ML031070329

R. E. Ginna Nuclear Power Plant - Response to LRA RAIs 4.2.1-1 & 4.2.2-1

ML031070329
Person / Time
Site:	Ginna
Issue date:	04/11/2003
From:	Mecredy R Rochester Gas & Electric Corp
To:	Arrighi R Document Control Desk, Office of Nuclear Reactor Regulation
References
BAW-2425, Rev 1
Download: ML031070329 (73)

Text

I Robert C. Mecredy Vice President Nuclear Operations Always at Your Servee April 11, 2003 U.S. Nuclear Regulatory Commission Document Control Desk Attn: Mr. Russell Arrighi (Mail Stop 0-12D-3)

Office of Nuclear Reactor Regulation Washington, D.C. 20555-0001

Reference:

"Request for Additional Information [RAI] for the Review of the RE. Ginna Nuclear Power Plant License Renewal Application [LRA], dated March 21, 2003"

Subject:

Response to LRA RAIs 4.2.1-1 and 4.2.2-1 R. E. Ginna Nuclear Power Plant Docket No. 50-244

Dear Mr. Arrighi:

In the referenced letter, RAI 4.2.1-1 requested RG&E to submit the equivalent margin analysis that was performed to demonstrate compliance with the Upper Shelf Energy (USE) requirements 1,

of Appendix G to IOCFR50. Enclosure 1 provides the requested report, BAW-2425, Revision E. Ginna for "Low Upper-Shelf Toughness Fracture Mechanics Analysis of Reactor Vessel of R.

Extended Life Through 54 Effective Full Power Years", June 2002.

Also in the referenced letter, RAI 4.2.2-1 requested information regarding our use of RGl.190 to perform neutron fluence calculations. Enclosure 2 provides Section 3 of WCAP-15885, Revision 0, "R.E. Ginna Heatup and Cooldown Limit Curves for Normal Operation", July 2002, which describes the radiation analysis and neutron dosimetry used for our reactor vessel calculations.

Very truly yours, Robert C. Mecredy tOo-11+

An equal opportunity employer 89 East Avenue I Rochester, NY 14649 tel (585) 546-2700 www.rge.corn E a n

xc w/enc: Mr. Russ Arrighi, Project Manager Office of Nuclear Reactor Regulation U.S. Nuclear Regulatory Commission One White Flint North 11555 Rockville Pike Rockville, MD 20852 xc: w/o enc: Mr. Robert L Clark (Mail Stop 0-8-C2)

Project Directorate I Division of Licensing Project Management Office of Nuclear Regulatory Regulation U.S. Nuclear Regulatory Commission One White Flint North 11555 Rockville Pike Rockville, MD 20852 Regional Administrator, Region I U.S. Nuclear Regulatory Commission 475 Allendale Road King of Prussia, PA 19406 U.S. NRC Ginna Senior Resident Inspector Mr. Denis Wickham Sr. Vice President Transmission and Supply Energy East Management Corporation P.O. Box 5224 Binghamton, NY 13902

ENCLOSURE I D0992044 BAW-2425. Rev. 1 June 2002 LOW UPPER-SHELF TOUGHNESS FRACTURE MECHANICS ANALYSIS OF REACTOR VESSEL OF R. E. GINNA FOR EXTENDED LIFE THROUGH '!

54 EFFECTIVE FULL POWER YEARS

. VNDOR DESIGN ANALYSS REVIEW

@Ao Memnorandum Required veppod ba A..

D Approved Memorandum Atlached  :

O Not Approved .Vndor Notified Approval of this design analysis does not relieve Prepared by Supplier from fuil Compliance with contract or Purchase ~r re lire ys.

Approved A s . / I H. P. Gunawardane By Date_________

NS&L Review /

Date- ,

(Required If pact on COL Values) C ROCHESTER GAS & ELECTRIC CORP. :.-.I-ROCHESTER. NY. l FRA-ANP Document No. 77-2425-01 Framatome ANP, Inc.

3315 Old Forest Road P. 0. Box 10935 Lynchburg, VA 24506-0935 Category-

,g,4c6.

Reviewed Page /

-1A __4___

of

LOW UPPER-SHELF TOUGHNESS FRACTURE MECHANICS ANALYSIS OF REACTOR VESSEL OF R. E. GINNA FOR EXTENDED LIFE THROUGH 54 EFFECTIVE FULL POWER YEARS BAW-2425, Rev. 1 FRA-ANP Document No. 77-2425-01 Prepared for Rochester Gas and Electric Corporation by Framatome ANP, Inc.

Lynchburg, Virginia This report is an accurate description of the low upper-shelf toughness fracture mechanics analysis performed for the reactor vessel at R. E. Ginna.

el q/1At0 H. P. Gunawardane, Engineer II Date Materials and Structural Analysis Unit This report has been reviewed and found to be an accurate description of the low upper-shelf toughness fracture mechanics analysis performed for the reactor vessel at R. E. Ginna.

K. K. Yoon, Yechnicat Consultant Date Materials and Structural Analysis Unit Verification of independent review.

, O,[ i A. D. McKim, Manager Date Materials and Structural Analysis Unit This report is approved for release.

J.-. Paijug, rqect Development Manager Date ii

EXECUTIVE

SUMMARY

Since it has been projected that the upper-shelf Charpy energy levels of reactor vessel beltline weld materials at R. E. Ginna may be less than 50 ft-lb at 54 effective full power years of service, a low upper-shelf fracture mechanics evaluation is required to demonstrate that sufficient margins of safety against fracture remain to satisfy the requirements of Appendix G to 10 CFR Part 50.

A low upper-shelf fracture mechanics analysis has been performed to evaluate the SA-847 circumferential reactor vessel weld at R. E. Ginna for ASME Levels A, B, C, and D Service Loadings, based on the evaluation acceptance criteria of the ASME Code, Section Xl, Appendix K.

The analysis presented in this report demonstrates that the limiting reactor vessel beltline weld at R. E. Ginna satisfies the ASME Code requirements of Appendix K for ductile flaw extensions and tensile stability using projected low upper-shelf Charpy impact energy levels for the weld material at 54 effective full power years of plant operation.

iii

RECORD OF REVISIONS Revision Affected Paaes Description Date 0 All Original release 05/29/02 1 All Updated analysis to conform to the 1995 Edition 06/10/02 of Appendix K to Section Xl of the ASME Code, with addenda through 1996.

Updated fluence and EFPY values based on latest fluence analysis.

Modified copper and nickel content to be consistent with licensing basis.

iv

CONTENTS Section Headinq Paqe

1. Introduction .......................................................... 1-1

2. Acceptance Criteria ........................................................ 2-1 2.1 Levels A and B Service Loadings (K-2200) ........................................................ 2-1 2.2 Level C Service Loadings (K-2300) ........................................................ 2-2 2.3 Level D Service Loadings (K-2400) ........................................................ 2-2

3. Material Properties and Reactor Vessel Design Data ................................... 3-1 3.1 J-lntegral Resistance Model for Mn-Mo-Ni/Linde 80 Welds .................... .................. 3-1 3.2 Reactor Vessel Design Data ........................................................ 3-2 3.3 Mechanical Properties for Weld Material ........................................................ 3-3 3.4 J-lntegral Resistance for SA-847 Weld Material ....................................................... 3-3

4. Analytical Methodology ......................................................... 4-1 4.1 Procedure for Evaluating Levels A and B Service Loadings ...................................... 4-1 4.2 Procedure for Evaluating Levels C and D Service Loadings ................... .................. 4-4 4.3 Temperature Range for Upper-Shelf Fracture Toughness Evaluations ..................... 4-4 4.4 Effect of Cladding Material ........................................................ 4-5

5. Applied Loads ......................................................... 5-1 5.1 Levels A and B Service Loadings ........................................................ 5-1 5.2 Levels C and D Service Loadings ................... ..................................... 5-1

6. Evaluation for Levels A and B Service Loadings ........................................................ 6-1

7. Evaluation for Levels C and D Service Loadings ...................................... 7-1

8. Summary of Results ........................................................ 8-1

9. Conclusion ......................................................... 9-1

10. References..................................................................................................................... 10-1 Appendix A. G. Wrobel (RG&E) letter to J. R. PalIjug (FRA-ANP), "RV Parameters," ......................... 11-1 v

LIST OF TABLES Page Table 3-1 Mechanical Properties for Beltline Materials ............................... ................ 3-2 Table 6-1 Flaw Evaluation for Levels A and B Service Loadings ........................................... 6-2 Table 6-2 J-lntegral vs. Flaw Extension for Levels A and B Service Loadings ....................... 6-3 Table 6-3 J-R Curves for Evaluation of Levels A and B Service Loadings ................. ............. 6-4 Table 7-1 K, vs. Crack Tip Temperature for TPSLB ............................................... 7-5 Table 7-2 Kk at 1/10 Wall Thickness ............................................... 7-5 Table 7-3 K.SCat 1/10 Wall Thickness with Aa = 0.10 in ............................................... 7-6 Table 7-4 J-lntegral vs. Flaw Extension for Levels C and D Service Loadings ....................... 7-7 Table 7-5 J-R Curves for Evaluation of Levels C and D Service Loadings ............................. 7-8 Table 7-6 Level D Service Loadings - Internal Pressure at Tensile Instability ........................ 7-9 LIST OF FIGURES Page Figure 1-1 Reactor Vessel Beltline Materials for R. E. Ginna ................................................. 1-2 Figure 2-1 Reactor Vessel Beltline Region with Postulated Longitudinal Flaw ......... ............... 2-3 Figure 2-2 Reactor Vessel Beltline Region with Postulated Circumferential Flaw ........ ........... 2-4 Figure 5-1 Steam Line Break without Offsite Power transient ................................................. 5-2 Figure 6-1 J-Integral vs. Flaw Extension for Levels A and B Service Loadings ......... .............. 6-5 Figure 7-1 K, vs. Crack Tip Temperature for Levels C and D Service Loadings .................... 7-10 Figure 7-2 J-lntegral vs. Flaw Extension for Levels C and D Service Loadings ..................... 7-11 vi

1. Introduction One consideration for extending the operational life of reactor vessels beyond their original licensing period is the degradation of upper-shelf Charpy impact energy levels in reactor vessel materials due to neutron radiation. Appendix G to 10 CFR Part 50, "Domestic Licensing of Production and Utilization Facilities," states in Paragraph IV.A.1.a that, "Reactor vessel beltline materials must have Charpy upper-shelf energy ... of no less than 75 ft-lb initially and must maintain Charpy upper-shelf energy throughout the life of the vessel of no less than 50 ft-lb, unless It is demonstrated in a manner approved by the Director, Office of Nuclear Reactor Regulation, that lower values of Charpy upper-shelf energy will provide margins of safety against fracture equivalent to those required by Appendix G of Section Xl of the ASME Code."

Materials with Charpy upper-shelf energy below 50 ft-lb are said to have low upper-shelf fracture toughness. Fracture mechanics analysis is necessary to satisfy the requirements of Appendix G to 10 CFR Part 50 for reactor vessel materials with upper-shelf Charpy impact energy levels that have dropped, or that are predicted to drop, below the 50 ft-lb requirement.

The base metal and weld materials used in the beltline regions of the R. E. Ginna reactor vessel are identified in Figure 1-1. Since it has been projected that the upper-shelf Charpy energy levels of the beltline weld materials may be less than 50 ft-lb at 54 effective full power years (EFPY) of service, a low upper-shelf fracture mechanics evaluation has been performed to satisfy the requirements of Appendix G to 10 CFR Part 50. A similar analysis is not required for the reactor vessel beltline forging materials since all applicable materials are predicted to have upper-shelf Charpy energy levels in excess of 50 ft-lb at 54 EFPY.

The present analysis addresses ASME Levels A, B, C, and D Service Loadings. For Levels A and B Service Loadings, the low upper-shelf fracture mechanics evaluation Is performed according to the acceptance criteria and evaluation procedures contained in Appendix K to Section Xl of the ASME Code [1]. The evaluation also utilizes the acceptance criteria and evaluation procedures prescribed in Appendix K for Levels C and D Service Loadings. Levels C and D Service Loadings are evaluated using the one-dimensional, finite element, thermal and stress models and linear elastic fracture mechanics methodology of Framatome ANP's PCRIT computer code to determine stress intensity factors for a worst case pressurized thermal shock transient.

1-1

Figure 1-1 Reactor Vessel Beitline Materials for R. E. Ginna (Forging) 1-2

2. Acceptance Criteria Appendix G to Section Xl of the ASME Code [1] provides analytical procedures for the prevention of non-ductile fracture in those areas of the pressure boundary that are comprised of materials with upper-shelf Charpy energy levels of at least 50 ft-lbs. These procedures utilize transition range fracture toughness curves with a fluence-based adjustment to crack tip temperature, and require that the component be operated at a sufficiently low pressure so as to preclude non-ductile failure. These same procedures, however, make no allowance when crack-tip temperatures are maintained above the transition range between cleavage and ductile type failures, where ductile tearing is the predicted mode of failure for ferritic reactor vessel materials. Accordingly, additional evaluation procedures were developed that utilize elastic-plastic fracture mechanics methodology and the concept of J-integral controlled crack growth.

Added to Section Xl of the ASME Code as Appendix K, these new analytical guidelines may be applied when crack tip temperatures are in the upper-shelf temperature region.

Acceptance criteria for the assessment of reactor vessels with low upper shelf Charpy energy levels are prescribed in Article K-2000 of Appendix K to Section Xl of the ASME Code [1].

These criteria, which apply to both longitudinal and circumferential flaws, as depicted in Figures 2-1 and 2-2, respectively, are summarized below as they pertain to the evaluation of reactor vessel weld metals.

