ML071640339

E-MAIL: (NPA-PD) LRA Letter Part 2

ML071640339
Person / Time
Site:	Pilgrim
Issue date:	05/01/2007
From:	Sanchez E Entergy Nuclear Generation Co
To:	Perry Buckberg NRC/NRR/ADRO/DLR/RLRA
References
TAC MD3698
Download: ML071640339 (35)

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Page 1 of 1 Perry Buckberg - LRA letter part 2 From: "Sanchez, Edward" <esanchl @entergy.com>

To: <phb I @nrc. gov>

Date: 5/1/2007 7:23:37 PM

Subject:

LRA letter part 2

Perry, Attached is part 2 of 2 of the 01 response letter.

Ed Sanchez Pilgrim Licensing file://C:\temp\GW} 00002.HTM 5/31/2007

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W-)0-0-0002, -TMP Page 1 Pg Mail Envelope Properties (4637CBE1.0E9: 9: 61673)

Subject:

LRA letter part 2 Creation Date 5/1/2007 7:22:33 PM From: "Sanchez, Edward" <esanch 1(entergy.com>

Created By: esanch 1(@,entergy.com Recipients nrc.gov OWGWPO01 .HQGWDO01 PHB 1 (Perry Buckberg)

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ATTACHMENT D to Letter 2.07.027 (9 pages)

Torus Room Concrete Base Mat Evaluation (Dr. Franz UIm, M.I.T.)

Preliminary Durability Performance Evaluation of Torus Base Mat in Pilgrim Station By: Franz-Josef Ulm Massachusetts Institute of Technology Cambridge, MA 02139 Date: April 30, 2Q07 Executive Summary: The objective of this report is a preliminary durability evaluation of the effect of the observed groundwater intrusion on the structural performance of the 8 ft thick Reactor Building base mat in Entergy's Pilgrim Station. Based on observations and documents shared by the Entergy team with the author, the following conclusions are drawn:

1. The groundwater migration through the 8ft. thick Reactor Building base mat is a highly localized phenomenon. It is caused by a 25ft hydraulic head difference, pushing groundwater through vertical joints and zones most likely weakened by tensions generated during setting and hydration following the construction. These localized zones are discontinuities equivalent to a vertical cylindrical hole of a maximum diameter of 4 mm (1/61h in). Such small discontinuities that originate from construction joints are inevitable in large-scale concrete engineering operations.

2. This highly localized nature of the zones through which water penetrates, does not compromise the overall structural performance of the Torus base mat: it does neither affect the bulk integrity of the concrete slab, nor the overall compressive and bending load bearing capacity of the reactor foundation.

3. Calcium leaching of the solid concrete is expected to take place in the localized zones through which water penetrates. While this localized calcium leaching does not affect the overall structural performance of the slab, it may contribute to further weakening the construction joints, and may eventually have degraded the grout in the annular space between the 3 in diameter hole and the 2 in diameter Williams rock anchors. A close-up inspection of the grout and bolt is recommended.

4. The lower pH-value of 9.3-9.4 of the water emerging from localized zones along the construction joints, compared to the typical pH- 12 of concrete's bulk pore solution, is consistent with the calcium leaching observation. Its localized occurrence does not compromise the corrosion protection of the steel reinforcement in the slab. A refined corrosion indicator analysis is recommended to confirm the prevention or minimization of reinforcement and anchor bolt corrosion.

5. Changes in environmental conditions (e.g., seasonal changes in water table or a seismic event) that affect the static head that drives the water migration through the concrete would impact the rate of water seepage into the torus room. These affects would be small since, as discussed in the report, the discontinuities in the concrete base mat that are allowing the Water seepage into the torus room are very small. Even if the current very low rate of water intrusion increased by an order of magnitude because of a change in static head there would be no impact on plant safety due to the large size of the torus room.

1. Objective The objective of this report is a preliminary durability evaluation of the Torus Mat in Entergy's Pilgrim Station. This report is based on observations and docunients shared by the Entergy team with the author during and following the visit on Monday, March 19, 2007 to Entergy's offices in the Pilgrim Station.

2. Observations The 8 ft thick Torus mat is part of the reactor building foundation, and surrounds the much thicker core vessel foundation. The Torus mat is divided into 16 Torus room bays, with bay #2 being the most south, and bay #10 the most north, where the ocean is situated.

At any point in time water may be observed on the floor of one or more Torus room bays, puddled at the low point under the invert of the Torus. The maximum water depth may be about 1/z inch or more, but generally limited to the area directly beneath the Torus shell. The wetted areas of the floor have not extended out from under the Torus, and pumping for removal has never been required. Some of remaining bays having no water may have crystal residue indicating they were once wet. Several other bays show no evidence to suggest that they were ever wet. These are bays #16, 1, 2, 3 ind 4, which are all situated on the south side of the reactor foundation, i.e. the furthest away from the sea.

Conditions associated with water seepage into the Torus room have been reported in the early 1980s but may have existed even before then. In the early 1980s seepage was reported at the junction of the drywell pedestal and the Torus room floor, in Bay #15. In the late 1990s, Torus room conditions were evaluated by Engineering. The most probable cause of water seepage was attributed to groundwater by-passing the waterproof membrane that encapsulates the Reactor Building 8 ft. thick base mat. A 25 ft. hydrostatic head forces water through the mat at discontinuities such as construction joints, anchor bolts holes or other features that result in reduced head losses.

This conclusion was based on the following observations and operating experience with membrane designs similar to Pilgrim:

• A portion of the floor in Bay #10, adjacent to the Torus saddle common to Bay #11 (situated on the north side), was dried, cleaned and isolated with a berm. Within the berm is a Williams rock anchor installed in the early 1980s to secure the Torus saddles for pipe break accident uplift conditions. Within about a day, water reappeared on the floor within the berm. There were no visible cracks on the floor surface, suggesting that the water was coming up from the anchor hole drilled in the slab, or from beneath the torus saddle base plate.

" The water was sampled and not found to be contaminated to an extent that would indicate an active plant system leak. Some radioactive contamination was found but this was believed to be the result of plant system leaks that occurred much earlier in the plants operating history.

" There were no active plant system leaks that could result in water on the Torus room floor.

" The seepage rate appears to be at steady state with evaporative losses, hence the floor surface within any bay never fills completely.

3. Methodology Based on these observation, and additional information about the construction history of the reactor building foundation, the tasks of this preliminary durability performance evaluation are:

(1) To identify the most likely cause of the observed seepage into some Torus room bays, and (2) To estimate the effect of this cause on the integrity and performance of the base mat foundation.

(3) Finally, some preliminary conclusions are drawn together with recommendations for further investigations.

• 4. Analysis 4.1 Correlating Construction History and Occurrence of Water Seepage The job drawings of the construction sequence and photos taken during the construction indicate that the Torus mat was constructed fitst before the massive core foundation was put in place. In particular, the construction of the Torus mat was divided in four concrete segments, staring with the North-West Corner, followed by the South-East Comer, the North-East Corner and finally the South-West Corner. Finally, the central core foundation was poured! As such there are a total number of four construction joints between the four mat segments, in addition to the joint between

the core foundation and the surrounding Torus base mat. It has been earlier suggested that those construction joints.

could be the preferential path for water seepage.

In order to check this suggestion, Figure 1 overlays in a plan view the pouring sequence with concrete joints and the bays in which water seepage has been observed. There appears to be a clear correlation between the construction sequence and the locus of occurrence of water. In particular:

(1) The water seepage observed in Bay #6 and #10, situated respectively on the North and the East side of the reactor foundation are situated in bays delineated to the inside by the core-mat construction joint, and separated in the center of the bay by one mat-to-mat construction segment joint.

(2) The water seepage observed in bays #7, #11 and #13 are all delineated to the center by a core-mat construction joint characterized by a kink forming a more-or-less sharp wedged geometrical discontinuity.

