 'Factran-A Fortran IV Code for Thermal Transients in UO2 Fuel Rod,' " September 30, 1986. 2. Foster, J. P., Sidener, S, "Westinghouse Improved Performance Analysis and Design Model (PAD 4.0)," WCAP-15063-P-A with errata, July 2000.

7-11 7.5 RETRAN RETRAN is used for studies of transient response of a pressurized water reactor (PWR) system to specified perturbations in process parameters. This code simulates a multi-loop system by a lumped parameter model containing the reactor vessel, hot and cold leg piping, reactor coolant pumps, steam generators (tube and shell sides), main steam lines, and the pressurizer. The pressurizer heaters, spray, relief valves, and safety valves may also be modeled. RETRAN includes a point neutron kinetics model and reactivity effects of the moderator, fuel, boron, and control rods. The secondary side of the steam generator uses a detailed nodalization for the thermal transients. The reactor protection system (RPS) simulated in the code includes reactor trips on high neutron flux, high neutron flux rate, overtemperature delta-T (OTT) and overpower delta-T (OPT), low RCS flow, high and low pressurizer pressure, high pressurizer level, and low-low steam generator water level. Control systems are also simulated including rod control and pressurizer pressure control. Parts of the safety injection system (SIS), including the accumulators, may also be modeled. RETRAN approximates the transient value of DNBR based on input from the core thermal safety limits. Conditional Requirement 5.1 Reference 1 states: "The transients and accidents that Westinghouse proposes to analyze with RETRAN are listed in this SER (Table 1) and the NRC staff review of RETRAN usage by Westinghouse was limited to this set. Use of the code for other analytical purposes will require additional justification." Justification The transients listed in Table 1 of the SER are: Feedwater system malfunctions Excessive increase in steam flow Inadvertent opening of a steam generator relief or safety valve Steam line break Loss of external load/turbine trip Loss of offsite power Loss of normal feedwater flow Feedwater line rupture Loss of forced reactor coolant flow Locked reactor coolant pump rotor/sheared shaft Control rod cluster withdrawal at power Dropped control rod cluster/dropped control bank Inadvertent increase in coolant inventory Inadvertent opening of a pressurizer relief or safety valve Steam generator tube rupture 7-12 The transients analyzed for Prairie Island using RETRAN are: Uncontrolled RCCA withdrawal at power (USAR 14.4.1) Excessive heat removal due to feedwater system malfunctions (USAR 14.4.6) Excessive load increase incident (USAR 14.4.7) Loss of reactor coolant flow (flow coast down and locked rotor) (USAR 14.4.8) Loss of external electrical load (USAR 14.4.9) Loss of normal feedwater (USAR 14.4.10) Loss of all AC power to the station auxiliaries (USAR 14.4.11) Steam line break (USAR 14.5.5) As each transient analyzed for Prairie Island using RETRAN matches one of the transients listed in Table 1 of the SER, additional justification is not required. Conditional Requirement 5.2 Reference 1 states: "WCAP-14882 describes modeling of Westinghouse designed 4-, 3, and 2-loop plants of the type that are currently operating. Use of the code to analyze other designs, including the Westinghouse AP600, will require additional justification." Justification The Prairie Island Nuclear Generating Plant consists of two 2-loop Westinghouse-designed units that were "currently operating" at the time the SER was written (February 11, 1999). Therefore, additional justification is not required. Conditional Requirement 5.3 Reference 1 states: "Conservative safety analyses using RETRAN are dependent on the selection of conservative input. Acceptable methodology for developing plant-specific input is discussed in WCAP-14882 and in Reference 2 [WCAP-9272-P-A]. Licensing applications using RETRAN should include the source of and justification for the input data used in the analysis." Justification The input data used in the RETRAN analyses performed by Westinghouse came from both Nuclear Management Company (NMC) and Westinghouse sources. Assurance that the RETRAN input data is conservative for Prairie Island is provided via Westinghouse's use of transient-specific analysis guidance documents. Each analysis guidance document provides a description of the subject transient, a discussion of the plant protection systems that are expected to function, a list of the applicable event acceptance criteria, a list of the analysis input assumptions (such as directions of conservatism for initial condition values), a detailed description of the transient model development method, and a discussion of the 7-13 expected transient analysis results. Based on the analysis guidance documents, conservative plant-specific input values were requested and collected from the responsible NMC and Westinghouse sources. Consistent with the Westinghouse Reload Evaluation Methodology described in WCAP-9272-P-A, the safety analysis input values used in the Prairie Island analyses were selected to conservatively bound the values expected in subsequent operating cycles. References 1 Letter from F. Akstulewicz (NRC) to H. Sepp (Westinghouse), "Acceptance for Referencing of Licensing Topical Report WCAP-14882-P, 'RETRAN-02 Modeling and Qualification for Westinghouse Pressurized Water Reactor Non-LOCA Safety Analyses,' " February 11, 1999. 2 WCAP-9272-P-A, "Westinghouse Reload Safety Evaluation Methodology," Davidson, S. L. (Ed.), et al., July 1985.