2.1 Levels A and B Service Loadings (K-2200)

(a) When evaluating adequacy of the upper shelf toughness for the weld material for Levels A and B Service Loadings, an interior semi-elliptical surface flaw with a depth 1/4 of the wall thickness and a length six times the depth shall be postulated, with the flaw's major axis oriented along the weld of concern and the flaw plane oriented in the radial direction. Two criteria shall be satisfied:

(1) The applied J-integral evaluated at a pressure 1.15 times the accumulation pressure (P.) as defined in the plant specific Overpressure Protection Report, with a factor of safety of 1.0 on thermal loading for the plant specific heatup and cooldown conditions, shall be less than the J-integral of the material at a ductile flaw extension of 0.10 in.

(2) Flaw extensions at pressures up to 1.25 times the accumulation pressure (P.) shall be ductile and stable, using a factor of safety of 1.0 on thermal loading for the plant specific heatup and cooldown conditions.

(b) The J-integral resistance versus flaw extension curve shall be a conservative representation for the vessel material under evaluation.

2.2 Level C Service Loadings (K-2300)

(a) When evaluating the adequacy of the upper shelf toughness for the weld material for Level C Service Loadings, interior semi-elliptical surface flaws with depths up to 1/10 of the base metal wall thickness, plus the cladding thickness, with total depths not exceeding 1.0 in., and a surface length six times the depth, 2-1

shall be postulated, with the flaw's major axis oriented along the weld of concern, and the flaw plane oriented in the radial direction. Flaws of various depths, ranging up to the maximum postulated depth, shall be analyzed to determine the most limiting flaw depth. Two criteria shall be satisfied:

(1) The applied J-integral shall be less than the J-integral of the material at a ductile flaw extension of 0.10 in., using a factor of safety of 1.0 on loading.

(2) Flaw extensions shall be ductile and stable, using a factor of safety of 1.0 on loading.

(b) The J-integral resistance versus flaw extension curve shall be a conservative representation for the vessel material under evaluation.

2.3 Level D Service Loadings (K-2400)

(a) When evaluating adequacy of the upper shelf toughness for Level D Service Loadings, flaws as specified for Level C Service Loadings shall be postulated, and toughness properties for the corresponding orientation shall be used. Flaws of various depths, ranging up to the maximum postulated depth, shall be analyzed to determine the most limiting flaw depth. Flaw extensions shall be ductile and stable, using a factor of safety of 1.0 on loading.

(b) The J-integral resistance versus flaw extension curve shall be a best estimate representation for the vessel material under evaluation.

(c) The extent of stable flaw extension shall be less than or equal to 75% of the vessel wall thickness, and the remaining ligament shall not be subject to tensile instability.

2-2

Figure 2-1 Reactor Vessel Beltline Region with Postulated Longitudinal Flaw 1 -Semi-Elliptical

.. . .Flaw I

I (Not to Scale) - --

2-3

Figure 2-2 Reactor Vessel Beltline Region with Postulated Circumferential Flaw Semi-Elliptical Flaw (Not to Scale) 2-4

3. Material Properties and Reactor Vessel Design Data An upper-shelf fracture toughness material model is presented below, as well as mechanical properties for the weld material and reactor vessel design data.

3.1 J-lntegral Resistance Model for Mn-Mo-Ni/Linde 80 Welds A model for the J-integral resistance versus crack extension curve (J-R curve) required to analyze low upper-shelf energy materials has been derived specifically for Mn-Mo-Ni/Linde 80 weld materials. A previous analysis of the reactor vessels of B&W Owners Group RVWG [2]

described the development of this toughness model from a large database of fracture specimens. Using a modified power law to represent the J-R curve, the mean value of the J-integral is given by:

J = 1000 C,(a)c- exp(C 3 AaC4) with In(C1) = a, + a2 Cu(pdY + e3 T + a4 In(BN)

C2 =d, +d2 In(C1)+d3 In(BN)

C3 =d4 +d In(C 1)+d. In(BN)

C4 = -0.4489 where Aa = crack extension, in.

Cu = copper content, wt%

j = fluence at crack tip, 1018 n/cm2 T = temperature, OF BN = specimen net thickness, in.

and a, = 1.81 a2 = -1.512 a3 = -0.00151 a4 = 0.3935 a7 = 0.1236

-d,- O--0.077--:.

d2= 0.1164 d3 = 0.07222 d4= -0.08124 d5 = -0.00920 d8 = 0.05183 3-1

A lower bound (-2Se) J-R curve is obtained by multiplying J-integrals from the mean J-R curve by 0.699 [2]. It was shown in a previous low upper-shelf fracture toughness analysis performed for B&W Owners Group plants [3] that a typical lower bound J-R curve is a conservative representation of toughness values for reactor vessel beltline materials, as required by Appendix K [1] for Levels A, B, and C Service Loadings. The best estimate representation of toughness required for Level D Service Loadings is provided by the mean J-R curve.

3.2 Reactor Vessel Design Data Pertinent design data for upper-shelf flaw evaluations in the beltline region of the reactor vessel are provided below for R. E. Ginna.

Design Pressure, Pd = 2485 psig (use 2500 psig) [2]

Inside radius, R, = 66 in. (2]

Vessel thickness, t = 6.5 in. [2]

Nominal cladding thickness, t, = 0.1875 in. [4]

Reactor coolant inlet temperature, Tc = 5280 F [5]

3.3 Mechanical Properties for Weld Material The beltline region weld SA-847 has been previously determined [2] to be the limiting weld for the reactor vessel at R. E. Ginna. Mechanical properties for the base and weld materials are presented in Table 3-1.

Reactor vessel base metal: SA-508, Grade 2, Class 1 low alloy steel forging {6]

(changed from Class 2 to Grade 2, Class 1 in 1995)

Description:

3/4Ni-12Mo-1/3Cr-V [7]

Carbon content: < 0.30% 16]

Linde 80 weld flux: SA-847 [2]

3-2

Table 3-1 Mechanical Properties for Beltline Materials Temp. E Yield Strength (ay) Ultimate Strength (q,)' a Material: Base Base Weld Base Weld Base Metal Metal SA-847 Metal SA-847 Metal Source: Code Code Actual Code Actual Code

[Ref.] [7] [7] [8] [7] [8 [7 (OF) (ksi) (ksi) (ksi) (ksi) (ksi) (infin/0F) 100 27800 50.0 95.00 80.0 99.8 6.50E-06 200 27100 47.5 89.60 80.0 99.8 6.67E-06 300 26700 46.1 86.01 80.0 99.8 6.87E-06 400 26100 45.1 84.77 80.0 99.8 7.07E-06 500 25700 44.5 84.26 80.0 99.8 7.25E-06 528 25560 44.2 84.11 80.0 99.8 7.30E-06 600 25200 43.8 83.74 80.0 99.8 7.42E-06 Also, Poisson's ratio, v, is taken to be 0.3.

The ASME transition region fracture toughness curve for Kk, used to define the beginning of the upper-shelf toughness region, is indexed by the initial RTNDT of the weld material. For SA-847, Initial RTNDT = -4.80F [9]

Margin = 48.30F [10]

3.4 J-lntegral Resistance for SA-847 Weld Material Values of J-integral resistance from the upper-shelf toughness model of Section 3.1 are dependent on the temperature and fluence at the crack tip location, the copper content of the weld material, and the size (thickness) of the fracture specimen. These parameters are listed below for the reactor vessel at R. E. Ginna.

Projected inside surface fluence-at-54 EFPY, - = 5.01 x 1019 ,Vcm 2 [9] '

Copper content of SA-847 weld material, Cu = 0.25 wt% [10]

Net specimen thickness, BN = 0.8 in. [2]

The ultimate strength values of the base and weld metals given here are not used in calculations 3-3

Crack tip temperature varies with plant operation. At 100% power normal operating conditions, the temperature at the crack tip, T, is taken to be the inlet temperature, or Crack tip temperature, T = Tc = 528 OF Fluence at the crack tip is determined using the attenuation equation from Regulatory Guide 1.99, Rev. 2 [11]:

(Pt = qt is e-0.24x where

= attenuated fluence at crack tip, n/cm2 011s = fluence at inside surface, n/cm 2 x = depth into the vessel wall, in.

Values of the J-integral resistance at a ductile flaw extension of 0.10 in., Jo.,, can then be defined for the following flaw depths:

Flaw Depth Extension Total Depth Fluence J-lntegral Resistance, Jo.,

a Aa x = a + Aa ot Mean Lower Bound (in.) (in.) (in.) (1018 n/cm2 ) (Ibhn) (lb/in) t4 = 1.625 0.1 1.725 33.12 853 596 V10 = 0.650 0.1 0.750 41.85 842 589 3-4

4. Analytical Methodology Upper-shelf toughness is evaluated through use of fracture mechanics analytical methods that utilize the acceptance criteria and evaluation procedures of Section Xl, Appendix K [11, where applicable. Since the R. E. Ginna reactor vessel contains only circumferential welds in the beltline region, only circumferentially oriented flaws need be addressed in the present analysis.

4.1 Procedure for Evaluating Levels A and B Service Loadings The applied J-integral is calculated per Appendix K, paragraph K-4210 [1], using an effective flaw depth to account for small scale yielding at the crack tip, and evaluated per K-4220 for upper-shelf toughness and per K-4310 for flaw stability, as outlined below.

(1) For a circumferential flaw of depth a, the stress intensity factor due to internal pressure is calculated with a safety factor (SF) on pressure using the following:

R, 05(m)5F K1, = (SF)r41 +- .j)() F2 where F2 = 0.885+ ).233(

+0 0.20 < ) 0.50 (2) For a circumferential flaw of depth a, the stress intensity factor due to radial thermal gradients is calculated using the following:

K=Cm(CR)t2 .5F, 0 < (CR) l 100 °F/hr where for SA-508, Class 2 steels the material coefficient Cm is defined in Appendix K (1] as:

Ea Cm = (1-v)d = 0.0051,

-CR= cooldown rate (0F/hr), and - . -- -

F3 =0.1181+ 0.5353( )J 1.273(atJ + 0.6046(.t-) 0.20* ( It)0.50 4-1

(3) The effective flaw depth for small scale yielding, ae, is calculated using the following:

KXe+Ki]

(4) For a circumferential flaw of depth ae, the stress intensity factor due to internal pressure is P( r I+ )~R(ea)05 F2 where F2 =0.885+ 0.233(-+0345(!jL), 0.205 ( <*0.50 (5) For a circumferential flaw of depth ae, the stress intensity factor due to radial thermal gradients is K;, = Cm(CR)t2 5 Fs, 0 < (CR)

• 100 0F/hr where F3 = 0.1 181 + 0.5353~L -1 .273( J+ 0.6046( L 1 0.20* (!L <0.50

( t 1) ( t -) (it -) (it I)

(6) The J-integral due to applied loads for small scale yielding is calculated using the following:

J. = 0100 (K~p +Ke )

E'

-where

- d E 1_v2 4-2

(7) Evaluation of upper-shelf toughness at a flaw extension of 0.10 in. is performed for a flaw depth, a = 0.25t + 0.1 Oin.,

using SF= 1.15 P = Pe where P., is the accumulation pressure for Levels A and B Service Loadings, such that JA < Jo.

where J= the applied J-integral for a safety factor of 1.15 on pressure, and a safety factor of 1.0 on thermal loading J= the J-integral resistance at a ductile flaw extension of 0.10 in.

(8) Evaluation of flaw stability is performed through use of a crack driving force diagram procedure by comparing the slopes of the applied J-integral curve and the J-R curve. The applied J-integral is calculated for a series of flaw depths corresponding to increasing amounts of ductile flaw extension. The applied pressure is the accumulation pressure for Levels A and B Service Loadings, P.,

and the safety factor (SF) on pressure is 1.25. Flaw stability at a given applied load is verified when the slope of the applied J-integral curve is less than the slope of the J-R curve at the point on the J-R curve where the two curves intersect.

4.2 Procedure for Evaluating Levels C and D Service Loadings Levels C and D Service Loadings are evaluated using the one-dimensional, finite element, thermal and stress models and linear elastic fracture mechanics methodology of the PCRIT computer code to determine stress intensity factors. The limiting transient for the R. E. Ginna vessel is discussed In BAW-2178-14J. The Ginna Station is an older vintageOplant; therefore its UFSAR did not present primary system analyses in terms of ASME service levels and service limits. As such, the available Ginna service levels C and D transients did not directly reflect worst-case fracture mechanics conditions. However, these transients appear to be bounded by the transients provided for other Westinghouse-designed plants. The analysis of Ref. 14]

determined that for Level C loading conditions, the Turkey Point Steam Line Break without Offsite Power transient (TPSLB), which is a service level D transient, bounded all Level C transients for Westinghouse-designed plants. For Level D loading conditions, i was also determined that the TPSLB was the limiting transient. Therefore this transient will be used for the Levels C and D low upper-shelf fracture toughness analysis of the R. E. Ginna vessel.

4-3

The evaluation is performed as follows:

(1) Utilize PCRIT to calculate stress intensity factors for a semi-elliptical flaw of depth 1/10 of the base metal wall thickness, as a function of time, due to internal pressure and radial thermal gradients with a factor of safety of 1.0 on loading.

The critical time in the transient occurs at that point where the stress intensity factor most closely approaches the upper-shelf toughness curve.

(2) At the critical transient time, develop a crack driving force diagram with the applied J-integral and J-R curves plotted as a function of flaw extension. The adequacy of the upper-shelf toughness is evaluated by comparing the applied J-integral with the J-R curve at a flaw extension of 0.10 in. Flaw stability is assessed by examining the slopes of the applied J-integral and J-R curves at the points of intersection.

(3) Verify that the extent of stable flaw extension is no greater than 75% of the vessel wall thickness by determining when the applied J-integral curve intersects the mean J-R curve.

(4) Verify that the remaining ligament is not subject to tensile instability. The internal pressure p shall be less than Pi, where PI is the internal pressure at tensile instability of the remaining ligament. For a circumferential flaw, PI is given by

[12]:

PI[ 1.07o[ 1-(AC /A) ]

where ay++au 2

A =t(t + t)

Arae ACt =-sa 4

and e = surface length of crack, six times the depth, a Rm = mean radius of vessel This equation for Pi includes the effect of pressure on the flaw face. This equation is valid for internal pressures not exceeding the pressure at tensile instability caused by the applied hoop stress acting over the nominal wall thickness of the vessel. This validity limit on pressure Pi is Pi <1.07ao[k.]