There appears to be, therefore, a strong indication in favor of the suggestion that the preferential paths for water seepage are construction joints. From a concrete material perspective, concrete construction joints are well-known to be the weak spot pf any concrete engineering application, due to so-called "wall effect", leading during pouring of the fresh concrete to a higher water-concentration compared to bulk concrete. This higher water concentration available for cement hydration leads to a higher final porosity of the concrete in the immediate surrounding (typically a fraction of Iin) of the joints than in the concrete bulk. Since the flow of water occurs through this porosity, concrete construction joints form a preferential path for fluid conduction, as analyzed later on. With regard to the occurrence of water in specific bay areas, the following additional observations are made:

(1) In bay #6 and # 10, it is most likely that the water seepage originates from the 'T' junction of the mat-mat joint with the core-mat joint.

(2) The fact .hat the core foundation was constructed only after the mat may have had some important implications on the stresses that were most likely generated during the concrete hydration. In fact, since hydration is an exothermic reaction, heat is generated during hydration, which leads to high temperature rises in massive concrete elements. This heat is transported through the concrete surface leading to a cooling over time of the concrete, until the concrete element reaches the ambient temperature [1]. Given the construction process, the base mat which had been poured first, must have been already in a state of cooling, when the much thicker core foundation was poured. Given the massive dimensions of the core foundation, the temperature in the core foundation is expected to have followed an almost adiabatic temperature rise. This temperature differential between the cooler mat and the hot core generates circumferential tension stresses in the mat immediately adjacent to the concrete joint. Indeed, a rough analysis Qf the stresses surrounding the core foundation shows that the circumferential (or hoop) stresses in the immediate vicinity of the joint are in tension if the temperature rise in the core had been twice the temperature rise in the mat. These stresses are amplified, in bay #7,#11,#13 by the presence of the wedge-shaped joint, leading to a significant stress amplification due to stress concentration which is typical for geometrical discontinuities. It is, therefore most likely that those wedge-shaped comers may have been additionally weakened during the differential hydration, forming some preferential path for water seepage.

(3) The fact that most water seepage observed occurred on the North side may well be related to the direction of water flow below the foundation, flowing towards the sea situated on the North side.

North Bay #11 Bay #10

1. Bay#12 C4, '~'4' 44-4-.~ 4.4 ~ .' . ............. 4

-> 4.4 .

.4.-,~. '.f ll./.'. ' '.....

..... 44-~

-3/4'(4

-4. 4 ~4 44 4.

!' ~>4444a,~..' . ........ .

Bay.#141 44444 3 Bay #~4.-4.444..'.4.-.. ..4,'44 [.44 4.............

• '444 4t'~

yBay #7 South Figure 1: Correlating concrete construction sequence with water seepage occurrence: The dashed lines represent construction joints (re-constructed from shop-drawings), the patches represent observed water seepage in specific bays.

It should be noted that those phenomena (higher porosity Of concrete close to joints, microcracks around geometrical discontinuities) are almost inevitable in massive concrete engineering applications due to the high exothermic nature of cement hydration. It is for this reason that concrete design codes specify a minimum amount of steel reinforcement, which needs to bridge concrete construction joints. This reinforcement ensures the structural performance of the slab, despite some localized concrete material weakness along construction joints. This reinforcement, however, cannot eliminate neither the higher material porosity, nor the occurrence of microcracks. It may eventually limit the microcrack opening and propagation; thus ensuring the structural performance of the concrete foundation.

4.2 Seepage through Concrete Bulk and Construction Joints The most likely source of the water seepage observed on the Torus floor is due to groundwater infiltration driven by the hydraulic head of the pore fluid under gravity forces. For purpose of analysis, the following assumptions are I

made:

• The porosity of the 8 ft concrete mat is assumed to be fully saturated by liquid water. This seems to*be a reasonable assumption, given, that the 8 ft thick concrete slab would take some centuries I to dry to a level in equilibrium with the ambient humidity conditions inside the Torus room (see, for instance, [2,3]).

" The 8 ft concrete slab is sufficiently homogeneous throughout its thickness. This allows one to condense the flow into a single material parameter, the intrinsic permeability of concrete. For bulk concrete, typically values for the intrinsic permeability reported in the open literature vary between 10'T-710t16 M 2 , depending on the cqncrete mix proportion and curing conditions (see, for instance, [4]). The concrete composition and massiveness of the Torus base mat is indicative of an intrinsic permeability of 10"17 M2.

Under these assumptions the mass flux rate through the 8 ft concrete slab can be estimated using Darcy's Law. For reference, we first estimate the mass flux rate through the concrete bulk. For an hydraulic head difference of 24.5 ft between the bottom of the slab and the surface and an intrinsic permeability of 10'17 m2, these calculations ýyield a mass flux (per unit surface) through the concrete bulk of Q=0.02 kg/ (m2 day) which amounts to 6.2 kg/ (m yr). Such a small amount of water (which generates daily a water film 17g.m thin) is expected to evaporate almost instantaneously, and it is a clear indication that the water flux through the concrete bulk porosity cannot be at the origin of the observed water seepage. It hints towards a very localized nature of the water penetration. Furthermore, the small value is . benchmark value for a first-order estimate of the size of the joint openings through which the water flow occurs.

Indeed, it was observed that in a previously dried area water reappeared within a day or two generating a water film on-average 1/4 in thick2. This observation translates into a water flux of Q=3.2-6.4 kg/(m 2day), which is substantially greater than the water flux through the concrete bulk.

One can attempt to link this high flux rate to the space through which the flow occurs, by considering the flow of water through a cylinder of radius a in a cross section A (see, for instance [5]). In this simple model representation of the, flow along the joints, the cylinder represents the joint opening, while the cross section represents the wetted surface in consequence of the flow through the cylinder. These calculations provide an upper-bound estimate of the cylinder radius representing the characteristic size of the joint opening, and yield values for the pore throat radius in functions of the wetted surface area. The results which are displayed in Figure 2, indicate that the equivalent cylindrical "joint" opening, is on the order of 1-2 mm (-1/25 - 1/12 in) for wetted surface areas of 5-20 m2 ( 200 ft2) which was observed in some bay areas.

It should be noted that this rough model overestimates the joint opening because it cumulates all possible joint openings into one single cylindrical shape. On the other hand, it provides an upper-bound estimate of the order of magnitude of the joint opening, which is expected to be in the sub-millimeter range,. but substantially greater than the typical capillary pore size of concrete which is in the micrometer range.

In all cases, the order of magnitude estimations of the water flow generated by the hydraulic head difference is clear evidence of a mal-functioning of the water-stop PVC membrane originally designed to control seepage at concrete construction joints.

I Drying of concrete is an extremely slow process. It takes some 10 years to dry a slab of 12 cm thickness exposed on both sides to ambient conditions. The drying duration increases with the square of the thickness. Hence, over the last 35 years, the drying front in the torus room may have reached a depth of x = (35/10)"' x 6 = 11.2 cm = 4.4 in, which is negligible (5%) compared to the base mat thickness of 8 ft. In return, this drying may have caused some microcracking (almost invisible to the eye) in the concrete surface, expanding roughly half the depth of the drying front [3]. This microcracking scales with the mass loss and is little affected by an increase of.the amount of steel reinforcement.

2 John Dyckman, email March 16, 2007.

0.0025 0.0015 c~0.001

- 0_0005

-hi ini 0

0 5 10 15 20 25 Wetted Surface Area A [m2]

Figure 2: Estimated equivalent cylindrical joint size as a function of the wetted surface area. The upper curve is based on water re-appearance within one day, the lower curve on water re-appearance in two days.

4.3 Effect of Calcium Leaching When concrete is put in contact with water having a lower calcium concentration than the equilibrium calcium concentration, calcium is leached from the concrete into the pore solution. The equilibrium concentration of calcium in the pore solution is roughly 480 mg/L (see e.g. [6-7]), meaning that if the calcium concentration in the pore solution is below this threshold value calcium is dissolved ("leached") from the solid into the pore solution. The consequence of leaching is a substantial increase in the porosity (due to the dissolution of Portlandite) and a substantial loss in mechanical stiffness and strength properties [6,8].