7-14 7.6 LOFTRAN Transient response studies of a PWR to specified perturbations in process parameters use the LOFTRAN computer code. This code simulates a multi-loop system by a model containing the reactor vessel, hot and cold leg piping, steam generators (tube and shell sides), the pressurizer and the pressurizer heaters, spray, relief valves, and safety valves. LOFTRAN also includes a point neutron kinetics model and reactivity effects of the moderator, fuel, boron, and rods. The secondary side of the steam generator uses a homogeneous, saturated mixture for the thermal transients. The code simulates the RPS, which includes reactor trips on high neutron flux, OTT and OPT, high and low pressurizer pressure, low RCS flow, low-low steam generator water level, and high pressurizer level. Control systems are also simulated including rod control, steam dump, and pressurizer pressure control. The SIS, including the accumulators, is also modeled. LOFTRAN can also approximate the transient value of DNB ratio based on input from the core thermal safety limits. Conditional Requirement 6.1 Reference 1 states: "LOFTRAN is used to simulate plant response to many of the postulated events reported in Chapter 15 of PSARs and FSARs, to simulate anticipated transients without scram, for equipment sizing studies, and to define mass/energy releases for containment pressure analysis. The Chapter 15 events analyzed with LOFTRAN are: Feedwater System Malfunction Excessive Increase in Steam Flow Inadvertent Opening of a Steam Generator Relief or Safety Valve Steam line Break Loss of External Load Loss of Offsite Power Loss of Normal Feedwater Feedwater Line Rupture Loss of Forced Reactor Coolant Flow Locked Pump Rotor Rod Withdrawal at Power Rod Drop Startup of an Inactive Pump Inadvertent ECCS Actuation Inadvertent Opening of a Pressurizer Relief or Safety Valve This review is limited to the use of LOFTRAN for the licensee safety analyses of the Chapter 15 events listed above, and for a steam generator tube rupture-"

7-15 Justification For Prairie Island, the LOFTRAN code was only used in the analyses of the dropped rod transient (USAR 14.4.3) and steam line break mass and energy releases (USAR 14.5.5.3.1). As each of these transients match one of the transients listed in the SER, additional justification is not required. References 1. Letter from C. O. Thomas (NRC) to E. P. Rahe (Westinghouse), "Acceptance for Referencing of Licensing Topical Report WCAP-7907-P, 'LOFTRAN Code Description,' " July 29, 1983.

7-16 7.7 TWINKLE TWINKLE is a multi-dimensional spatial neutron kinetics code. The code uses an implicit finite-difference method to solve the two-group transient neutron diffusion equations in one, two, and three dimensions. The code uses six delayed neutron groups and contains a detailed multi-region fuel-cladding- coolant heat transfer model for calculating pointwise Doppler and moderator feedback effects. The code handles up to 8,000 spatial points and performs steady-state initialization. Aside from basic cross-section data and thermal-hydraulic parameters, the code accepts as input basic driving functions such as inlet temperature, pressure, flow, boron concentration, control rod motion, and others. The code provides various outputs, such as channelwise power, axial offset, enthalpy, volumetric surge, pointwise power, and fuel temperatures. It also predicts the kinetic behavior of a reactor for transients that cause a major perturbation in the spatial neutron flux distribution. The TWINKLE licensing topical report, WCAP-7979-P-A (Reference 1), was approved by the U.S. Atomic Energy Commission (AEC) via an SER from D. B. Vassallo (AEC) to R. Salvatori (Westinghouse), dated July 29, 1974. The TWINKLE SER does not identify any conditions, restrictions, or limitations that need to be addressed for application to Prairie Island. References 1. Letter from D. B. Vassallo (AEC) to R. Salvatori (Westinghouse), "Acceptance for Referencing of Licensing Topical Report WCAP-7979-P, 'TWINKLE - A Multi-Dimensional Neutron Kinetics Computer Code,' " July 29, 1974.

7-17 7.8 ADVANCED NODAL CODE (ANC) ANC is an advanced nodal code capable of two-dimensional and three-dimensional neutronics calculations. ANC is the reference model for certain safety analysis calculations, power distributions, peaking factors, critical boron concentrations, control rod worths, reactivity coefficients, etc. In addition, three-dimensional ANC validates one-dimensional and two-dimensional results and provides information about radial (x-y) peaking factors as a function of axial position. It can calculate discrete pin powers from nodal information as well. The ANC licensing topical report, WCAP-10965-P-A (Reference 1), was approved by the NRC via an SER from C. Berlinger (NRC) to E. P. Rahe (Westinghouse), dated June 23, 1986 (Reference 1). The ANC SER does not identify any conditions, restrictions, or limitations that need to be addressed for application to Prairie Island. References 1. Letter from C. Berlinger (NRC) to E. P. Rahe (Westinghouse), "Acceptance for Referencing of Licensing Topical Report WCAP-10965-P, 'ANC - A Westinghouse Advanced Nodal Computer Code,' " June 23, 1986.

7-18 7.9 GOTHIC The Prairie Island GOTHIC containment evaluation model consists of a single lumped-parameter node; the diffusion layer model is used for heat transfer to all structures in the containment. The model was described in WCAP-16219 (Reference 1). This evaluation model was approved by the NRC in Reference 2. The model and sample analyses in WCAP-16219 used Version 7.1 patch 1 of GOTHIC 7 (QA). The long-term containment temperature analysis and steam line break containment response use the most recent release of GOTHIC, Version 7.2a. Changes in the GOTHIC code versions are detailed in Appendix A of the GOTHIC User Manual (Reference 3). All the GOTHIC 7 models used in the analyses presented within this report use the same limitations as approved in Reference 2. Minor differences that don't impact the restrictions within Reference 2 are outlined in subsections 5.3.2.3, 5.3.3.3, and 5.4.1.3.3 of this report. References 1. WCAP-16219-NP, "Development and Qualification of a GOTHIC Containment Evaluation Model for the Prairie Island Nuclear Generating Plants," Ofstun, R., April 2004. 2. Letter from Mahesh L. Chawla (NRC), "Prairie Island Nuclear Generating Plant, Units 1 and 2 - Issuance of Amendments Re: (TAC Nos. MC4245 and MC4246)," Docket Nos. 50-282 and 50-306, Nuclear Regulatory Commission ADAMS Accession No. ML052000046, August 2005. 3. NAI 8907-02 Rev. 17, "GOTHIC Containment Analysis Package User Manual," Version 7.2a(QA), January 2006.