4-4

4.3 Temperature Range for Upper-Shelf Fracture Toughness Evaluations Upper-shelf fracture toughness is determined through use of Charpy V-notch impact energy versus temperature plots by noting the temperature above which the Charpy energy remains on a plateau, maintaining a relatively high constant energy level. Similarly, fracture toughness can be addressed in three different regions on the temperature scale, i.e. a lower-shelf toughness region, a transition region, and an upper-shelf toughness region. Fracture toughness of reactor vessel steel and associated weld metals are conservatively predicted by the ASME initiation toughness curve, KC, in lower-shelf and transition regions. Inthe upper-shelf region, the upper-shelf toughness curve, Kjc, is derived from the upper-shelf -integral resistance model described in Section 3.1. The upper-shelf toughness then becomes a function of fluence, copper content, temperature, and fracture specimen size. When upper-shelf toughness is plotted versus temperature, a plateau-like curve develops that decreases slightly with increasing temperature. Since the present analysis addresses the low upper-shelf fracture toughness issue, only the upper-shelf temperature range, which begins at the intersection of Kc and the upper-shelf toughness curves, Kjc, is considered.

4.4 Effect of Cladding Material The PCRIT code utilized in the flaw evaluations for Levels C and D Service Loadings does not consider stresses in the cladding when calculating stress intensity factors for thermal loads. To account for this cladding effect, an additional stress intensity factor, Kidad, is calculated separately and added to the total stress intensity factor computed by PCRIT.

The contribution of cladding stresses to stress intensity factor was examined previously [4]. In this low upper-shelf fracture toughness analysis performed for B&W Owners Group Reactor Vessel Working Group plants, it was shown that the limiting weld was the Zion-1 WF-70 weld and the limiting transient was the Turkey Point Steam Line Break without Offsite Power. The Zion vessel had the highest projected fluence and was as thick or thicker than any other vessel.

The thicknesses of the reactor vessels for R. E. Ginna and Zion are 6.5' and 8.44',

respectively. The nominal cladding thickness is 3/16" for both vessels. From a thermal stress perspective, it is conservative to consider the thicker vessel. For the Zion vessel, the maximum value of KiCd8d,at any time during the transient and for any flaw depth, was determined to be 9.0 ksi~in. This bounding value is therefore used as the stress intensity factor for Kfidad in this R. E.

Ginna low upper-shelf fracture toughness analysis.

4-5

5. Applied Loads The Levels A and B Service Loadings required by Appendix K are an accumulation pressure (internal pressure load) and a cooldown rate (thermal load). Since Levels C and D Service Loadings are not specified by the Code, Levels C and D pressurized thermal shock events are reviewed and a worst case transient is selected for use in flaw evaluations.

5.1 Levels A and B Service Loadings Per paragraph K-1300 of Appendix K [1], the accumulation pressure used for flaw evaluations should not exceed 1.1 times the design pressure. Using 2.5 ksi as the design pressure, the accumulation pressure is 2.75 ksi. The cooldown rate is also taken to be the maximum required by Appendix K, 100 'F/hour.

5.2 Levels C and D Service Loadings As discussed in Section 4.2, the conservative Turkey Point Steam Line Break without Offsite Power transient (TPSLB) is used for the PCRIT analysis of Levels C and D service loadings.

Pressure and temperature time histories for this transient are shown in Figure 5-1. The PCRIT analysis of this transient was of sufficient duration to capture the peak value of stress intensity factor over time.

5-1

600 550 500 C 450 i 400 E

! 350 300 250 200 0 200 400 600 800 1000 1200 1400 Time (sec) 2500-2300-2100 1900 ff1700 -

(A 1500 1300-1100-900 700 0 200 400 600 800 1000 1200 1400 Time (see)

Figure 5-1 Steam Line Break without Offsite Power transient (TPSLB) 5-2

6. Evaluation for Levels A and B Service Loadings Initial flaw depths equal to 1/4 of the vessel wall thickness are analyzed for Levels A and B Service Loadings following the procedure outlined in Section 4.1 and evaluated for acceptance based on values for the J-integral resistance of the material from Section 3.4. The results of the evaluation are presented in Table 6-1, where it is seen that the minimum ratio of material J-integral resistance (Jo.1) to applied J-integral (J.) is 5.79 which is significantly higher than the minimum acceptable value of 1.0.

The flaw evaluation for the controlling weld (SA-847) is repeated by calculating applied J-integrals for various amounts of flaw extension with safety factors (on pressure) of 1.15 and 1.25 in Table 6-2. The results, along with mean and lower bound J-R curves developed in Table 6-3, are plotted in Figure 6-1. An evaluation line at a flaw extension 0.10 in. is also included to confirm the results of Table 6-1 by showing that the applied J-integral for a safety factor of 1.15 is less than the lower bound J-integral resistance of the material. The requirement for ductile and stable crack growth is also demonstrated by Figure 6-1 since the slope of the applied J-integral curve for a safety factor of 1.25 is considerably less than the slope of the lower bound J-R curve at the point where the two curves intersect.

6-1

ITable 6-1 Flaw Evaluation for Levels A & B Service Loadings I i Dimensional data: Material data:

R = T= 528 F

'66 in.

t= 6'. in. E= 25560 ksl a0 = 1.6250 in. v= 0.3 Aa = 0.1000 in. E'= 28088 ksi a= 1.725d in.

aft = 0.2654 ( 0. .2 a/t 0.5 )

Loading data: Geometry factors for initial flaw depth (w/o plasticity correction):

Pd = 2.50 ksi F, = 1.0529 for pressure loading and axial flaws P. = 2;75 ksi F2 = 0.9711 for pressure loading and circumferential flaws SF = 1.15 F3 = 0.1818 for thermal loading and both flaw types CR = 100 F/hr Cm = 0.0051 (ksi-hr)/(in 2-0 F) i .,*I Weld Orient. Kip

, K, ay ae a/t Fj' or F2' F3' Klq' KW J. J0o at V4 Jo.1 /J1 (ksNin' (ksNin) (ksl) (in.) (ksi'in) (ksi4in) (lbfin) (lbfin) r SA-847 C 43145 9.99 84.11 1.7464 0.2687 0.9725 0.1818 43.78 9.98 103 596 1 5.79

.I i

i I i j

I t

i

i. I t

.j IiI iiI i

6-2

Table 6-2 J-lntegral versus Flaw Extension for Levels A & B Service Loadings a= 66 In.

• P. = 2.75 ksi t= 6.5 In.  ! CR = 100 F/hr ao = 1.6250 in. I, Cm = 0.0051 (ksi-hr)/(inW-F) 2 ay= 84.11 ksi SF ' 1.15 SF = 1.25 ha a K p KitK, Kip Keea Kip' Kit' (in.) (in.) (ksi inj (ksivin) (in.) (ksblin) (kshlin) (lbrin) (ksblin) (ksivin) (in.) (ksivin) (ksivin) (Ibfin) 0.000 1.625 41.89 9.99 1.6452 42.21 9.99 97 45.54 9.99 1.6481 45.93 9.99 111 0.025 1.65 42.28 9.99 1.6705 42.60 9.99 98 45.96 9.99 1.6735 46.36 9.99 113 0.050 1.675 42.67 9.99 1.6958 43.00 9.99 100 46.38 9.99 1.6988 46.79 9.99 115 0.075 1.7 43.06 9.99 1.7211 43.39 9.99 101 46.80 9.99 1.7242 47.21 9.99 116 0.100 1.725 43.45 9.99 1.7464 43.78 9.98 103 47.23 9.99 1.7495 47.64 9.98 118 0.125 1.75 43.83 9.98 1.7717 44.17 9.98 104 47.64 9.98 1.7749 48.06 9.98 120 0.150 1.775 44.22 9.98 1.7970 44.56 9.97 106 48.06 9.98 1.8003 48.49 9.97 122 0.175 1.8 44.66 9.97 1.8223 44.95 9.97 107 48.48 9.97 1.8256 48.91 9.97 123 0.200 1.825 44.99 9.97 1.8476 45.33 9.96 109 48.90 9.97 1.8510 49.33 9.96 125 0.225 1.85 45.37 9.96 1.8730 45.72 9.95 110 49.32 9.96 1.8763 49.75 9.95 127 0.250 1.875 45.75 9.95 1.8983 46.11 9.94 112 49.73 9.95 1.9017 50.18 9.94 129 0.275 1.9 46.14 9.94 1.9236 46.50 9.93 113 50.15 9.94 1.9271 50.60 9.92 130 0.300 1.925 46.52 9.93 1.9489 46.88 9.91 115 50.56 9.93 1.9524 51.02 9.91 132 0.325 1.95 46.96 9.91 1.9742 47.27 9.90 116 50.98 9.91 1.9778 51.44 9.90 134 0.350 1.975 47.28 9.90 1.9995 47.65 9.88 118 51.39 9.90 2.0032 51.86 9.88 136 0.375 2 47.66 9.88 2.0248 48.04 9.87 119 51.80 9.88 2.0285 52.28 9.86 137 0.400 2.025 48.04 9.87 2.0501 48.42 9.85 121 52.22 9.87 2.0539 52.70 9.85 139 0.425 2.05 48.42 9.85 2.0755 48.81 9.83 122 52.63 9.85 2.0793 53.11 9.83 141 0.450 2.075 48.80 9.83 2.1008 49.19 9.81 124 53.04 9.83 2.1046 53.53 9.81 143 0.475 2.1 49.18 9.81 2.1261 49.58 9.79 125 53.46 9.81 2.1300 53.95 9.78 145 0.500 2.125 49.56 9.79 2.1514 49.96 9.76 127 53.87 9.79 2.1554 54.37 9.76 146 6-3

Table 6-3 J-R Curves for Evaluation of Levels A and B Service Loadings Plant: R. E. GINNA T= 528 F t= 6.5 In.

ao = 1.625 in.

t = 50.10 1018 n/cm2 @ inside surface Cu= 0.25 Bn = 0.80 in (na a X In C1 C1 C2 C3 J-R (lb/in)

(in.) (in.) (1018 n/cm2 ) Mean Low 0.001 1.6260 33.9124 0.34060 1.40579 0.10053 -0.09594 83 58 0.002 1.6270 33.9043 0.34062 1.40582 0.10053 -0.09594 158 110 0.004 1.6290 33.8880 0.34065 1.40587 0.10054 -0.09594 257 180 0.007 1.6320 33.8636 0.34071 1.40594 0.10054 -0.09594 351 245 0.010 1.6350 33.8392 0.34076 1.40601 0.10055 -0.09594 415 290 0.015 1.6400 33.7987 0.34084 1.40613 0.10056 -0.09594 490 342 0.020 1.6450 33.7581 0.34093 1.40626 0.10057 -0.09594 544 381 0.030 1.6550 33.6772 0.34110 1.40650 0.10059 -0.09594 622 435 0.040 1.6650 33.5965 0.34128 1.40674 0.10061 -0.09594 677 474 0.050 1.6750 33.5159 0.34145 1.40699 0.10063 -0.09594 720 503 0.070 -

1.6950 -

33.3554 0.34180 1.40747 0.10067 -0.09595 785 549 0.100 1.7250 33.1161 0.34232 1.40820 0.10073 -0.09595 = 853 596 0.120 1.7450 32.9576 0.34266 1.40869 0.10077 -0.09595 887 620 0.140 1.7650 32.7998 0.34301 1.40918 0.10081 -0.09596 917 641 0.160 1.7850 32.6427 0.34335 1.40966 0.10085 -0.09596 942 658 0.200 1.8250 32.3308 0.34404 1.41064 0.10093 -0.09597 984 688 0.250 1.8750 31.9452 0.34490 1.41185 0.10103 -0.09598 1026 718 0.300 1.9250 31.5641 0.34576 1.41307 0.10113 -0.09598 1061 742 0.350 1.9750 31.1876 0.34662 1.41428 0.10123 -0.09599 1090 762 0.400 2.0250 30.8156 0.34748 1.41549 0.10133 -0.09600 1116 780 0.450 2.0750 30.4480 0.34833 1.41670 0.10143 -0.09601 1139 796 0.500 2.1250 30.0848 0.34919 1.41791 0.10153 -0.09601 1159 810 6-4

Figure 6-1 J-Integral vs. Flaw Extension for Levels A & B Service Loadings 800 700 600 500 Levels A and a

/ l B Marginl

'E 400 0-

/___ _ LLower Bound J-R Curve 300

- - - Japp w/ SF=1.25

---

Japp w/ SF=1.15 200 l_ Evaluation Line for SF=1.15 1... .. . . .. :... . . . . . . . .

100 A

0.00 0.05 0.10 0.15 0.20 0.25 Flaw Extension, Aa (in.)

6-5

7. Evaluation for Levels C and D Service Loadings A flaw depth of 1/10 of the base metal wall thickness is used to evaluate the Levels C and D Service Loadings. Table 7-1 presents applied stress intensity factors, K,, from the PCRIT pressurized thermal shock analysis of the steam line break transient described in Section 5.2, along with total stress intensity factors after including a contribution of 9.0 ksiIin from cladding, as discussed in Section 4.4. The stress intensity factor calculated by the PCRIT code is the sum of thermal, residual stress, deadweight, and pressure terms. Table 7-1 also shows the variation of crack tip temperature with time for the TPSLB event. To determine the critical time in the transient for the Levels C and D flaw evaluation, allowable stress intensity factors are calculated for both the transition and upper-shelf toughness regions. Transition region toughness is obtained from the ASME Section Xl equation for crack initiation [13],

Kk = 33.2 + 2.806 exp[0.02(T- RTNDr+ 1000F)]

using an RTNDr value of 277.10 F from PCRIT for a flaw depth of 1/10 of the wall thickness, where:

Kc = transition region toughness, ksi'in T = crack tip temperature, OF Upper-shelf toughness is derived from the J-integral resistance model of Section 3.1 for a flaw depth of 1/10 of the wall thickness, a crack extension of 0.10 in., and a fluence value of 41.85 x 10 n/cm2 as follows:

K_ lJofE jc 000(1-v2) where KJC = upper-shelf region toughness, ksilin Jo, = J-integral resistance at Aa = 0.1 in.