Water collected frpm the Torus room and analyzed chemically3 showed a calcium concentration of 230 mg/L, which is smaller than the equilibrium calcium concentration. As a consequence, it cannot be excluded that calcium leaching occurred along the preferential path of groundwater intrusion. In favor of this suggestion is the observation that evaporation residues furthermore show a 31mw% of calcium4.

Fortunately, calcium leaching is a very slow process: the calcium leaching front advances 0.115 mm/'lday; which means that over the last 35 years the leaching may have dissolved the calcium in a layer maximum 13 mm (-1/2 in) thick. In other words, in the life span of the power plant, calcium leaching has no effect on the bulk integrity and structural performance of the concrete foundation. On the other hand, it cannot be excluded that calcium leaching may have contributed to the weakening of the construction joints over the years. In fact, the calcium leaching may have contributed to the joint opening through which groundwater (at a lower calcium concentration than the equilibrium concentration) penetrates. The way by which calcium leaching is most likely to affect the joint opening is sketched in Figure 3, showing dissolution fronts originating from the bottom side around a joint (left). The figure shows that the dissolution generates a vertical wedge-shaped dissolution pattern around the joint, which propagates upward through the slab in a self-similar fashion. These dissolution fronts scale linearly with the water-velocity in the joint and the joint opening [9]; which means that the higher the water flow through the joint, the more advanced the vertical position of the degraded zone.

Measurements of the flow rate through the joints could make it possible to estimate the current height of the leaching front in the 8 ft base mat. In the absence of such measurements, it is not possible to exclude that leaching by groundwater may have reached the grout of the Williams rock anchored (2ft below surface), leaching the calcium of the grout in the annular space between the 3 in diameter hole and the 2 in Williams rock anchor.

3 Northeast laboratory Services Report, dated 03/19/2007.

4 Northeast Laboratory Services Report, dated 03/15/2007.

_41'.

(a) (b) (c) (d)

Figure 3: Dissolution Fronts around a joint through which water flows with a lower calcium concentration than the equilibrium concentration. For a flow velocity of V=10 cm/d and a joint opening of 0.4 nmm, (a) 27 years; (b) 158 years; (c) 318 years; (d) 438 years. The vertical height in this figure is 5 cm (reproduced from

[9]).

4.4 Effect of pH Value Compared to the groundwater (pH - 6.7)5, the water collected in the Torus bay has a pH value of 9.3 - 9.46. This pH value is below..the typical pH- 12 value found in the highly basic interstitial pore solution of concrete, and it is consistent with the observation of a lower-than-equilibrium calcium concentration in the water seepage. Given the very low flow rates through the bulk of the concrete slab, it is reasonable to expect that a high pH- 12 value prevails in the pore solution of the concrete bulk, while the slightly lower pH value of 9.3 - 9.4 only occurs locally in the water penetrating through the joint openings. In all cases, both pH values are in a range that should prevent or minimize reinforcement corrosion, To ascertain this suggestion, a refined chemical analysis of the water may be helpful, one in which not only the pH value is measured (which represents the H' cation- concentration), but as well other quantities, such as the CI7/OW concentration ratio in water. Such corrosion indicators have recently been identified to provide a more comprehensive measure of the onset of corrosion, as it requires a critical amount of CF-to start the corrosion (see, for instance, [10) and references cited herein).

5. Summary of Analysis and Recommendation The analysis of observations and documents relating to the groundwater intrusion into some bays of the Torus room allows for the following preliminary conclusions:

1. The observed groundwater penetration is a highly localized phenomenon. It is caused by the high hydraulic head difference, pushing groundwater through vertical joints and zones most likely weakened by tensions 5 SAIC Report dated July 17, 2006.

6 Northeast laboratory Services Report, dated 03/19/2007

generated during setting and hydration following the construction. These zones are discontinuities equivalent to a vertical cylindrical hole of a 'maximumdiameter of 4 mm (1/6h in).

2. The calcium concentration of the collected groundwater is smaller than the equilibrium calcium concentration in cementitious materials. As a consequence calcium leaching is expected to take place. The highly localized nature of the water penetration ensures that this leaching has no effect neither on the bulk integrity of thie concrete, nor on the overall structural performance of the reactor foundation. (In fact, a vertical cylinder of 4 mm diameter in a - 142 ft x 142 ft foundation slab, will not compromise neither its compressive load distributiop, function nor the bending capacity of the slab.)

3. Thus, the effects of calcium leaching are localized around the vertical joints and other vertical discontinuities present in the slab. Depending on the flow velocity, leaching may have affected the grout in the annular space between the 3 in diameter hole and the 2 in diameter Williams rock anchor, compromising the grout's stiffness and strength. It is recommended to inspect whether the grout has been chemically degraded.

4. The lower pH-value of the collected groundwater is consistent with the observation of a lower-than-equilibrium calcium concentration. Since the phenomenon is localized around some weak spots along the construction joints, it is unlikely that this locally lower pH value substantially increases the risk of reinforcement corrosion. To further prevent and minimize the risk of reinforcement corrosion, a refined chemical analysis is recommended to measure relevant corrosion indicators, such as the CI-/OH-concentration ratio, and to compare those corrosion indicators with now well-established corrosion thresholds

[10]. A close-up inspection of the Williams rock anchors may complement this evidence of a non-detectable corrosion risk.

In summary, the analysis provides evidence that the observed groundwater penetration does not compromise the structural performance of the Torus base mat. It is recommended that any corrective measurement should be based on a prior identification of the exact location of the weak points along the construction joints, and a detailed analysis of the degradation state of both the grout and the bolt of the William rock anchor system.

References

[I ] Ulm FJ, Coussy 0 Couplings in early-age concrete: From material modeling to structural design INTERNATIONAL JOURNAL OF SOLIDS AND STRUCTURES 35 (31-32): 4295-4311 NOV 1998

[2] Acker P, UImFJ Creep and shrinkage of concrete: physical origins and practical measurements NUCLEAR ENGINEERING AND DESIGN 203 (2-3): 143-158 JAN 2001

[3] Ulm FJ, Rossi P, Schaller I, Chauvel D Durability scaling of cracking in HPC structures subject to hygromechanical gradients JOURNAL OF STRUCTURAL ENGINEERING-ASCE 125 (6): 693-702 JUN 1999

[4] Claisse PA, Ganjian E, Adham TA In situ measurement of the intrinsic permeability of concrete MAGAZINE OF CONCRETE RESEARCH 55 (2): 125-132 APR 2003

[5] Dormieux L, Kondo D, Ulm FJ Microporomechanics

'J. WILEY & SONS, CHICHESTER, UK

[6] Ulm FJ, Torrenti JM, Adenot F Chemoporoplasticity of calcium leaching in concrete JOURNAL OF ENGINEERING MECHA.NICS-ASCE 125 (10): 1200-1211 OCT 1999

[7] Mainguy M, Coussy 0 Propagation fronts during calcium leaching and chloride penetration JOURNAL OF ENGINEERING MECHANICS-ASCE 126 (3): 250-257 MAR 2000

[8] Ulm FJ, Heukamp FH, Germaine JT Residual design strength of cement-based materials for nuclear waste storage systems NUCLEAR ENGINEERING AND DESIGN 211 (1): 51-60 JAN 2002

[9] Mainguy M, Ulm FJ, Heukamp FTH Similarity properties of demineralization and degradation of cracked porous materials INTERNATIONAL JOURNAL OF SOLIDS AND STRUCTURES 38 (40-41): 7079-7 100 OCT 2001

[10] Alonso C, Andrade C, Castellote M, et al.

Chloride threshold values to depassivate reinforcing bars embedded in a standardized OPC mortar CEMENT AND CONCRETE RESEARCH 30 (7): 1047-1055 JUL 2000

ATTACHMENT E to Letter 2.07.027 (22 pages)

Structural Integrity Associates Fluence Evaluation for PNPS

StructuraiIntegrity Associates, Inc. File No.: PNPS-27Q-301 CALCULATION PACKAGE Project No.: PNPS-27Q PROJECT NAME: Evaluation of Fluence Issues for Pilgrim Nuclear Power Station CONTRACT NO.: 4500555337, Rev. 1 CLIENT: EntergyýNuclear Northeast PLANT: Pilgrim Nuclear Power Station CALCULATION TITLE: Evaluation of Fluence Issues for PNPS Document Affected Project Manager Preparer(s) &

Domevont Pafed Revision Description Approval Checker(s)

Revision Pages Signature & Date Signatures & Date 1 3, 8, 9, 10, & 12 Update Sections 1.0, 4.5, T. J. Griesbach T. J. Griesbach 5.0, 6.0, and 7.0 4/30/07 4/30/07 G. L. Stevens 4/30/07 Contains References to Proprietary Information Page 1 of 22 F0306-OI RO

StructuralIntegrityAssociates, Inc.