Toughness values are given in Tables 7-2 and 7-3 for the transition and upper-shelf regions, respectively, as a function of temperature.

Figure 7-1 shows the variation of applied stress intensity factor, K,, transition range toughness, Kk, and upper-shelf toughness, Kjc with temperature. The small triangles on the K, curve indicate points in time at which PCRIT solutions are available. In the upper-shelf toughness range,- the Kr-curve is closest to -the lower bound -Kk curve at-3.6- minutes -into the transient.

This time is selected as the critical time in the transient at which to perform the flaw evaluation for Levels C and D Service Loadings.

Applied J-integrals are calculated for the controlling weld (SA-847) for various flaw depths in Table 7-4 using stress intensity factors from PCRIT for the steam line break transient (at 3.6 min.) and adding 9.0 ksivin to account for cladding effects. Stress intensity factors are converted to J-integrals by the plain strain relationship, Japplied (a) = 1 KtEE000 (a) (1-0) 7-1

Table 7-4 lists flaw extensions vs. applied J-integrals. As the Ginna vessel is 6.5 in. thick, the initial flaw depth of 1/10 of the wall thickness is 0.65 in. Flaw extension from this flaw depth is calculated by subtracting 0.65 in. from the built-in PCRIT flaw depths. The results, along with mean and lower bound J-R curves developed in Table 7-5, are plotted in Figure 7-2. An evaluation line is used at a flaw extension 0.10 in. to show that the applied J-integral is less than the lower bound J-integral of the material, as required by Appendix K [1]. The requirements for ductile and stable crack growth are also demonstrated by Figure 7-2 since the slope of the applied J-integral curve is considerably less than the slopes of both the lower bound and mean J-R curves at the points of intersection.

Referring to Figure 7-2, the Level D Service Loading requirement that the extent of stable flaw extension be no greater than 75% of the vessel wall thickness is easily satisfied since the applied J-integral curve intersects the mean J-R curve at a flaw extension that is only a small fraction of the wall thickness (less than 1%).

The last requirement is that the internal pressure p shall be less than Pi, the internal pressure at tensile instability of the remaining ligament. Table 7-6 gives the results of the calculations for PI for flaw depths up to 1.04 in. The calculated validity limit of Pi is 9.69 ksi. The calculated values of Pi shown in Table 7-1 exceed this validity limit for all flaw depths of concern. The internal pressure p is much less than the validity limit of Pi; therefore the remaining ligament is not subject to tensile instability.

7-2

Table 7-1 Kevs. Crack Tip Temperature for TPSLB aft =1/10 a = 0.650 in.

PCRIT Clad Total Time Temp Kisum K Kg 0.00 547.00 29.31 9.0 38.31 0.10 546.90 25.21 9.0 34.21 0.20 545.50 23.65 9.0 32.65 0.30 541.70 24.42 9.0 33.42 0.40 536.20 27.16 9.0 36.16 0.50 529.40 30.07 9.0 39.07 0.60 521.90 32.87 9.0 41.87 0.70 514.10 35.55 9.0 44.55 0.80 506.20 38.04 9.0 47.04 0.90 498.40 40.37 9.0 49.37 1.00 490.80 42.59 9.0 51.59 1.10 483.40 44.75 9.0 53.75 1.20 476.10 46.82 9.0 55.82 1.30 469.00 48.83 9.0 57.83 1.40 462.10 50.83 9.0 59.83 1.50 455.20 52.70 9.0 61.70 1.60 448.50 54.40 9.0 63.40 1.70 442.20 56.00 9.0 65.00 1.80 436.00 57.53 9.0 66.53 1.90 430.10 59.00 9.0 68.00 2.00 424.40 60.40 9.0 69A0 2.10 418.80 61.65 9.0 70.65 2.20 413.60 62.79 9.0 71.79 2.30 408.60 63.85 9.0 72.85 2.40 403.90 64.83 9.0 73.83 2.50 399.40 65.74 9.0 74.74 2.60 395.10 66.61 9.0 75.61 2.70 391.00 67.43 9.0 76A3 2.80 387.10 68.21 9.0 77.21 2.90 383.30 68.91 9.0 77.91

-3.100 - - 379.70 6951 --- 9.0 -- 78.51 3.10 376.40 70.02 9.0 79.02 3.20 373.20 70.48 9.0 79A8 3.30 370.30 70.89 9.0 79.89 3.40 367.50 71.25 9.0 80.25 3.50 364.90 71.57 9.0 80.57 3.60 362.40 71.85 9.0 80.85 3.80 357.90 72.30 9.0 81.30 4.00 353.80 72.63 9.0 81.63 7-3

Table 7-1 K,vs. Crack Tip Temperature for TPSLB (continued) alt =1/10 a = 0.650 in.

PCRIT Clad Total Time Temp Kisum K, Ken 4.20 350.10 72.91 9.0 81.91 4A0 346.70 73.10 9.0 82.10 4.60 343.60 73.24 9.0 82.24 4.80 340.80 73.32 9.0 82.32 5.00 338.20 73.35 9.0 82.35 5.20 335.80 73.34 9.0 82.34 5.40 333.50 73.29 9.0 82.29 5.60 331.40 73.22 9.0 82.22 5.80 329.40 73.12 9.0 82.12 6.00 327.50 73.01 9.0 82.01 6.20 325.80 72.88 9.0 81.88 6A0 324.10 72.74 9.0 81.74 6.60 322.40 72.59 9.0 81.59 6.80 320.90 72.42 9.0 81.42 7.00 319.40 72.25 9.0 81.25 7.20 318.00 72.06 9.0 81.06 7.40 316.70 71.86 9.0 80.86 7.60 315.40 71.66 9.0 80.66 7.80 314.10 71.45 9.0 80.45 8.00 312.90 71.23 9.0 80.23 8.20 311.80 71.04 9.0 80.04 8.40 310.70 70.84 9.0 79.84 8.70 309.10 70.53 9.0 79.53 9.00 307.60 70.22 9.0 79.22 9.50 305.20 69.67 9.0 78.67 10.00 303.10 69.04 9.0 78.04 10.50 301.00 68.16 9.0 77.16 11.00 299.10 67.13 9.0 76.13 12.00 295.70 65.17 9.0 74.17 14.00 289.70 61.36 9.0 70.36 16.00 284.50 - 57.86 9.0 66.86 19.00 277.40 53.14 9.0 62.14 22.00 271.20 49.02 9.0 58.02 25.01 265.70 45.44 9.0 54.44 29.01 259.70 41.24 9.0 50.24 33.35 255.00 37.34 9.0 46.34 7-4

Table 7-2 Kl at 1/10 Wall Thickness Kc Curve at a = 1/10T RTNDT = 277.1 F T T-RTNDT K1c (F) (ksiNin) 200 -77.1 37.6 210 -67.1 38.6 220 -57.1 39.8 230 -47.1 41.3 240 -37.1 43.1 250 -27.1 45.3 260 -17.1 47.9 270 -7.1 51.2 280 2.9 55.2 290 12.9 60.0 300 22.9 66.0 310 32.9 73.2 320 42.9 82.1 330 52.9 92.9 340 62.9 106.2 350 72.9 122.3 360 82.9 142.0 370 92.9 166.1 380 102.9 195.6 390 112.9 231.5 400 122.9 275.4 410 132.9 329.0 420 142.9 394.5 430 152.9 474.5 440 162.9 572.2 450 172.9 691.6 7-5

Table 7-3 K at 1/10 WaIl Thickness with ba = 0.10 in.

Kj, Curve with Ha = 0.10 in.

Fluence = 50.10 x 1018 n/cm2 at inside surface

= 41.85 x 1018 n/cm2 at tl0 + 0.1" Aa = 0.10 in.

Cu = 0.25 Wt-%

E = 25560 ksi V= 0.30 C 4 = -0.4489 Lower Lower Mean Bound Mean Bound T InC1 C1 C2 C3 JO.1 JO1 Kjc Kic (F) (Ibtin) (lb/in) (kshlin) (ksi4in) 200 0.82050 2.27163 0.15639 -0.10035 1195 836 183.2 153.2 250 0.74500 2.10644 0.14760 -0.09965 1133 792 178.4 149.2 300 0.66950 1.95326 0.13881 -0.09896 1074 751 173.7 145.2 350 0.59400 1.81122 0.13003 -0.09826 1019 712 169.1 141.4 400 0.51850 1.67951 0.12124 -0.09757 966 675 164.7 137.7 450 0.44300 1.55737 0.11245 -0.09688 916 640 160.4 134.1 500 0.36750 1.44412 0.10366 -0.09618 868 607 156.1 130.5 550 0.29200 1.33910 0.09487 -0.09549 823 575 152.0 127.1 600 0.21650 1.24172 0.08609 -0.09480 780 545 148.0 123.8 7-6

Table 7-4 J-lntegral vs. Flaw Extension for Levels C and D Service Loadings Time = 3.60 min. E= 25560 ksi Crack tip at t:10 t 6.5 in. v= 0.3 (alt)*40 a Aa Temp. Kisum w w Kite, Japp (in.) (in.) (F) (Ib/in) 1 0.1625 308.60 43.50 9.0 52.5 98 2 0.3250 327.20 59.20 9.0 68.2 166 3 0.4875 345.20 67.13 9.0 76.1 206 4 0.6500 0.0000 362.40 71.85 9.0 80.9 233 5 0.8125 0.1625 378.80 74.70 9.0 83.7 249 6 0.9750 0.3250 394.30 76.30 9.0 85.3 259 7 1.1375 0.4875 409.00 76.94 9.0 85.9 263 8 1.3000 0.6500 422.70 77.00 9.0 86.0 263 9 1.4625 0.8125 435.50 76.38 9.0 85.4 260 10 1.6250 0.9750 447.30 75.54 9.0 84.5 254 12 1.9500 1.3000 468.30 72.84 9.0 81.8 238 14 2.2750 1.6250 485.80 69.35 9.0 78.4 219 16 2.6000 1.9500 500.10 65.66 9.0 74.7 198 18 2.9250 2.2750 511.60 61.35 9.0 70.4 176 20 3.2500 2.6000 520.70 56.66 9.0 65.7 153 22 3.5750 2.9250 527.80 52.02 9.0 61.0 133 24 3.9000 3.2500 533.10 47.23 9.0 56.2 113 26 4.2250 3.5750 537.20 42.82 9.0 51.8 96 28 4.5500 3.9000 540.10 38.97 9.0 48.0 82 30 4.8750 4.2250 542.20 35.67 9.0 44.7 71 32 5.2000 4.5500 543.70 32.64 9.0 41.6 62 Note: At Aa = 0.10 in., Jpp = 243 lb/in.

7-7

Table 7-5 J-R Curves for Evaluation of Levels C and D Service Loadings Plant: R. E. GINNA Time = 3.60 min.

T= 362.4 F t= 6.5 in.

so = 0.65 in.

Ais = 50.10 1018 n/cm2 @ inside surface Cu = 0.25 Bn= 0.80 in Aa a In C1 C1 C2 C3 J-R (lb/in)

(in.) (in.) (1018 n/cm2 ) Mean Low 0.001 0.6510 42.8532 0.57351 1.77449 0.12764 -0.09808 83 58 0.002 0.6520 42.8429 0.57353 1.77452 0.12764 -0.09808 163 114 0.004 0.6540 42.8224 0.57357 1.77458 0.12765 -0.09808 272 190 0.007 0.6570 42.7916 0.57362 1.77468 0.12765 -0.09808 379 265 0.010 0.6600 42.7608 0.57367 1.77477 0.12766 -0.09808 454 317 0.015 0.6650 42.7095 0.57376 1.77493 0.12767 -0.09808 544 380 0.020 0.6700 42.6583 0.57385 1.77509 0.12768 -0.09808 610 427 0.030 0.6800 42.5560 0.57403 1.77541 0.12770 -0.09808 707 494 0.040 0.6900 42.4540 0.57421 1.77572 0.12772 -0.09808 777 543 0.050 0.7000 42.3522 0.57439 1.77604 0.12774 -0.09808 831 581 0.070 0.7200 42.1494 0.57474 1.77667 0.12778 -0.09809 915 640 0.100 0.7500 41.8470 0.57528 1.77762 0.12785 -0.09809 1005 703 0.120 0.7700 41.6467 0.57563 1.77825 0.12789 -0.09810 1052 735 0.140 0.7900 41.4472 0.57599 1.77888 0.12793 -0.09810 1091 763 0.160 0.8100 41.2488 0.57634 1.77952 0.12797 -0.09810 1126 787 0.200 0.8500 40.8547 0.57705 1.78078 0.12805 -0.09811 1184 828 0.250 0.9000 40.3673 0.57794 1.78236 0.12816 -0.09812 1243 869 0.300 0.9500 39.8858 0.57882 1.78394 0.12826 -0.09812 1292 903 0.350 1.0000 39.4101 0.57971 1.78551 0.12836 -0.09813 1333 932 OAOO 1.0500 38.9400 0.58059 1.78709 0.12847 -0.09814 1370 958 OA50 1.1000 38.4755 0.58147 1.78866 0.12857 -0.09815 1403 981 0.500 1.1500 38.0165 0.58235 1.79024 0.12867 -0.09816 1432 1001 7-8

Table 7.6 Level D Service Loadings - Internal Pressure at Tensile Instability flaw depth a (in.) PI (ksi) 0.065 20.32 0.130 20.29 0.195 20.25 0.260 20.19 0.325 20.11 0.390 20.03 0.455 19.94 0.520 19.84 0.585 19.73 0.650 19.62 0.715 19.50 0.780 19.37 0.845 19.25 0.910 19.12 0.975 18.98 1.040 18.84 7-9

'I-Figure 7-1 K, vs. Crack Tip Temperature for Levels C & D Service Loadings 220 200 180 160 140 120 100 80 60 40 20 0

325 375 425 475 525 575 Crack Tip Temperature ( 0F) 7-10

Figure 7-2 J-Integral vs. Flaw Extension for Levels C & D Service Loadings 1600 1400 1200 Margin, ,.

1000 C

l i l- Mean J-R Curve I...

800 7__ / /Lower

- _ _ Bound J-R Curve

.62 C / - -- Japplied for SLB at 3.60 min.

Evaluation Line for Level C 600 ILeve I Margin 400 200 a

0.00 0.05 0.10 0.15 0.20 0.25 0.30 0.35 0.40 0.45 0.50 Flaw Extension, Aa (in.)

7-11

8. Summary of Results A low upper-shelf fracture mechanics analysis has been performed to evaluate the SA-847 circumferential reactor vessel weld at R. E. Ginna for projected low upper-shelf energy levels at 54 EFPY, considering Levels A, B, C, and D Service Loadings of the ASME Code.

Evidence that the ASME Code, Section Xl, Appendix K [1] acceptance criteria have been satisfied for Levels A and B Service Loadings is provided by the following:

(1) Figure 6-1 shows that with a factors of safety of 1.15 on pressure and 1.0 on thermal loading, the applied J-integral (J1) is less than the J-integral of the 1 = 5.79 which material at a ductile flaw extension of 0.10 in. (Jo.,). The ratio Jo.11J is significantly greater than the required value of 1.0.

(2) Figure 6-1 shows that with a factors of safety of 1.25 on pressure and 1.0 on thermal loading, flaw extensions are ductile and stable since the slope of the applied J-integral curve is less than the slope of the lower bound J-R curve at the point where the two curves intersect.

Evidence that the ASME Code, Section Xl, Appendix K [1] acceptance criteria have been satisfied for Levels C and D Service Loadings is provided by the following:

(1) Figure 7-2 shows that with a factor of safety of 1.0 on loading, the applied J-integral (J1) is less than the J-integral of the material at a ductile flaw extension of 0.10 in. (Jo.,). From Tables 7-4 and 7-5, the ratio JO.,1J1 = 703/243 = 2.89, which is much greater than the required value of 1.0.

(2) Figure 7-2 shows that with a factor of safety of 1.0 on loading, flaw extensions are ductile and stable since the slope of the applied J-integral curve is less than the slopes of both the lower bound and mean J-R curves at the points of intersection.

(3) Figure 7-2 shows that flaw growth is stable at much less than 75% of the vessel wall thickness. It has also been shown that the remaining ligament is sufficient to preclude tensile instability by a large margin.

8-1

9. Conclusion The limiting R. E. Ginna reactor vessel beltline weld satisfies the acceptance criteria of Appendix K to Section Xl of the ASME Code [1 for projected low upper-shelf Charpy impact energy levels at 54 effective full power years of plant operation.

9-1

10. References

1. ASME Boiler and Pressure Vessel Code, Section Xl, 1995 Edition with Addenda through 1996.

2. BAW-2192PA, Low Upper-Shelf Toughness Fracture Mechanics Analysis of Reactor Vessels of B&W Owners Reactor Vessel Working Group For Level A & B Service Loads, April 1994.

3. BAW-2275, Low Upper-Shelf Toughness Fracture Mechanics Analysis of B&W Designed Reactor Vessels for 48 EFPY, August 1996.

4. BAW-2178PA, Low Upper-Shelf Toughness Fracture Mechanics Analysis of Reactor Vessels of B&W Owners Reactor Vessel Working Group For Level C & D Service Loads, April 1994.

5. R. E. Ginna Nuclear Power Plant, Updated Final Safety Analysis Report, Revision 16, Volume 8, April 2001

6. WCAP-7254, Rochester Gas and Electric Robert E. Ginna Unit No. 1 Reactor Vessel Radiation Surveillance Program, May 1969.

7. ASME Boiler and Pressure Vessel Code,Section II, Part D, 1995 Edition with Addenda through 1996.

8. WCAP-1 3902, Analysis of Capsule S from the Rochester Gas and Electric Corporation R. E. Ginna Reactor Vessel Radiation Surveillance Program, December 1993.

9. G. Wrobel (RG&E) letter to J. R. Paljug (FRA-ANP), "RV Parameters," dated May 24, 2002 (attached as Appendix A).

10. USNRC Reactor Vessel Integrity Database (RVID) Version 2.0.1, U. S. Nuclear Regulatory Commission, July 2000.

11. Regulatory Guide 1.99, Revision 2, U.S. Nuclear Regulatory Commission, May 1988.

12. ASME Boiler and Pressure Vessel Code, Appendix K, Section Xl, 2001 Edition.

13. EPRI NP-719-SR, T.U. Marston, Flaw Evaluation Procedures: ASME Section Xl, Electric Power Research Institute, Palo Alto, California, August 1978.

10-1

11. Appendix A - G. Wrobel (RG&E) letter to J. R. Paijug (FRA-ANP), "RV Parameters,"

MY-24-2002 16: 18 RG&E GINRA ENGINEERING 716 771 3904 P,02 I A SubsMidiry of RGS E&WW GroLp Inc.

ROCHESTER GAS ANDELECTRCCORPORATION89 EAsT AvlUF ROCHESTER. NY. 744-j007 t5-46-27 GEOROEWROBEL UOM U&:~Ro May 24,2002 To: Joe Pajug From George Wrobel St~ect: RV Paramets The reference for peal EOL neutron fluenwc at the redor vessel belie weld is WCAP-158S5, Rev. 0,Q 'E.

Gima,Hea"up and Cooldown Limit Curves for Normal Operafion", May 2002.

In Table 6 ofthatWCAP, neutro fuence (E>1.0 MeV) is 4.85 E + 19 at 52 EFPY, and S.OIE + I9 at 54 EFPY (attadhnw 1). he value of 5.5 E + 19 used in the RV EMA is consezvative wit respect to these values.

The reference for initial RT? of 4.8-F is the RG&E "RCS Pressurc and Temperature Limits Report PThR",

Rev. 3, 2/VI2001, Table PTLR-6 (attachmue 2).

11-1

r MRY-24-2m2 10: 16 RGE.E G1INN ENGINEERING 716 771 3904 14 P.03 I

l 4 Table 6 Sumnay ef Calculated Maxim=u Pressure Vessel Exposure Clad/Basc Metal Inteac Neon zenceVE> i.oMvJ Cumulative Neutron Fluence ln/cmal Operating Time

.EFF 0.0 Degres 15.0 Degrees 30.0 Dezrecs 45.0 Degree 24.8 (EOC 29) 2.68e+19 1.69e+19 1.22c+19 1.09&+19 28 2.94c+19 1.t5c+19 1.34c+19 1.20c+19 32 3.26e+19 2.05Sc+19 1.48e+19 1.33c+19 36 3.57c+19 2.25e+19 1.63e 19 1.46e+19 40 3.89e+19 2.45c+19 1.77ci19 1.60e*19 44 4.2ie+19 2.65c+19 192e+19 1.73c+19 48 4.53ce+19 2.S9e+19 2.07e+19 1.86e+19 52 4.85e+19 3.05e 19 2.21e+19 2.00c419 54 5.01e+19 3.15c+19 2.23c+19 2.06e+19 Iron Atom Displac cnts CumulaiveIron _cements Atom Di Idoal Opersting IEREY 0.0 Derees 15.0 Degrees 30.0 berecg 45.0 Degees 24.8 (EOC 29) 4.37e-02 2.25e-02 2.01-e02 1.77c.02 28 4.79e-02 3.12O02 2.20e402 I.94e-02 32 S.31e-02 3.46c.02 2.44"02 2.16e-02 36 S.52c.02 3.79c-02 2.68e-02 2.37-402 40 6.34e-02 4.12c-02 2.91.02 2.59c-02 44 6.86e-02 4.46c-02 3.15.e02 2.I0e-02 48 7.38e-02 4.79c-02 3.39e*02 3.02e-02 S2 7.89c-02 S.12c-02 3.63c-02 3.23e-02 54 S.l5c-02 5.29c-02 3.75e-02 3.34e02 11-2

mny-24-2Em2 16:1e RGRE GINN ENINEERING 716 771 3904a so-so P.4 Table PTLR - 6 Material II calculation of Adjusted Reference Temperatures al 28 EFPY for the Umnfing Reactor Vessel Paranter Vakues Opersmgn; Tffne 2 EFPY Material Cire. Weld Ckorc Weld Location 1/T 314-T 160.7 150.7 Chemstrny Factor (CF), *Fi')

Ru=nca M,1v1U nkmz (E > 1.0MeV)M) S.1l .e6 Flhenca Factor (FF) 120 1.00 liaNFT= CF x FFoF 1028 160.7 initial RTNDt (1). OF -4.8 -4.8 Mnin ,M Flb) 4&3 48.

Am . i * (OFxFF). M. .F~t3(c) 238.3 204.2 (a) values from Table PTLR - 3.

(b) Value calculated using Table PTLR - 6 values-(c) Reference 1.

11-3

ENCLOSURE 2 Westinghouse Non-Proprietary Class 3 WCAP-15885 July 2002 Revision 0 R. E. Ginna Heatup and Cooldown Limit Curves for Normal Operation OF U-

• @Westinghouse

9 3 RADIATION ANALYSIS AND NEUTRON DOSIMETRY

3.1 INTRODUCTION

This section describes a discrete ordinates S. transport analysis performed for the R. E. Ginna reactor to determine the neutron radiation environment within the reactor pressure vessel and surveillance capsules. In this evaluation, fast neutron exposure parameters in terms of fast neutron fluence (E > 1.0 MeV) and iron atom displacents (dpa) were established on a plant and fuel cycle specific basis for the first twenty nine reactor operating cycles. In addition, neutron dosimetry sensor sets from the first four surveillance capsules withdrawn from the R. E. Ginna reactor were re-analyzed using current dosimetry evaluation methodology.

The results of these dosimetry re-evaluations provided a validation of the plant specific neutron transport calculations. The validated calculations were then used to project future fluence accumulation through operating periods extending to 54 effctive full power years (efpy).

The use of fast neutron fluence (E > 1.0 MeV) to correlate measured material property changes to the neutron exposure of the material has traditionally been accepted for development of damage trend curves as well as for the implementation of trend curve data to assess vessel condition. In recent years, however, it has been suggested that an exposure model that accounts for differences in neutron enea-y spectra between surveillance capsule locations and positions within the vessel wall could lead to an improvcment in the uncertainties associated with damage trend curves as well as to a more accurate evaluation of damage gradients through the reactor vessel wall.

Because of this potential shift away ftom a threshold fluence toward an energy dependent damage function for data correlation, ASTM Standard Practice E853, "Analysis and Interpretation of Light-Water Reactor Surveillance Results," recommends reporting displacements per iron atom (dpa) along with fluence (E > 1.0 MeV) to provide a data base for future reference. The energy dependent dpa function to be used for this evaluation is specified in ASTM Standard Practice E693, "Characterizing Neutron Exposures in Iron and Low Alloy Steels in Terms of Displacements per Atom." The application of the dpa parameter to the assessment of embrittlement gradients through the thickmess of the reactor vessel wall has already been promulgated in Revision 2 to Regulatory Guide 1.99, "Radiation Embriement of Reactor Vessel Materials." Therefore, in keeping with the philosophy espoused in the current standards governing pressure vessel exposure evaluations, dpa data is also included in this section.

All of the calculations and dosimetry evaluations described in this report were based on the latest available nuclear cross-section data derived from ENDFIB-VI and made use of the latest available calculational tools. Furthermore, the neutron transport and dosimetry evaluation methodologies follow the guidance and meet the requirements of Regulatory Guide 1.190, "Calculational and Dosimetry Methods for.Deermining Pressure Vessel Neutron Fluence.W411) Additionally, the methods used to determine ttewessure vessel neutron exposure are consistent with the NRC approved methodology described in WCAP-14040-NP-A, "Methodology Used to Develop Cold Overpressure Mitigating System Setpoints and RCS Heatup and Cooldown Limit Curves," January 1996.J

10 3.2 NEUTRON TRANSPORT CALCULATIONS In performing the fist neutron exposure evaluations for the R. E. Ginna surveillance capsules and reactor vessel, plant specific forward transport calculations were carried out using the following tredimmensional flux synthesis technique:

0(r6,0z) = g5(rO)

• _r__)

where V(r,9,z) is the synthesized three-dimensional neutron flux distribution, +(rO) is the transport solution in re geometry, C(rz) is the two-dimensional solution for a cylindrical reactor model using the actual axial core power distribution, and +(r) is the one-dimensional solution for a cylindrical reactor model using the same source per unit height as that used in the rG two-dimensional calculation.

For the R. E. Ginna analysis, all of the transport calculations were carried out using the DORT two-dimensional discrete ordinates code Version 3. [131 and the BUGLE-96 cross-section librar 41 . The BUGLE-96 library provides a 67 group coupled neutron-gamma ray cross-section data set produced specifically for light water reactor application. In these analyses, anisotropic scattering was treated with a Ps legendre expansion and the angular discretization was modeled with an S16 order of angular quadrature.

A plan view of the rO model of the R E. Ginna reactor geometry at the core midplane is shown in Figure 1.

Since the reactor exhibits octant symmetry only a 0° to 450 sector is depicted. In addition to the core, reactor internals, pressure vessel and primary biological shield, the model also included explicit representations of the surveillance capsules, the pressure vessel cladding, and the insulation located external to the pressure vessel.