Table of Contents 1.0 INTR O DU CTION .............................................................................................................. 3 2.0 TECHN ICAL APPROA CH ..................................................................................................... 3 3.0 ASSUMPTIONS / DESIGN INPUTS ...................................................................................... 4 4.0 CA L CU L A T ION S ......................................................................................................................... 4 5.0 RESULTS OF AN A LY SIS ..................................................................................................... 9

6.0 CONCLUSION

S AND DISCUSSIONS ................................................................................ 10 7.0 REFE REN CE S ........................................................................................................................... 11 List of Tables Table 1: Beltline Curve A for 54 EFPY with Bias Correction Factor on Fluence ................... 13 Table 2: PNPS ART Calculations for 54 EFPY with Bias Correction Factor on Fluence ....... 14 Table 3: PNPS Charpy Upper Shelf Energy Values for 54 EFPY (Without 1.78 Bias Correction Factor on Fluence) ................................................................................. 15 Table 4: PNPS Charpy Upper Shelf Energy Values for 54 EFPY (With 1.78 Bias Correction Factor on Fluence) ................................................................................. 16 Table 5: PNPS Maximum Projected Fluence and USE Drop for Vessel Beltline Materials .......... 17 Table 6: Effects of Irradiation on RPV Axial Weld Properties ............................................. 18 Table 7: Effects of Irradiation on RPV Circumferential Weld Properties ............................. 19 List of Figures Figure 1: Pressure Test P-T Curve (Curve A) for 54 EFPY with Bias CF on Fluence ................. 20 Figure 2: Calculated Hydrotest Temperature and 1/4t ART versus Fluence ............................... 21 Figure 3: Predicted Decrease in Shelf Energy as a Function of Pct. Copper and Fluence ............ 22 File No.: PNPS-27Q*301 Page 2 of 22 Revision: I Contains References to Proprietary Information F0306-01 RO

Structural IntegrityAssociates, Inc.

1.0 INTRODUCTION

The recent fluence re-evaluation for the Pilgrim Nuclear Power Station (PNPS) reactor pressure vessel (RPV) using the EPRI RAMA code required an increased fluence bias correction factor (CF) of 1.78 to adjust for the benchmarking discrepancy with the cycle 4 surveillance capsule dosimetry results [1]. A rigorous technical explanation for this bias has not been determined. As a result, the NRC will not accept the PNPS fluence calculations for application to future plant operation without further justification. In response to this, Entergy Nuclear Northeast (ENN) has requested SI to perform an additional evaluation to demonstrate adequate vessel life prediction through the extended 60-year operating period with respect to the fluence projections.

Increasing fluence has an effect on the toughness of the RPV materials. This is measured by an increase in the adjusted reference temperature (ART) and a decrease in the upper shelf energy (USE) of the RPV beltline materials. The PNPS FSAR identifies the vessel as being controlling for all reactor pressure boundary carbon steel components [17]. The ASME Code [2] and I OCFR50, Appendix G [3] give criteria for maintaining pressure boundary integrity including the effects of materials degradation due to irradiation damage. Additional evaluations for equivalent margins have been submitted to the Nuclear Regulatory Commission (NRC) and approved for use by boiling water reactors (BWRs) for Charpy USE drop. These equivalent margin analyses for USE are published in BWRVIP-74-A [4]. In addition, BWRVIP-05 [5] provides a technical basis for alternative inspection requirements of the RPV shell welds to eliminate inspections of circumferential welds. The methods and criteria in these documents form the basis for demonstrating vessel integrity margins, including the effects of plant aging due to fluence.

ENN performed an integrated plant assessment (IPE) to extend the operating license of PNPS.

This included a review of the time-limited aging analyses (TLAA) and exemptions to 10CFR50 for the period of extended operation [6]. Increasing fluence is one aspect considered in the TLAAs. The calculated fluence in the vessel using the method from Reference [1] is projected through 54 effective full power years (EFPY) without a bias correction factor. The results of that study are now being reevaluated using assumed fluences greater than the previously calculated results. This analysis for PNPS uses the established methods and Criteria for evaluating embrittlement for fluence levels exceeding the previously projected end-of-license fluence in the vessel.

2.0 TECHNICAL APPROACH The shift in the ART and a decrease in the USE for ferritic materials are predicted by Regulatory Guide 1.99, Revision 2 [7]. The embrittlement trend curves are a function of File No.: PNPS-27Q-301 Page 3 of 22 Revision: I I Contains References to Proprietary Information F0306-O1RO

StructuralIntegrityAssociates, Inc.

copper (Cu) content, nickel (Ni) content, and fluence; different trend curves apply for welds and base metals. The materials in the RPV that must be monitored for irradiation effects are the regions where significant fluence levels are projected (> lxl017 n/cm 2, E > 1 MeV), and those materials are characterized as beltline materials. Analyses of all of the beltline materials for the PNPS vessel determine the weld or plate that is the limiting RPV beltline material. The properties of that limiting beltline material are then used to calculate the operating heatup, cooldown;- and pressure test curves. Those calculations were performed previously for the PNPS RPV for a fluence up to 54 EFPY using fluence projections with and without the 1.78 bias correction factor [8, 9]. The calculations show that there is no RPV integrity concern for 54 EFPY even with the bias corrected fluence.

To further demonstrate that the fluence uncertainty issue for PNPS is not a concern, additional analyses are being performed in this calculation assuming even greater fluence levels in the RPV beyond the 54 EFPY predicted fluence values with the 1.78 bias correction factor. The fluence levels are assumed to increase until a criterion for operability can no longer be maintained. When that limit is determined, the calculated factor on fluence is an indication of the conservatism against brittle fracture of the RPV (or some other criteria) in order to accommodate the observed uncertainty in the fluence calculations.

3.0 ASSUMPTIONS / DESIGN INPUTS

1. The pressure for the pressure test is normal operating pressure (1,035 psig) from Reference [10].

2. The maximum test temperature for the hydrotest is 2127F per the PNPS Technical Specifications

[11]. (Note that this is an operational limit, not a brittle fracture limit.)

4.0 CALCULATIONS 4.1 Maximum Fluence to Perform Hydrotest Irradiation by high energy neutrons raises the RTNDT of the reactor vessel materials. The ART is defined as RTNDT + ARTNDT + Margin in accordance with Regulatory Guide 1.99, Rev. 2 [7].

The pressure-temperature (P-T) curves are developed from the ART value for the vessel material. The calculated hydrotest pressure vs. temperature curve (Curve A) results for 54 EFPY are shown in Table I and in Figure 1 [8]. The PNPS projected values for ARTNDT and ART at 54 EFPY were calculated with the 1.78 bias correction factor on fluence [9]. The projected values of ART are shown in Table 2. The hydrotest pressure is the normal operating pressure, which is 1,035 psig [10]. The system hydrostatic test temperature is calculated to meet the requirements of ASME Section XI, Appendix G, Article G-2400 [2]. The system hydrostatic test should be performed at a temperature not lower than the highest required temperature for any component in the system. For PNPS, the limiting component is the beltline File No.: PNPS-27Q-301 Page 4 of 22 Revision: I Contains References to Proprietary Information F0306-OIRO

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material with the highest ART value at the quarter-thickness (1/4t) location. From Table 2, the limiting materials are the lower intermediate shell longitudinal welds #1 and #3.