From a neutronic standpoint the inclusion of the surveillance capsules and associated support structure in the analytical model is significant. Since the presence of the capsules and structure has a marked impact on the magnitude of the neutron flux as well as on the relative neutron and gamma ray spectra at dosimetry locations within the capsules, a meaningful evaluation of the radiation environmt internal to the capsules can be made only when these perturbation effects are properly accounted for in the analysis.

In developing the re analytical model of the reactor geometry shown in Figure 1, nominal design dimensions were employed for the various structural components. Likewise, water temperatures and, hence, coolant density in the reactor core and downcomer regions of the reactor were taken to be represetave of full power operating conditions. The reactor core itself was treated as a homogeneous nmxture of fuel, cladding, water, and miscellaneous core structures such as fuel assembly grids, guide tubes, etc. The rO geometric mesh description of the reactor-model shown in Figure 3.2-1 Vvsisted of 170 radial by 67 azimuthal intervals. Mesh sizes were chosen to assure that proper convergence of the inner iterations was achieved on a pointwise basis. The pointwise inner iteration flux convergence criterion utilized in the re calculations was set at a value of 0.001.

A section view of the rz model of the L E. Ginna reactor is shown in Figure 2. The model extended radially fromnthe centerline of the reactor core out to a location interior to the primary biological shield and over an axial span from an elevation 1 foot below the active fuel to approximately 1 foot above the active fuel. As in the case of the r,9 model, nominal design dimensions and full power coolant densities were

11 employed in the calculations. hi this case, the homogenous core region was treated as an equivalent cylinder with a volume equal to that of the active core zone. The stainless steel former plates located between the core baffle and core barrel regions were also explicitly included in the model. The rz geometric mesh description of the reactor model shown in Figure 2 consisted of 153 radial by 90 axial intervals.

Mesh sizes were chosen to assure that proper convergence of the inner iterations was achieved on a pointwise basis. The pointwise inner iteration flux convergence criterion utilized in the rz calculations was also set at a value of 0.001.

The one-dimensional radial model used in the synthesis procedure consisted of the same 153 radial mesh intervals included in the rz model. Thus, radial synthesis factors could easily be determined on a meshwise basis throughout the entire geometry.

The core power distributions used in the plant specific transport analysis for the R. E. Ginna reactor were taken from the appropriate fuel cycle design reports for Cycles 1 through 29. The data extracted from the design reports represented cycle average relative assembly powers, burnups, and axial distributions.

Therefore, the calculated results provided data in terms of fuel cycle averaged neutron flux which, when multiplied by the appropriate fuel cycle length, in urn, yielded the incremental fast neutron exposure for each fuel cycle. In constructing, the core source distributions, the energy distribution of the source was based on an appropriate fission split for uranium and plutonium isotopes; and from that fission split, composite values of energy release per fission, neutron yield per fission, and fission spectrum were determined. Fluence prqjections beyond the end of Cycle 29 were based on the assumption that the core power distribution averaged over Cycles 26 through 29 would be representative of future plant operation.

Cycles 26 through 29 were designed as 18 month fuel cycles using the low leakage fuel management concept.

The maximum calculated fast neutron fluence (E > 1.0 MeV) and dpa exposure values for the R. E. Ginna pressure vessel are provided in Table 6. As presented, these data represent the maximum exposure of the pressure vessel clad/base metal interface at azimuthal angles of 0, 15, 30, and 45 degrees relative to the core cardinal axes. The data tabulation includes the plant specific calculated fluence at the end of cycle twenty nine (the last cycle completed at the R E. Ginna plant) and projections for future operation to 28, 32, 36, 40, 44, 48, 52, and 54 EFPY Similar data applicable to the intermediate shell to nozzle shell circm nal weld as wel as to the nozzle shell course located above the top of the active fuel stack are given in Table 7.

The results of the updated fluence calculations for the four surveillance capsules withdrawn to date from the R. E. Ginna reactor are provided in Table 8. These calculated values of neutron fluence should be used to specify the neutron exposure of the irradiated test specimens for use in materials damage correlations.

Updated lead factors for the R E. Ginna surveillance capsules are provided in Table 9. Thevcapsule lead factor is defined as the ratio of the calculated fluene at the jeomctric center of the surveilHce capsule to the co ng maximum calculated fluence at the pressure vessel clad/base metal interface.

In Table 9, the lead factors for capsules that have been withdrawn from the reactor (V, R, T, and S) were based on the calculated fluence values for the irradiation period corresponding to the time of withdrawal for the individual capsules. For the capsules remainin in the reactor (P and N), the lead factors correspond to the calculated fluence values at the projected end of cycle twenty nine, the last fuel cycle completed at the time of analysis. The lead factors provided in Table 9 should be used as the basis for the development of fiuture capsule withdrawal schedules for the R E. Ginna reactor.

12 R. E. Ginna re Reactor Gomnetry at the Core. M#idane 0

0 N

i'saa'0 1-I 49 03 0

6 a1

13 Figure 2 R. E. Gima rz Geometry 200*

1520 100

- 0a

-1800

-200 * * ~ *

• 0 40 80 120 It0 Soo 240 IAzb (an) a a -,

r 14 Table 6 Summary of Calculated Maximumr Pressure Vessel Exposure CladMase Metal Interfaice Neutrm Fluence [E > 1.0 MeV]

Cumulative Neutron Fluence In/cn?2 Operating Time EFjP. 0.0 Degrees 15.0 Dezrees 30.0 Degrees 45.0 Degrees 24.8 (EOC 29) 2.68e+19 1.69e+19 1.22e+19 1.09e+19 28 2.94e+19 1.85e+19 1.34e+19 1.20e+19 32 3.26e+19 2.05e+19 1.48e+19 1.33e+19 36 3.57e+19 2.25e+19 1.63e+19 1.46e+19 40 3.89e+19 2.45e+19 1.77e+19 1.60e+19 44 4.21e+19 2.65e+19 1.92e+19 1.73e+19 48 4.53e+19 2.85e+19 2.07e+19 1.86e+19 52 4.85e+19 3.05e+19 2.21e+19 2.00e+19 54 5.0le+19 3.15e+19 2.28e+19 2.06e+19 Iron Atom Displacements Cumulative Iron Atom Dis lacements Idpal Operating Time JEFPY 0.0 Degrees 15.0 Degrees 30.0 Decrees 45.0 Degrees 24.8 (EOC 29) 4.37e-02 2.85e-02 2.Ole-02 1.77e-02 28 4.79e-02 3.12e-02 2.20e-02 1.94e-02 32 5.3le-02 3.46e-02 2.44e-02 2.16e-02 36 5.82c-02 3.79e-02 2.68e-02 2.37c-02 40 6.34e-02 4.12e-02 2.91c-02 2.59e-02 44 6.86e-02 4.46e-02 3.15e-02 2.80e-02 48 7.38e-02 4.79e-02 3.39e-02 3.02e-02 52 7.89e-02 5.12e-02 3.63e-02 3.23e-02 54 8.1 5e-02 5.29e-02 3.75e-02 3134e-02

15 Table 7 Su of Calaclated Maximum Exposure of the Intermediate to Nozzle Shell Circumferential Weld and the Nozzle Shell Course Clad/Base Metal Interface Neutron Fluence [E > 1.0 MeV Cumulative Neutron Fluence inicmLl Operating Time XEFLP. 0.0 Degrees 15.0 Degrees 30.0 Degrees 45.0 Degrees 24.8 (EOC 29) 1.05e+18 6.64e+17 4.78e+17 4.26e+17 28 1.16e+18 7.28e+17 5.25e+17 4.68e+17 32 1.28e+18 8.07c+17 5.83e+17 5.21e+17 36 1.41e+18 8.86e+17 6.41e+17 5.74e+17 40 l.54e+l8 9.65e+17 6.98e+17 6.27e+17 44 1.66c+18 1.04e+18 7.56e+17 6.80e+17 48 1.79e+18 1.12e+18 8.14e+17 7.32e+17 52 1.92e+18 1.20e+18 8.72e+17 7.85e+17 54 1.98e+18 1.24e+18 9.00e+17 8.12c+17 Iron Atom Displacements Cumulative Iron Atom Disp lacements Idpal Operating Time EFY 0.0 Degrees 15.0 Degrees 30.0 Degrees 45.0 Degrees 24.8 (EOC 29) 1.83e-03 1.19e-03 8.38e-04 7.35e-04 28 2.01e-03 1.31e-03 9.20e-04 8.09c-04 32 2.23e-03 1.45e-03 1.02e-03 9.0le-04 36 2.45e-03 l.59c-03 1.12e-03 9.92e-04 40 2.67e-03 1.73e-03 1.22e-03 1.08e-03 44 2.89e-03 1.87e-03 1.33e-03 1.17e-03 48 3.1le-03 2.02e-03 1.43e-03 1.27e-03 52 3.33e-03 2.16e-03 1.53e-03 ).36c-03 54 3.44e-03 2.23e-03 l.58e-03 , "1.40e-03

16 Table 8 Calculated Surveillance Capsule Exposure Irradiation Time Fluence (E > 1.0 MeV) Iron Displacements Capsule IEFPYI In/cm2 l Idpal V 1.4 5.87e+18 1.07e"02 R 2.6 1.02e+19 1.85e-02 T 6.9 1.69e+19 2.94e-02 S 17.0 3.64e+19 6.38e-02 Table 9 Calculated Surveillance Capsule Lead Factors Capsule ID And Location Status Lead Factor V (130) Withdrawn EOC 1 2.96 R (130) Withdrawn EOC 3 2.97 T (230 ) Withdrawn EOC 9 1.82 S (33°) Withdrawn EOC 22 1.79 P (230) In Reactor 1.91 N (33 0 ) In Reactor 1.81 Note: Lead fictors for capsulas reanin in the reactor are based on cycle specific ccposure calculations through fael cycle twenty nine.

4 -,

17 3.3 NEUTRON DOSIMETRY EVALUATIONS 3.3.1 Sensor Reaction Rate Determinations In this section, the results of the evaluations of the fur neutron sensor sets withdrawn as a part of the R E.

Ginna Reactor Vessel Materials Surveillance Program are presented. The capsule designation, location within the reactor, and time of withdrawal of each of these dosimetry sets were as follows:

Azinuthal Withdrawal hradiation Capsule ID Location Time Time MeMbsi V 130 End of Cycle 1 4.46e+07 R 13° End of Cycle 3 8.05e+07 T 230 End of Cycle 9 2.17e+08 S 330 End of Cycle22 5.36e+08 The type and radial locations of the neutron sensors within the capsules arc sunmarized as follows:

Radius Sensor Type [Cm]

Copper 158.11 158.11 Iron - Core Side Charpy 159.11 Iron - Vessel Side Charpy 158.91 Nickel 158.35 Uranium 238 158.35 Neptunium 237 159.11 Bare Cobalt-Aluminum 159.11 Cd Cov. Cobalt-Aluminum The copper, nickel, and cobalt-aluminum monitors, in wire form, were placed in holes drilled in spacers at several axial levels within the capsules. The cadmium shielded uranium and neptumium fission monitors were accomnodated within the dosimeter block located nea the center of the capsule. The iron sensors were obtained by cutting small samples from individual charpy specimens taken from several locations within the surveillance capsules.

18 The use of passive monitors such as those listed above does not yield a direct measure of the energy dependent neutron flux at the point of interest. Rather, the activation or fission process is a measure of the integrated effect that the time and energy dependent neutron flux has on the target material over the course of the irradiation period. An accurate assessment of the average neutron flux level incident on the various monitors may be derived from the activation measurements only if the irradiation paramers are well kown. In particular, the following variables are of interest:

• The measured specific activity of each monitor,

• The physical charactestics of each monitor,

• The operating history of the reactor,

• The energy response of each monitor, and

• The neutron energy spectrum at the monitor location.

The radiometic counting of each of the R. E. GPna dosimetry data sets was accomplished by Westinghouse using established ASTM procedures. Following sample preparation and wghing, the activity of each monitor was dermined by means of a high resolution gamma spectrmneter. For the copper, iron, nickel, and cobalt-alummum sensors, these analyses were performed by direct counting of each of the individual samples. In the case of the uranium and neptumium fission sensors, the analyses were carried out by direct counting preceded by disolution and chemical separation of cesium frm the sensor material.

The irradiation history of the reactor over the irradiation period experienced by Capsules V, P., T, and S was obtained on a monthly basis from reactor startup to the end of the dosimetry evaluation period. For the sensor sets utilized in the surveillance capsules, the half-lives of the product isotopes are long enough that a monithly histogram describing reactor operation has proven to be an adequate represetation for use in radioactive decay corrections for the reactions of interest in the exposure evaluations.

Having the measured specific activities, the operating history of the reactor, and the physical ca isics of the sensors, reaction rates referenced to full power operation were determined from the following equation:

A NFY -CJ (-eJ )et e where:

A = measured specific activity (dps/g)

R = reaction rate averaged over the irradiation period and referenced to operation at a core power level of Pad (rps/nucleus).

No number of target clement atoms per gran of sensor.

19 F = weight fiaction of the target isotope in the sensor material.

Y = number of product atoms produced per reaction.

P = average core power level during irradiation period j (MW).

P, = maximum or reference core power level of the reactor (MW).

= calculated ratio of +(E > 1.0 MeV) during irradiation period j to the time weged average +(E > 1.0 MeV) over the entire irradiation period.

A

= decay constant of the product isotope (s').

length of irradiation period j (s).

td = decay time following irradiation period j (s).

and the summation is carried out over the total number of monthly intervals comprising the irradiation period.

In the above equation, the ratio P/Pf accounts for month by month variation of power level within a given fuel cycle. The ratio q is calculated for each fuel cycle using the methodology descnbed in Section 3.2 of this report and accounts for the change in sensor reaction rates caused by variations in flux level due to changes in core power spatial distributions from fuel cycle to fuel cycle. For a single cycle irrdiation Cq =

1.0. However, for multiple cycle irradiations, particularly those employing low leakage fuel management, the additional Cj correction must be utilized. THis additional correction can be quite significant for sensor sets that have been irradiated for many fuel cycles in a reactor that has transitioned from non-low leakage to low leakage fiel management.

Prior to using the measured reaction rates in the least squares adjustment procedure discussed later in this section, additional corrections were made to Urn measurements to account for the presence of U1 5 impurities in the sensors as well as to adjust for the build-in of plutonium isotopes over the course of the irradiation. These corrections were location and fluence dependent and were derived from the plant specific discrete ordinates analysis described in Section 3.2. Corrections were also made to the U2M and Np2" sensor reaction rates to account for gamma ray induced fission reactions that occurred over the course of the irradiation. These photo-fission corrections were, likewise, location dependent and were based on the transport calculations described in Section 3.2.

Results of the sensor reaction rate determinatons for Capsules V, R, T, and S are given in Tables 10 hough 14. In Tables 10 through 13, the measured specific activities, gradient corrected specific activities, and decay corrected reaction rates are listed for Capsules V, R, T, and S, respectively. A summary of the reaction rates for each capsule is provided in Table 14. The data listed in Table 14 are indexed to the geometric center of the respective capsules and included all corrections for UIP impurities, Pu build-in, and photo-fission effects.

Ile

20 Table 10 Measured Sensor Specific Activities and Reaction Rates Capsule V Radially Radially Adjusted Adjusted Average Measured Saturated Saturated Reaction Reaction Radius Activity Activity Activity Rate Rate Sample ID Foil ID [cm.] [dpslgl [dpslg] [dpslg] [rpslatoml [rpslatomJ CU Top 158.11 7.38E+04 4.63E+05 4.43E+05 6.77E-17 CU Top-Mid 158.11 6.77E+04 4.25E+05 4.07E+05 6.21E-17 CU Bot-Mid 158.11 7.48E+04 4.70E+05 4.49E+05 6.B6E-17 CU Bottom 158.11 8.13E+04 5.10E+05 4.89E+05 7.45E-17 6.82E-17 FE W-1 158.11 2.47EI06 5.00E+06 4.82E+06 7.64E-15 FE R-1 158.11 2.57E+06 5.20E+06 5.02E+06 7.95E-15 FE S-6 158.11 2.18E+06 4.41E+06 4.25E+06 6.74E-15 FE P-7 158.11 2.57E+06 5.20E+06 5.02E+06 7.95E-15 FE W-2 159.11 2.04E+06 4.13E+06 4.78E+06 7.58E-15 FE R-3 159.11 1.95E+06 3.95E+06 4.57E+06 7.25E-15 FE S-8 159.11 2.02E+06 4.09E+06 4.74E+06 7.51E-15 FE P-9 159.11 2.10E+06 4.25E+06 4.92E+06 7.BOE-15 7.55E-15 Ni Middle 158.11 2.38E+07 6.51E+07 6.21E+07 8.90E-15 8.90E-15 U Middle 158.35 2.30E+05 7.26E+06 7.26E+06 4.77E-14 3.91E-14 NP Middle 158.35 1.23E+06 3.88E+07 3.88E+07 2.48E-13 244E-13 Notes:

• The average U-238(nf) reaction rate of 2.91E-14 includes the correction of a fctor of 0.861 to account for plutonium build-in and an additional fator of 0.950 to account for photo-fission effects in the sensor.

• The average Np-237(nf) reaction rate of 2.44E-13 includes the correction of a factor of 0.983 to account for the photo-fission effects in the sensor.

{ 4,i

21 Table 11 Measured Sensor Specific Activities and Reaction Rates Capsule R Radially Radially Adjusted Adjusted Average Measured Saturated Saturated Reaction Reaction Radius Activity Activity Activity Rate Rate Sample ID Foil ID [cm.J [dpslg] [dpslg] [dpslg] [rpslatomJ [rpslatom]

74-2204 Top 158.11 1.08E405 4.42E+05 4.23E+05 6.45E-17 74-2207 Top-Mid 158.11 9.68E+04 3.96E+05 3.79E+05 5.78E-17 74-2213 Bot-Mid 158.11 1.15E.05 4.70E+05 4.50E+05 6.87E-17 74-2216 Bottom 158.11 I1.15E+05 4.70E+05 4.50E+05 6.87E-17 6.49E-17 74-2202 W-13 158.11 2.08E+06 5.19E.06 5.OOE+06 7.93E-15 74-2200 R-14 158.11 1.98E+06 4.94E+06 4.76E+06 7.55E-15 74-2198 P-18 158.11 2.06E+06 5.14E.06 4.95E+06 7.85E-15 74-2203 W-14 159.11 I1.63E.06 4.07E+06 4.71E+06 7.46E-15 74-2201 R-15 159.11 I1.70E.06 4.24E4+06 4.91E+06 7.78E-15 74-2199 P-19 159.11 1.85E.06 4.61 E+06 5.34E+06 8.47E-15 7.84E-15 74-2210 Middle 158.11 5.83E+06 7.36E+07 7.03E+07 1.01E-14 1.01E-14 74-2220 Middle 158.35 4.32E+05 7.79E+06 7.79E+06 5.12E-14 4.11E-14 74-2219 Middle 158.35 4.25E+06 7.66E.07 7.66E+07 4.89E-13 4.81E-13 74-2205 Top 159.11 3.09E.07 I1.26E+08 1.22E+08 7.97E-12 74-2208 Top-Mid 159.11 3.14E+07 1.28E.08 1.24E+08 8.10E-12 74-2211 Middle 159.11 2.96E+07 1.21 E.08 1.17E+08 7.64E-12 74-2214 Bot-Mid 159.11 2.94E+07 1.20E4+08 1.16E+08 7.59E-12 74-2217 Bottom 159.11 2.94E.07 1.20E+08 1.16E+08 7.59E-12 7.78E-12 74-2206 Top 159.11 1.19E.07 4.87E407 5.69E+07 3.71E-12 74-2209 Top-Mid 159.11 1.18E407 4.83E407 5.64E+07 3.68E-12 74-2212 Middle 159.11 I1.07E.07 4.38E.07 5.1 1E+07 3.34E-12 74-2215 Bot-Mid 159.11 I1.24E+07 5.07E+07 5.92E+07 3.87E-12 74-2218 Bottom 159.11 I1.24E+07 5.07E+07 5.92E+07 3.87E-12 3.69E-12 Notes:

• The average U-238(nf) reaction rate of 4. 11E-14 includes the correction of a factor of 0.845 to account for plutonium build-in and an additional factor of 0.950 to account for photo-fission effects in the sensor.

• The average Np-237(nf) reaction rate of 4.81E-13 includes the correction of a factor of 0.983 to account for the photo-fission effects in the sensor.

4

22 Table 12 Measured Sensor Specific Activities and Reaction Rates Capsule T Radially Radially Adjusted Adjusted Average Measured Saturated Saturated Reaction Reaction Radius Activity Activity Activity Rate Rate Sample ID Foil ID [cm.] [dpslg] [dpslg] [dps/g] [rpslatom] [rpslatom]

81-1392 Top 158.11 1.60E+05 3.51E+05 3.35E+05 5.11E-17 81-1395 Top-Mid 158.11 1.40E+05 3.07E+05 2.93E+05 4.47E-17 81-1402 Bot-Mid 158.11 1.66E+05 3.64E+05 3.48E+05 5.30E-17 81-1415 Bottom 158.11 1.74E+05 3.82E+05 3.64E+05 5.56E-17 5.11E-17 8143390 S-22 158.11 1.14E+06 3.36E+06 3.19E+06 5.06E-15 81-3392 P-28 158.11 1.27E+06 3.74E+06 3.56E+06 5.64E-15 8143394 W-21 158.11 1.30E+06 3.83E+06 3.64E+06 5.77E-15 8143391 S-23 159.11 1.01E+06 2.97E+06 3.43E+06 5.44E-15 814393 P-29 159.11 1.03E+06 3.03E+06 3.50E+06 5.55E-15 813395 W-22 159.11 1.10E+06 3.24E+06 3.74E+06 5.92E-15 5.56E-15 81-1399 Middle 158.11 8.62E+05 5.25E+07 5.01E+07 7.17E-15 7.17E-15 81-1388 Middle 158.35 7.41E+05 5.34E+06 5.34E+06 3.51E-14 2.74E-14 81-1389 Middle 158.35 6.09E+06 4.39E+07 4.39E+07 2.80E-13 2.75E-13 81-1390 Top 159.11 3.17E+07 6.96E+07 6.60E+07 4.31E-12 81-1393 Top-Mid 159.11 3.06E+07 6.72E+07 6.37E+07 4.16E-12 81-1396 Middle 159.11 3.03E+07 6.65E+07 6.31E+07 4.12E-12 81-1400 Bot-Mid 159.11 3.27E+07 7.18E+07 6.81E+07 4.44E-12 81-1403 Bottom 159.11 3.07E+07 6.74E+07 6.39E+07 4.17E-12 4.24E-12 81-1391 Top 159.11 1.21E+07 2.66E+07 3.O6E+C07 2.OOE-12 81-1394 Top-Mid 159.11 1.13E+07 2.48E+07 2.86E+07 1.87E-12 81-1397 Middle 159.11 1.16E+07 2.55E+07 2.94E+07 1.92E-12 81-1401 Bot-Mid 159.11 1.26E+07 2.77E+07 3.19E+07 2.08E-12 81-1404 Bottom 159.11 1.20E+07 2.63E+07 3.04E+07 1.98E-12 1.97E-12 Notes:

• Ihe average U-238(nf) reaction rate of 2.74E-14 includes the correction of a factor of 0.820 to account for plutonium build-in and an additional fiator of 0.955 to account for photo-fission effects in the sensor.

• The average Np-237(nf) reaction rate of 2.75E-13 includes the correction of a factor of 0.984 to account for the photo-fission effects in the sensor.

4O-

23 Table 13 Measured Sensor Specfic Activities and Reaction Rates Capsule S Radially Radially Adjusted Adjusted Average Measured Saturated Saturated Reaction Reaction Radius Activity Activity Activity Rate Rate Sample ID Fall ID [cm.l [dpsfg] [dpslg] [dpsfg] [rpsiatom] [rlatomJ 93-3163 Top 158.11 2.06E+05 3.06E+05 2.92E+05 4.45E-17 93-3166 Top-Mid 158.11 1.82E+05 2.70E+05 2.58E+05 3.93E-17 93-3172 Bot-Mid 158.11 1.98E+05 2.94E+05 2.81E+05 4.28E-17 93-3175 Bottom 158.11 2.18E+05 3.24E+05 3.09E+05 4.71E-17 4.34E-17 93-4326 P-31 158.11 1.62E+06 2.93E+06 2.79E.06 4.42E-15 4.42E-15 93-3169 Middle 158.11 8.51E+06 4.27E+07 4.06E+07 5.81E-15 5.81E-15 93-3159 Middle 158.35 1.40E+06 4.63E+06 4.63E+06 3.04E-14 2.19E-14 93-3160 Middle 158.35 1.11 E+07 3.67E+07 3.67E+07 2.34E-13 2.30E-13 93-3161 Top 159.11 3.55E+07 5.27E+07 S.05E+07 3.29E-12 93-3164 Top-Mid 159.11 3.71E+07 5.51E+07 5.28E+07 3.44E-12 93-3167 Middle 159.11 3.39E+07 5.03E+07 4.82E+07 3.15E-12 93-3170 Bat-Mid 159.11 3.60E+07 5.35E+07 5.12E+07 3.34E-12 93-3173 Bottom 159.11 3.45E+07 5.12E+07 4.91E+07 3.20E-12 3.29E-12 93-3162 Top 159.11 1.43E+07 2.12E+07 2.47E+07 1.61E-12 93-3165 Top-Mid 159.11 1.37E+07 2.03E+07 2.37E+07 1.54E-12 93-3168 Middle 159.11 1.31E+07 1.95E+07 2.26E+07 1.48E-12 93-3171 Bot-Mid 159.11 1.45E+07 2.15E+07 2.50E+07 1.63E-12 93-3174 Bottom 159.11 1.35E+07 2.OOE+07 2.33E+07 1.52E-12 1.56E-12 Notes:

• The average U-238(nsf) reaction rate of 2. 19E-14 includes the correction of a fictor of 0.755 to account for plutonium build-in and an additional factor of 0.953 to account for photo-fission effects in the sensor.

• The average Np-237(nf) reaction rate of 2.30E-13 includes the correction of a fictor of 0.983 to account for the photo-fission ceffects in the sensor.

24 Table 14 Sunmnary of Sensor Reaction Rates fom Capsules V,R, T, and S Measured Reaction Rate Irpsenucl Sensor Reaction Capsule V Capsule R sule T Capsule S Cu-63(n4)Co-60 6.82e-17 6.49e-17 5.11e-17 4.34e-17 Fe-54(np)Mn-54 7.55e-15 7.84e-15 5.56e-15 4.42e-15 Ni-58(np)Co-58 8.90e-15 1.Ole-14 7.17e-15 5.81e-15 U-238(nf)Cs-137 Cd Covered 3.91e-14 4.11e-14 2.74e-14 2.19e-14 Np-237(nf)Cs-137 Cd Covered Rejected 4.8le-13 2.75e-13 2.30e-13 Co-59(ny) Co-60 None 7.78e-12 4.24e-12 3.29e-12 Co-59(ny) Co-60 Cd Covered None 3.69e-12 1.97e-12 1.56e-12

. v

25 3.4 LEAST SQUARES EVALUATION OF SENSOR SETS Least squares adjustment methods provide the capability of combining the measurement data with the neutron transport calculation resulting in a Best Estimate neutron energy spectrum with associated uncertainties. Best Estimates for key exposure parameters such as +(E > 1.0 MeV) or dpa/s along with their uncertainties are then easily obtained from the adjusted spectrum. in general, the least squares methods, as applied to surveillance capsule dosimetry evaluations, act to reconcile the measured sensor reaction rate data, dosimetry reaction cross-sectibns, and the calculated neutron energy spectrum within their respective uncertainties. For example, R in =X(a+/-c)(09,+/-6Q) relates a set of measured reaction rates, R; to a single neutron spectrum, X through the multigroup dosimeter reaction cross-section, cr,, each with an uncertainty 6. The primary objective of the least squares evaluation is to produce unbiased estimates of the neutron exposure parameters at the location of the measurement.