The maximum calculated ART value for these welds at 54 EFPY is 122.7°F. This corresponds to a 1/4t fluence value of 1.46x10' 8 n/cm 2, including the 1.78 bias correction factor. The hydrotest temperature at this fluence is 152.5°F. This hydrotest temperature is interpolated linearly from the values from Table 1 as follows:

Hydrotest Hydrotest Temperature Pressure

(°F) (psig) 150 1,007 152.5 1,035 155 1,063 The temperature difference between the 1/4t ART value and the hydrotest temperature is calculated to be 29.8°F. This temperature difference is assumed to be constant for increasing fluence and ART values, so the maximum fluence to conduct the hydrotest can be calculated from the maximum achievable temperature to perform the hydrotest, which is 212'F for PNPS

[11]. The 1/4t fluence and corresponding 1/4t ART for the limiting welds are increased until the hydrotest temperature of 212'F is reached. From the table below it is noted that the maximum 1/4t fluence of 4.12x 1018 n/cm 2 corresponds to a 1/4t ART value of 182.2°F for a hydrotest temperature of 212'F, the maximum temperature to perform the hydrotest at PNPS.

Calculation of Hydrotest Maximum Temperature and Fluence S14t 1/4t Hydrotest Temp. Fluence Fluence ART Temp. Difference Ratio (n/cm2) (OF) (OF) (OF) 1.46E+18 1.22.7 152.5 29.8 1.00 2.00E+18 139.6 169.4 29.8 1.37 3.OOE+18 162.9 192.7 29.8 2.05 4.O0E+18 180.4 210.2 29.8 2.74 4.12E+18 182.2 212 29.8 2.82 maximum fluence to conduct hydrotest < 212°F 4.50E+18 187.8 217.6 29.8 3.08 5.00E+18 194.4 224.2 29.8 3.42 The calculated hydrotest temperature and 1/4t ART values versus fluence are shown in Figure 2.

A fluence ratio of 2.82 is the ratio of the maximum 1/4t fluence at-the limiting vessel beltline welds compared to the 54 EFPY fluence with the 1.78 bias correction factor. In other words, the fluence with the 1.78 bias correction factor would have to be increased by an additional factor of 2.82 before the limiting hydrotest temperature of 212F is reached.

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4.2 Maximum Fluence to Maintain Charpy Upper Shelf Energy Appendix G of 10CFR50 requires that reactor vessel beltline materials "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." Regulatory Guide 1.99, Rev. 2, RadiationEmbrittlement of Reactor Vessel Materials, defines the method for predicting upper shelf energy drop in terms of a percentage from the unirradiated value. Figure 3 shows the predicted Charpy upper shelf energy for welds and base metals as a function of copper content and fluence.

The predicted Charpy upper shelf energy (CvUSE) values for PNPS at 54 EFPY were determined previously for the PNPS license renewal project [6]. The predicted CvUSE values based on the Regulatory Guide 1.99 Position 1 method are shown in Table 3. The predicted values for CvUSE using the 54 EFPY fluences with the 1.78 bias correction factor are shown in Table 4.. It is noted that all projected USE values are above 50 ft-lbs, even with the 1.78 bias correction factor on fluence. The USE limit shows a minimum fluence ratio of 4.9 for the projected fluence to reach 50 ft-lbs for the lower intermediate shell axial welds, as shown in Table 5. Because the USE values are always greater than 50 ft-lbs., the equivalent margin method of BWRVIP-74-A is not required.

4.3 Maximum Fluence Bounded by the Reactor Vessel Weld Failure Probability The BWVIP recommendations for inspection of reactor vessel shell welds in BWRVIP-05 [5]

are based on generic analyses supporting a Safety Evaluation Review (SER) conclusion that the generic plant axial weld failure rate is no greater than 5x10-6 per reactor year [ 12] at the end of 40 years. BWRVIP-05 showed that this axial weld failure rate is orders of magnitude greater than the 40-year end-of-life circumferential weld failure probability, and used this analysis to justify relief from inspection of the circumferential welds as described above.

PNPS received relief from the circumferential weld inspections for the remainder of the original 40-year operating term [13]. The basis for this relief request was a plant specific analysis that showed the limiting conditional failure probability for the PNPS circumferential welds at the end of the original operating term were less than the values calculated in the BWRVIP-05 SER [12].

Table 6 contains a comparison of the PNPS reactor vessel limiting axial weld parameters to those used-in the NRC analysis. The data in column two (CE) is from Table 2.6-5 of the NRC SER for BWRVIiP-05 [12]. The data in the third column (PNPS) is the projected 54 EFPY data for PNPS without the 1.78 bias correction factor on fluence [6]. (For consistency with the NRC evaluation, the RTNDT is calculated without the margin term.) The data in column four (PNPS with Bias CF) is the projected 54 EFPY data for PNPS with the 1.78 bias correction factor on fluence.

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"Columnfive (PNPS Limit) shows the maximum fluence and RTNDT to assure that the limiting axial weld remains bounded by the value of 172.4°F determined from the CEOG and accepted in the SER [12]. The maximum ID fluence of 8.48x10s 8 n/cm 2 gives a fluence ratio of 4.18 compared to the 54 EFPY fluence with the 1.78 bias correction factor.

Table 7 contains a comparison of the PNPS reactor vessel limiting circumferential weld parameters to those used in the NRC analysis. The data in column two (CE) is from Table 2.6-5 of the NRC SER for BWRVIP-05 [12]. The data in the third column (PNPS) is the projected 54 EFPY data for the PNPS circumferential weld without the 1.78 bias correction factor on fluence [6]. The data in column four (PNPS with Bias CF) is the projected 54 EFPY data for the PNPS circumferential weld with the 1.78 bias correction factor on fluence. Column five (PNPS Limit) shows the maximum fluence and RTNDT to assure that the PNPS circumferential weld remains bounded by the value of 128.5°F determined from 2

9 the CEOG and accepted in the SER [12]. The maximum ID fluence of 1.14x101 n/cm gives a fluence ratio of 7.35 compared to the 54 EFPY fluence with the 1.78 bias correction factor.

PNPS obtained relief from the examination of RPV circumferential welds related to the augmented shell weld examination requirements contained in 10CFR50.55a(g)(6)(ii)(A)(5).

The reduction in scope of these inspections from essentially 100 percent of all RPV shell welds to examination of essentially 100 percent of the axial welds and essentially zero percent of the circumferential welds was based on the NRC staff determination that the conditional probability of failure for these welds was within the acceptable limits at the expiration of the current operating license [ 13]. The results given in Tables 6 and 7 show that the bounding reactor vessel weld conditional failure probabilities can be maintained well beyond the 54 EFPY projected fluences and ART values for the PNPS vessel. The relatively large calculated fluence ratios shown in these tables indicate that the criteria for relief from the circumferential vessel weld inspections will not be the limiting factor for fluence margin in the PNPS RPV.

4.4 Effect of Fluence on Evaluation of N2 Nozzles The fluence levels in the N2 nozzles are relatively low compared to the peak fluence in the beltline. These fluences shown in the table below were obtained from the RAMA code fluence calculation [9, 14].

54 EFPY Fluence @ 114t 54 EFPY Fluence @ 1/4t (w/o 1.78 bias CF) (with 1.78 bias CF)

(nfcmA2) (n/cmA2)

Recirc. inlet (N2) nozzles 2.02E+17 3.60E+17 Limiting Axial Welds :, 8.18E+17 1.46E+18 File No.: PNPS-27Q-301 Page 7 of 22 Revision: 1 Contains References to Proprietary Information F0306-01RO

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The effect of the increasing fluence on the calculated ART values for the limiting weld and the N2 nozzle' is shown below. The ART values for the A508-2 nozzle forgings was estimated using upper bound Cu = 0.35, Ni = 0.85, and an initial RTNDT of 0°F [14].

54 EFPY ART @ 1/4t 54 EFPY ART @ 1/4t (w/o 1.78 bias CF) (with 1.78 bias CF)

(0F) (°F)

Recirc. inlet (N2) nozzles 77.0 94.7 Limiting Axial Welds 95.3 122.7 Structural Integrity Associates recently performed an evaluation of the recirc. inlet nozzles using best estimate copper and nickel chemistry values of Cu = 0.15 wt%, Ni = 0.85 wt% [ 14].