For the least squares evaluation of the R. E. Ginna surveillance capsule dosimetry, The FERRET code&P was employed to combine the results of the plant specific neutron transport calculations and sensor set reaction rate to determine best estimate values of exposure parameters (*(E > 1.0 MeV) and dpa) along with associated uncertainties for the three in-vessel capsules withdrawn to date.

The application of the least squares methodology requires the following input:

I - The calculated neutron energy spectrum and associated uncertainties at the measurement location.

2 - The measured reaction rates and associated uncertiunty for each sensor contained in the multiple foil set.

3 - The energy dependent dosimetry reaction cross-sections and associated uncertainties for each sensor contained in the multiple foil sensor set.

For the R. E. Ginna application, the calculated neutron spectrum was obtained from the resylts of plant specific neutron transport calculations described in Section 3.2 of this report. The senqoraction rates were derived from the measured specific activities using the procedures described in Section 3.3. The dosimetry reaction cross-sections and uncertainties were obtained from the SNLRML dosimetry cross-section library The SNLRML ibrary is an evaluated dosimetry reaction cross-section compilation recommended for use in LWR evaluations by ASTM Standard E1018, "Application of ASTM Evaluated Cross-Section Data File, Matrix E 706 (IB)".

26 The uncertainties associated with the measured reaction rates, dosimetry cross-sections, and calculated neutron spectrum were input to the least squares procedure in the form of variances and covariances. The assignment of the input uncertainties followed the guidance provided in ASTM Standard E 944, "Application of Neutron Spectrum Adjustment Methods in Reactor Surveillance.

The following provides a summary of the uncertainties associated with the least squares evaluation of the R. E. Ginna surveillance capsule sensor sets:

Reaction Rate Uncertainties The overall uncertainty associated with the measured reaction rates includes components due to the basic measurement process, the irradiation history corrections, and the corrections for cmpeting reactions. A high level of accuracy in the reaction rate determinations is assured by ulizing laboratory procedures that conform to the ASTM National Consensus Standards for reaction rate de ions for each sensor type.

After combining all of these uncertainty commpoents, the sensor reaction rates derived from the counting and data evaluation procedures were assigned the following net uncertainties for input to the least squares evaluation:

Reaction Uncertainty Cu6 (nIA)CoP 5%

Fes'(n,p)Mn 5%

Ni (n,p)Co' 5%

UIP(ft)Cs S37 10%

Np'(n,f)Cs137 10%

Co59(n,y)Co 60 5%

Dosinetnr Cross-Section Uncertainties The reaction rate cross-sections used in the least squares evaluations were taken from the SNLRML library. This data library provides reaction cross-sections and associated uncertainties, including covariances, for 66 dosimetry sensors in common use. Both cross-sections and uncertainties are provided in a fine multigroup structure for use in least squares adjustment applications. These cross-sections were compiled from the most recent cross-section evaluations and they have been tested with respect to ther accuracy and consistency for least squares evaluations. Further, the library has been empirically tested for use in fission spectra determination as well as inthe fluenc and energych afon of 14 MeV neutron sources. Detailed discussions of the contents of the SNLRML library along with the evaluation process for each of the sensors is provided in Reference 16. .

For sensors included in the RI E. Ginna surveillance capsules, the following uncertainties in the fission spectrum averaged cross-sections are provided in the SNLRM doumentatIn package.

27 These tabulated ranges provide an indication of the dosimetry cross-section uncertainties associated with the sensor sets used in LWR irradiations.

Calculated Neutron Spect The neutron spectrum input to the least squares adjustment procedure was obtained directly from the results of plant specific transport calculations for each surveillance capsule location. The spec at each location was input in an absolute sense (rather thaan as simply a relative spectral shape). Therefore, within the constraints of the assigned uncertainties, the calculated data were treated equally with the measurements.

While the uncertainties associated with the reaction rates were obtained from the measurement procedures and counting benchmarks and the dosimetry cross-section uncertainties were supplied directly with the SNLRML ibary, the uncertainty matrix for the calculated spectrum was constructed from the following relationship:

r= +R * *R. s where R. specifies an overall fractional normalization uncertainty and the fractional uncertainties R1 and Rg specify additional random groupwise uncertainties that are correlated with a correlation matrix given by.

P. = = [ 19-] * . + 8 ef where (g -g')!

H=(2r 2 The first term in the correlation matrix equation specifies purely random uncertainties, while the second term describes the short range correlations over a group range y(9 specifies the strength of the latter term).

The value of &is 1.0 when g = g' and 0.0 otherwise.

28 The set of parameters defining the input covariance matrix for the R E. Ginna calculated spectra was as follows:

Flux Normalization Uncertainty (R) 15%

Flux Group Uncertainties (Re, Rs.)

(E > 0.0055 MeV) 15%

(0.68 eV < E < 0.0055 MeV) 29%

(E < 0.68 52%

5V)

Short Range Correlation (0)

(E > 0.OO55 MeV) 0.9 (0.68 eV < E < 0.0055 MeV) 0.5 (E < 0.68 eV) 0.5 Flux Group Corrlation Range (y)

(E >0.OO55 MeV) 6 (0.68 eV < E < 0.0055 MeV) 3 (E < 0.68 eV) 2 Results of the least squares evaluation of the four sensor sets withdrawn from the R. E. Ginna reactor are provided in Tables 15 and 16. In Table 15, measured, calculated, and best estimate sensor reaction rates are given for Capsules V, R, T, and S. The improvement in the fit of the adjusted spectra to the measuem is evident for all four capsule data sets. Prior to the application of the adjustment procedure M/C ratios for individual foil reactions ranged from 0.75 to 1.32, while after the adjustment M/BE ratios ranged from 0.91 to 1.12. Thus, demonsng a significant improvement in the data fits.

In Table 16, the calculated and best estimate exposure rates and integrated exposures of Capsules V, R, T, and S are given. Data are provided in terms of both fluence (E > 1.0 MeV) and iron atom displacements.

• ace0/.

29 Table 15 Comparison of Measured, Calculated, and Best Estimate Reaction Rates at the Surveillance Capsule Center Surveillance Capsule V Reaction Rate IrpsaomL Best Reaction Measured Calculated Estimate MIC MtBE Cua(na)Coe 6.82e-17 6.69c-17 6.54e-17 1.02 1.04 Fe"(n,p)Mn" 7.55e-15 8.14e-l5 7.34e-15 0.93 1.03 Ni5(n,p)Co's 8.90c-15 1.14e-14 9.74e-15 0.78 0.91 U2s(n,f)Cs'7 Cd 3.91e-14 4.37e-14 3.77e-14 0.89 1.04 Np25(n,f)CsL" Cd CoP(n,y)Coe CoP9(nr)Co0 Cd Surveillance Capsule R Reaction Rate Irps a Lom Best Reaction Measured Calculated Estimate M/C MIBE CuO3 (nz)Co'0 6.49e-17 6.41e-17 6.39e-17 1.01 1.02 Fe"(n,p)Mn" 7.84e-15 7.79e-15 7.74e-15 1.01 1.01 Ni(n,p)Coe 1.Ole-14 1.09e-14 1.06e-14 0.93 0.95 UIr(n,f)Csu3? Cd 4.1le-14 4.18e-14 4.22e-14 0.98 0.97 NpD7 (n,f)CsU37 Cd 4.8le-13 3.64e-13 4.28e-13 1.32 1.12 Co5 9(n,y)Co60 7.78e-12 9.22c-12 7.84e-12 0.84 0.99 Co"(n,y)Coe0 Cd 3.69e-12 3.66e-12 3.68e-12 1.01 1.00 Surveillance Capsule T t:

Reaction Rate IrpsIaol Best Reaction Measured Calculated Estimate MIC MIBE Cuo(n,aeCo' 0 5.lle-17 S.09c-17 5.05e-17 1.00 1.01 Fe"(n,p)Mn 5.56e-15 5.5le-15 t5le-15 1.01 ve Nig8 (n,p)Cogs 7.17e-15 7.58e-15 7.46e-15 0.95 0.96 Ulfn(n,fCsU37 Cd 2.74e-14 2.70e-14 2.76e-14 1.01 0.99

,. Np (n,)Cs Cd 2.75e-13 2.14e-13 2.49e-13 1.29 1.10 9

Co0 (n,y)Co 0 4.24e-12 5.Ole-12 4.27e-12 0.85 0.99 Co' 9 (n,)Coe0 Cd 1.97e-12 1.89e-12 1.96e-12 1.04 1.01

30 Table 15 (contined)

Comparison of Measured, Calculaed, and Best Estimate Reaction Rates at the Surillance Capsule Center Surveillance Capsule S Reacton Rate IrPsatomL Best Reaction Measured Calculated Estimate MIC M/BE Cu0 (n,a)Co' 0 4.34e-17 4.25e-17 421e-17 1.02 1.03 Fe"(n,p)Mn" 4.42e-15 4.66c-15 4.45e-15 0.95 0.99 Nigs(n,p)Coe 5.8l-15 6.43e-15 6.05e-15 0.90 0.96 U=(n,f)Csl' Cd 2.19e-14 2.33e-14 2.24e-14 0.94 0.98 Npw7 (nf)Csm Cd 2.30e-13 1.S8e-13 2.06e-13 1.22 1.12 Co59(nT)Co' 0 3.29e-12 4.36e-12 3.32e-12 0.75 0.99 Co"(n,y)Coe0 Cd 1.56e-12 1.70e-12 1.56e-12 0.92 1.00

31 Table 16 Comparison of Calculated and Best Estimate Exposure Parameters at the Surveillane Capsule Center Time Averaged Exposure Rates O(E> 1.0 M ) [nlcm-si Best Calculated Estimate Uncertainty BE/C Capsule V 1.32e+11 1.13e+ll 7% 0.86 Capsule R 1.26e+11 1.30e+11 6% 1.03 Capsule T 7.79e+10 8.1le+10 6% 1.04 Capsule S 6.78e+10 6.62e+10 6% 0.98 Iron Atom Displ ae etsdt Best Calculated Estimate Uncertainty BE/C Capsule V 2.40e-10 2.05e-10 90/ 0.86 Capsule R 2.30e-10 2.37e-10 7% 1.03 Capsule T 1.36e-10 1.40e-10 7% 1.03 Capsule S 1.19e-10 1.16e-10 7% 0.97 Integrated Capsule Exposure 4 (E > 1.0 eV) In/cm2 Best Calculated Estimate Uncertainty BE/C Capsule V 5.87c+18 5.03e+18 7% 0.86 Capsule R 1.02e+19 1.05e+19 6% 1.03 Capsule T 1.69e+19 1.76e+19 6% 1.04 Capsule S 3.64e+19 3.55e+19 6% 0.98 Iron Atom Disp lacements Idpal Best Calculated Estimate Uncertainty BE/C .

Capsule V 1.07e-02 9.lSe-03 90.%

Capsule R 1.85e-02 1.91e-02 7% 1.03 Capsule T 2.94e-02 3.04e-02 7%/ 1.03 Capsule S 6.38e-02 6.22c-02 7% 0.97

32 3.5 COMPARISON OF MEASUREMENTS AND CALCULATIONS In this section, comparisons of the measurement results from the four surveillnce capsules withdrawn to date with the corresponding analytical predictions at the measurement locations are provided. These comparisons are given on two levels. In the first instance, calculations of individual sensor reaction rates are compared directly with the corresponding values obtained from the measured specific activities. In the second case, calculations of fast neutron exposure rates in terms of $(E > 1.0 MeV) and dpa/s are compared with the best estimate results obtained from the least squares evaluation of the three capsule dosimetry results. These two levels of comparison yield consistent and similar results with all measurement to calculation comparisons falling within the 20% limits specified as the acceptance criteria in Regulatory Guide 1.190.

In the case of the direct comparison of measured and calculated sensor reaction rates, the MWC comparisons for fast neutron reactions range from 0.78-1.32 for the 19 samples included in the data set. In the comparisons of best estimate and calculated fast neutron exposure parameters, the cresponding BEIC comparisons range from 0.85 -1.04 for the four surveillance capsules withdrawn to date.

Based on these comparisons, it is concluded that the data comparisons validate the use of the calculated fast neutron xposures provided in Section 3.2 of this report for use in the assessment of the condition of the materials comprising the beltline region of the R. E. Ginna reactor pressure vessel.

. 0A/

33 Table 17 Comparison of Measured and Calculated Neutron Sensor Reaction Rates For In-Vessel Surveillance Capsules V,R, T, and S M/C Ratio Capsule Cu-63(na) Fe-54(np) Ni-58(np) U-238(nf) Np-237(nf) Average  % std dev V 1.02 0.93 0.78 0.89 0.91 10.9 R 1.01 1.01 0.93 0.98 1.32 1.05 14.8 T 1.00 1.01 0.95 1.01 1.29 1.05 12.7 S 1.02 0.95 0.90 0O.94 1.22 1.01 12.7 Average 1.01 0.97 0.89 0.96 1.28 1.01 13.3

% std dev 0.8 4.2 8.4 5.5 .3.9 __ _

Note: The average and % std dev values in bold fice type represent the average and standard deviation of the entire 19 sample threshold foil data set.

Table 18 Comparison of Best Estimate and Calculated Fast Neutron Exposure Rates For I-Vessel Surveillance Capsules V, R, T, and S BE/C Ratio Capsule Neutron Fluence Iron Atom Displacements V 0.86 0.86 R 1.03 1.03 T 1.04 1.03 S 0.98 0.97 Average 0.98 0.97

% std dev 8.7 8.6 4-'4