Using these best estimate values, the calculated ART values for the nozzles are as follows:

54 EFPY ART @ 1/4t 54 EFPY ART @ 1/4t (wlo 1.78 bias CF) (with 1.78 bias CF)

_(*F) (*F)

Re'circ. inlet (N2) nozzles 39.9 56.4 Limiting Axial Welds 95.3 122.7 From the comparison of the ART values for the recirc. inlet nozzles and the limiting axial welds, the recirc. inlet nozzle embrittlement levels are well below the projected ART values for the limiting axial welds. This is mainly because of significantly lower fluences at the height of the nozzles compared to the active core region. Thus, there is no impact of fluence uncertainty for this evaluation, and it is determined that the nozzles will not become the limiting beltline materials for P-T limits or hydrotest conditions as fluence levels are increased.

4.5 Effect of Fluence on RPV Internals 4.5.1 Top Guide 2

BWRVIP-26 calculated the minimum top guide fluence for 32 EFPY (40 years) as 4xl 0 1 n/cm 2 [15]:- The threshold for IASCC is 5xl 0 20 n/cm 2, and the PNPS top guide fluence will exceed this threshold [6]. Therefore, PNPS must manage IASCC of the top guide assembly.

PNPS has implemented the inspection recommendation in BWRVIP-26 through the BWR Vessel Internals Program [16]. The BWR Vessel Internals Program will adequately manage the effects of aging on the top guide for the period of extended operation. The top guide does not affect the operating P-T limit curves, and there is no criterion on fluence that would further limit the operation of the top guide structure.

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4.5.2 Core Shroud The core shroud is a BWR component that is known to be susceptible to aging effects. Section 3.8.12 of the PNPS License Renewal Project, TLAA and Exemption Evaluations [6] addresses the time limited aging analyses of the core shroud. A review of the analyses related to the core shroud found that the only TLAA involves the fatigue analysis and calculation of cumulative usage factors (CUFs) for the shroud repair. The core shroud does not affect the operating P-T

'limit curves, and there is no criterion on fluence that would further limit the operation of the core shroud structure.

5.0 RESULTS OF ANALYSIS The effects of increased fluence beyond the projected 54 EFPY fluence calculations for the PNPS RPV are summarized below for each of the potential aging effects. The results are compared to determine the minimum acceptable fluence ratio. This is the fluence multiplier that could be achieved compared to the 54 EFPY fluence with the 1.78 bias correction factor, and is the measure of tolerance on fluence before a limit is reached that would exceed a Code limit, regulatory criterion, or service limit.

Effect of Fluence on Acceptable Fluence Ratio Hydrotest Temperature 2.82*

Charpy Upper Shelf Energy 4.86 RPV Axial Weld Failure Probability 4.18 RPV Circ. Weld Failure Probability 7.35 Evaluation of N2 Nozzles Bounded by beltline

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6.0 CONCLUSION

S AND DISCUSSIONS Fluence contributes to changes in the vessel beltline material properties. These changes are measured by the shift in RTNDT or the drop in USE of the ferritic materials (i.e., welds, plates, and forgings). The analyses using projected fluence values for license renewal (54 EFPY) for PNPS show no limitations due to embrittlement concerns for the vessel. Considering increasing fluence levels, the RPV analyses demonstrate that the Code and regulatory criteria can be met for operation well beyond this maximum fluence level by a factor of 2.82 (or greater) on the 54 EFPY fluence including a bias correction factor of 1.78.

The limiting condition for the vessel is the temperature required to perform the ASME Code hydrotest. The temperature to perform the hydrotest is prescribed by ASME Section XI, Article.G-2400 that requires a safety factor of 1.5 on the pressure stress intensity to prevent brittle fracture of the vessel during this test. The maximum temperature limit for the hydrotest of 212'F in the PNPS Technical Specifications is an administrative limit; it may be possible to perform the test at higher temperatures which would allow for even higher fluence levels.

These analyses demonstrate that there is a considerable tolerance on the acceptable range of fluence. This is exemplified by the difference between the fluence for the maximum predicted levels of embrittlement and the limiting criteria for operability, a difference large enough to accommodate the uncertainties on the calculated fluence for PNPS.

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7.0 REFERENCES

1. ENT-FLU-00 1-R-001, "Pilgrim Nuclear Power Station Reactor Pressure Vessel Fluence Evaluation at End of Cycle 15 and 54 EFPY," TransWare Enterprises Inc.,

December 2005, Proprietary, (SI File No. PNPS-22Q-201).

2. ASME Boiler and Pressure Vessel Code,Section XI, Appendix G, 1998 Edition, 2000 Addenda.

3. Code of Federal Regulations, 10 CFR, Part 50, Appendix G, Federal Register: January 2005.

4. BWRVIP-74-A, "BWR Reactor Vessel Inspection and Flaw Evaluation Guidelines for License Renewal, BWR Vessel and Internals Project," EPRI Proprietary, 1008872, June. 2003, (SI File No. BWRVIP-01-274AP).

5. BWRVIP-05, "BWR Reactor Pressure Vessel Shell Weld Inspection Recommendations (BWRVIP-05)," BWR Vessel and Internals Project," EPRI Proprietary, EPRI TR-105697, September 1995. (SI File No. BWRVIP-01-205P).

6. PNPS License Renewal Project, TLAA and Exemption Evaluations, LRPD-03, Revision 0, January 2006, (SI File No. PNPS-27Q-202).

7. NRC Regulatory Guide 1.99, Rev. 2, "Radiation Embrittlement of Reactor Vessel Materials," May 1988.

8. SI Calculation PNPS-03Q-301, Revision 1, "Development of Pressure Test (Curve A)

P-T Curves," 1/30/06.

9. SI Calculation PNPS-22Q-301, "ARTNDT and ART Evaluation," 1/20/06.

10. Email from Bryan Ford to Raymond Pace, Timothy J. Griesbach, and Gary L. Stevens,

Subject:

Maximum Pressure Test Temperature, dated 3/8/07, (SI File No. PNPS-27Q-205).

11. PN-PS Technical Specifications, Revision 274, Amendment No.'s 224 and 225, Limiting Conditions for Operation, 3.14 Special Operations, A. Inservice Hydrostatic and Leak Testing Operation, (SI File No. PNPS-27Q-206).

12. BWRVIP-05 SER (Final), USNRC letter from Gus C. Lainas to Carl Terry, Niagara Mohawk Power Company, BWRVIP Chairman, Final Safety Evaluation of the BWR Vessel and Internals Project BWRVIP-05 Report, (TAC No. M93925), July 28, 1998, (SI File No. BWRVIP-01-205P).

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13. Letter from J. Boska (NRC) to M. Bellamy (ENGC), Pilgrim Nuclear Power Station -

Pilgrim Relief Request No. 28, Relief from ASME Code,Section XI, Examinations of Reactor Pressure Vessel Circumferential Shell Welds (TAC No. MB6074), April 11, 2003, (SI File No. PNPS-27Q-208).

14. SI Calculation PNPS-22Q-302, "N2 Nozzle Evaluation," 2/21/06.

15. BWRVIP-26, "BWR Top Guide Inspection and Flaw Evaluation Guidelines (BWRVIP-26)," EPRI Proprietary, EPRI Report TR-107285, December 1996, (SI File No.

BWRVIP-0 I-226P).

16. Engineering Report PNPS-EP-06-00001, Revision 0, "Reactor Vessel Internals Inspection Program," (SI File No. PNPS-27Q-207)

17. Pilgrim Nuclear Power Station Final Safety Analysis Report, Section 3.3, Reactor Vessel Internals Mechanical Design, and Section 4.2, Reactor Vessel and Appurtenances Mechanical Design, (SI File No. PNPS-27Q-209).

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Table 1: Beltline Curve A for 54 EFPY with Bias Correction Factor on Fluence [81 Pressure-Temperature Curve Calculation

-(Pressure Test = Curve A)

(NOTE: THE ARTNDT includes a calculated bias on fluence of 1.78.)

Inputs: Plant = 1ýflgfih Component = GBein Vessel thickness, t = 5.31 inches, so Jt 2.352 'inch Vessel Radius, R = 1.91l inches A R T NDT 122.7 *F . . . . . .> " EU*J' PY ,"

Cooldown Rate, CR 0 ý°F/hr 12 KiT ,0 ksilinch (From Appendix G, for cooldown rate above)

AT1/4t 00 °F (no thermal for pressure test)

Safety Factor 1 J(for pressure test)

Mm = 1 (From Appendix G, for inside surface axial flaw)

Temperature Adjustment = -°F Height of Water for a Full Vessel = ;inches Pressure Adjustment 18.3 psig (hydrostatic pressure for a full vessel at 70°F)

Hydro Test Pressure = ,565 psig Flange RTNDT 10.0 'F Fluid Calculated Adjusted Adjusted Temperature 1/4t P re ssure Temperature Pressure for T Temperature Kic Kip P for P-T Curve P-T Curve

(°F) (°F) (ksi*lnchlI 2 ) (ksi*inch"') (psig) ('F) (psig) 70.0 70.0 40.43 26.95 0 70.0 0 70.0 70.0 40.43 26.95 601 70.0 583 75.0 75.0 41.19 27.46 612 75.0 594 80.0 80.0 42.03 28.02 625 80.0 606 85.0 85.0 42.95 28.64 638 85.0 620 90.0 90.0 43.98 29.32 654 90.0 635 95.0 95.0 45.11 30.08 671 95.0 652 100.0 100.0 46.37 30.91 689 100.0 671 105.0.. 105.0 47.75 31.84 710 105.0 691 110.0 110.0 49.28 32.86 733 110.0 714 115.0 115.0 50.97 33.98 758 115.0 739 120.0 120.0 52.84 35.23 785 120.0 767 125.0 125.0 54.91 36.61 816 125.0 798 130.0 130.0 57.19 38.13 850 130.0 832 135.0 135.0 59.72 39.81 888 135.0 869 140.0 140.0 62.51 41.67 929 140.0 911 145.0 145.0 65.59 43.73 975 145.0 957 150.0 150.0 68.99 46.00 1026 150.0 1,007 155.0 155.0 72.76 48.51 1082 155.0 1,063 160.0' 160.0 76.92 51.28 1143 160.0 1,125 165.0 165.0 81.52 54.34 1212 165.0 1,193

.170.0 170.0 86.60 57.73 1287 170.0 1,269 175.0 175.0 92.21 61.48 1371 175.0 1,352 180.0 180.0 98.42 65.61 1463 180.0 1,445 185.0 185.0 105.28 70.19 1565 185.0 1,547 190.0 190.0 112.86 75.24 1678 190.0 1,659 195.0 195.0 121.24 80.83 1802 195.0 1,784 File No.: PNPS-27Q-301 Page 13 of 22 Revision: 1 Contains References to Proprietary Information F0306-O1RO

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Table 2: PNPS ART Calculations for 54 EFPY with Bias Correction Factor on Fluence [91 IPilgrim RPV MaterialART Calculations (54 EFPY)

(NOTE: This table covers all RPV matenals bwthan exposednuence, E. I M1eV,of greaterthan .Ox10 " ntcm'.) includes .78calculatedbias on fluence Estimated Chemistry Chemistry Adjustments For 114t Description Piece Code Heat Initial RTNo- Factor ARTCDT Margin Terms ARTT No. No. No. (=F) Cu(wt%) Ni(wt%) ("F) ('F) oA(7F) a,(°F) (°F)

Lower Shell#1 I 337-01A G-3109-2 C-2957-2 0 0.10 0.47 65.0 30.5 15.3 0.0 61.1 LowerShell#2 337-01B G-3109-1 C-2957-1 -3 0.10 0.48 65.0 30.5 15.3 0.0 58.1 LowerShell#3 337-01C G-3109-3 C-2973-1 -4 0.11 0.63 74.5 35.0 17.0 0.0 65.0 Lower-int. Shell #1 337-03A G-3108-3 C-2945-2 -12 0.10 0.66 65.6 34.3 17.0 0.0 56.3 Lower-inLShell#2 337-03B G-3108-1 C-2921-2 -30 0.14 0.60 100.0 52.2 17.0 00 562 Lower-Int. Shell #3 337-03C G-3108-2 C-2945-1 -7 0.10 0.65 65.5 34.2 17.0 0.0 61.2 Estimated Chemistry Chemistry Adiustments For 114t Description Seam Heat Flux Type & InItiai RT.r Factor ARTem Margin Terms ART T No. No. Lot No. (IF) Cu (wt%) NI(wt%) ("F) ("F) ua (*F) a,(*F) (=F)

L. Int.Shei Long. Weld #1 1-338A 27204112008 Linde1092#3774 -48 0.219 0.996 231.1 114.7 28.0 0.0 122.7 L. It Shell Long. Weld #2 1-3388 27204/12008 Linde 1092 #3774 -48 0.219 0.996 231.1 78.4 28.0 0.0 86.4 L. Int.ShellLong.Weld#3 1-338C 27204112008 Linde1092#3774 -48 0.219 0.996 231.1 114.7 28.0 0.0 122.7 L. ht./L. Shell Girth Weld 1-344 21935 Linde 109243869 -50 0.183 0.704 172.2 75.4 28.0 0.0 81.4 Lower Shell Long. Weld #1 2-338A 27204 Linde 1092 #3714 -34 0.203 1.018 226.8 83.5 28.0 0.0 105.5 Lower Shell Long. Weld #2 2-3388 27204 Linde 1092 #3714 -34 0.203 1.018 226.8 96.5 28.0 0.0 118.5 Lower Shell Long. Weld #3 2-338C 27204 Linde 1092 #3714 -34 0.203 1,018 226.8 87.8 28.0 0.0 109.8 Fluence information lsee Note 2):

Calculated Fluence Bias= 1.78 Wall Thickness (inches) Fluence at ID Attenuation, 114t Fluence @ 114t Fluence Factor, FF 2 2 Location Full (f) 14t (nlcm ) e""04. (nlcm ) 002M.-0.la9sat LowerShelI#1 5.531 1.383 1.80E+18 0.718 1.29E+18 0.470 LowerShell#2 . 5.531 1.383 1.80E+18 0.718 1.29E+18 0.470 LowerShell#3 5.531 1.383 1.80E+18 0.718 1.29E+18 0.470 Lower-int. Shell#1 5.531 1.383 2.28E+18 0.718 1.63E+18 0.522 Lower-int Shell #2 5.531 1.383 2.28E+18 0.718 1.63E+18 0.522 Lower-int. Shell#3 5.531 1.383 228E+18 0.718 1.63E+18 0.522 L. int.SheltLong. Weld #1 5.531 1.383 2.03E+18 0.718 1.468+18 0.496 L. It Shell Long. Weld #2 5.531 1.383 9.20E+17 0.718 6.60E+17 0.339 I L. Int Shell Long. Weld #3 L. int./L. Shell Girth Weld Lower Shell Long. Weld #1 Lower Shell Long. Weld #2 Lower Shell Long. Weld #3 5.531 5.531 5.531 5.531 5.531 1.383 1.383 1.383 1.383 1.383 2.03E+18 1.55E+18 1.08E+18 1.45E+18 1.20E+18 0.718 0.718 0.718 0.718 0.718 1.46E+18 1.11E+18 7.77E+17 1.04E+18 8.58E÷17 0.496 0.438 0.368 0.425 0.387 Notes: 1. Material information taken from SLAReport No. SIR-00-082. Revision 0, "Updated Evaluation of Reactor Pressure Vessel Materials Properties for Pilgrim Nuclear Power Station," August 2000. Tables 3-1 through 3-12.

2. Fluence values from Transware Report No. ENT-FLU-001-R-001, Revision 0. "Pilgrim Nuclear Power Station Reactor Pressure Vessel Fluence Evaluation." Tables 7-3 and 7-4, and are multiplied bya calculated bias of 1.78.

3. RPV minimum thickness = 5 17/32" per Section 3.3.2 of SIR-00-082, Revision 0.

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Table 3. PNPS Charpy Upper Shelf Energy Values for 54 EFPY (Without 1.78 Bias Correction Factor on Fluence) [61 Material Description 54 EFPY Projection 1/4 T ffuenoe % Drop in Reactor Vessel Material Unirradiated (1019 n/cm4) USE USE (1D4 T)

Beltllne Region Location Matl Type Identification Heat # %Cu CvUSE Lower Intermediate Shell A533B G-3108.1 C-2921-2 0.14 81 0.084 12.79% 70.6 Lower Intermediate Shell A533B G-3108-2 C-2945-1 0.10 80 0.084 10.57% 71.6 Lower Intermediate Shell A533B G-3108-3 C-2945-2 0.10 81 0.084 10.57% 72.4 Lower Shell A533B G-3109-1 C-2957-1 0.10 76 0.001 9.79% 68.6 Lower Shell A533B G-3109-2 C-2957-2 0.10 79 0.061 9.79% 71.3 Lower Shell A533B G-3109-3 C-2973-1 0.11 72 0.061 10.31% 64.6 Lower Int/Lower Shell Circ Weld Linde 1092 1-334 21935 0.18 75 0.057 16.39% 62.7 Lower Int Shell Axial Welds Unde 1092 1-336A, B,C 27204-12008 0.22 75 0,076 19.52% 60.4 Lower Shell Axial Welds Linde 1092 2-338AB,C 27204 0.20 75 0.050 16.87% 62.3 File No.: PNPS-27Q-301 Page 15 of 22 Revision: 1 Contains References to Proprietary Information F0306-O1 RO

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Table 4. PNPS Charpy Upper Shelf Energy Values for 54 EFPY (With 1.78 Bias Correction Factor on Fluence)

Material Description 54 EFPY Projection (with 1.78 bias CF on fluence)

Reactor Vessel Matl Mati Heat# %Cu Unirr. 1/4t fluence  % Drop USE Beltline Region Location Type Ident. CvUSE (1OAJ 9 n/cm2) in USE @ 1/4t Lower Intermediate Shell A533B G-3108-1 C-2921-2 0.14 81 0.129 14.3 69.4 Lower Intermediate Shell A533B G-3108-2 C-2945-1 0.10 80 0.129 11.7. 70.6 Lower Intermediate Shell A533B G-3108-3 C-2945-2 0.10 81 0.129 11.7 71.5 Lower Shell A533B G-3109-1 C-2957-1 0.10 76 0.163 12.3 66.7 Lower Shell A533B G-3109-2 C-2957-2 0.10 79 0.163 12.3 69.3 Lower Shell A533B G-3109-3 C-2973-1 0.11 72 0.163 13.1 62.6 Lower Int./Lower Shell Circ. Weld Linde 1092 1-334 21935 0.183 75 0.111 19.6 60.3 Lower In. Shell Axial Welds Linde 1092 1-338A,B,C 27204/12008 0.219 75 0.146 23.2 57.6 Lower Shell Axial Welds Linde 1092 2-338A,B,C 27204 0.203 75 0.104 20.5 59.6 File No.: PNPS-27Q-301 Page 16 of 22 Revision: 1 Contains References to Proprietary Information F0306-01 RO

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Table 5. PNPS Maximum Projected Fluence and USE Drop for Vessel Beltline Materials Material Description Maximum Projected Fluence and USE Drop Reactor Vessel Matl Matl %Cu Unirr. 1/4t fluence Max. % Drop Min. USE Fluence Beltline Region Location Type Ident. CvUSE (10^19 n/cm2) in USE @ 1/4t Ratio Lower Intermediate Shell A533B G-3108-1 0.14 81 > 6.0 38.3 50.0 > 46.5 Lower Intermediate Shell A533B G-3108-2 0.10 80 > 6.0 37.5 50.0 > 46.5 Lower Intermediate Shell A533B G-3108-3 0.10 81 > 6.0 38.3 50.0 > 46.5 Lower Shell A533B G-3109-1 0.10 76 > 6.0 34.2 50.0 > 36.8 Lower Shell A533B G-3109-2 0.10 79 > 6.0 36.7 50.0 > 36.8 Lower Shell A533B G-3109-3 0.11 72 > 6.0 30.6 50.0 > 36.8 Lower int./Lower Shell Circ. Weld Linde 1092 1-334 0.183 75 1.11 33.3 50.0 10 Lower Int. Shell Axial Welds Linde 1092 1-338AB,C 0.219 75 0.71 33.3 50.0 4.86*

Lower Shell Axial Welds Linde 1092 2-338A,BC 0.203 75 0.86 33.3 50.0 8.3

• limiting fluence ratio to reach 50 ft-lbs CvUSE = (0.71E 19)/(0.146E 19) = 4.86 File No.: PNPS-27Q-301 Page 17 of 22 Revision: I Contains References to Proprietary Information F0306-01RO

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Table 6. Effects of Irradiation on RPV Axial Weld Properties Limiting Axial Welds - Lower Int. Long. Welds #1 and #3 Wire Heat/Lot (27204/12008, Lot No. 3774)

Pldnt CE (CEOG) PNPS PNPS with PNPS Limit

_______ _______ __ ____ ______Bias CF _ _ _ _ _ _ _

Parameter Description USNRC Data for Data for Data for Limiting Plant- axial weld axial weld axial weld Specific Data (no bias CF) (1.78 bias CF) (limiting fluence)

EFPY 64 54 54 >54 Initial (unirradiated) reference 0 -48 -48 -48 temperature (RTndt), OF Neutron fluence al the end of the requested relief period (Peak 4.OOE+ 18 1.14E+18 2.03E+ 18 8.48E+18*

Surface Fluence in the Beltline),

n/cmA2 FF = Fluence factor (calculated 0.746 0.444 0.573 0.954 per Reg. Guide 1.99, Rev. 2)

Weld Copper content, wt. % 0.219 0.219 0.219 0.219 Weld Nickel content, wt% 0.996 0.996 0.996 0.996 CF = Chemistry Factor 231.1 231.1 231.1 231.1 Increase in reference temperature 172.4 102.9 132.4 220.4 (ARTndt), °F (= FF*CF)

Mean adjusted reference 172.4 54.9 84.4 172.4 temperature (ART), 'F

(= RTndt +ARTndt)

• Fluence ratio = (8.48E 18)/(2.03E 18) = 4.18 File No.: PNPS-27Q-301 Page 18 of 22 Revision: 1 Contains References to Proprietary Information F0306-O1RO

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Table 7. Effects of Irradiation on RPV Circumferential Weld Properties Limiting Circ. Weld - Lower Int.-to-Lower Shell Circ. Weld 1-344 Wire Heat/Lot (21935, Lot No. 3869)

Plant CE (CEOG) PNPS PNPS with PNPS Limit Bias CF Parameter Description USNRC Data for Data for Data for Limiting Plant- circ. weld circ. weld circ. weld Specific Data (no bias C.F) (1.78 bias CF) (limiting fluence)

EFPY .64 54 54 >54 Initial (unirradiated) reference 0 -50 -50 -50 temperature (RTndt), OF Neutron fluence at the end of the requested relief period (Peak 4.OOE+18 8.69E+17 1.55E+18 1.14E+19*

Surface Fluence in the Beltline),

n/cmA2 FF = Fluence factor (calculated 0.746 0.389 0.510 1.037 per Reg. Guide 1,99, Rev. 2)

Weld Copper content, wt. % 0.183 0.183 0.183 0.183 Weld Nickel content, wt% 0.704 0.704 0.704 0.704 CF = Chemistry Factor 172.2 172.2 172.2 172.2 Increase in reference temperature 128.5 67.1 87.9 178.5 (ARTndt), 'F (= -FF*CF)

Mean adjusted reference 128.5 17.1 37.9 128.5 temperature (ART), °F

(= RTndt +ARTndt)

• Fluence ratio = (l.14E19)/(l.55E18) = 7.35 File No.: PNPS-27Q-"301 Page 19 of 22 Revision: 1 Contains References to Proprietary Information F0306-O1RO

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Figure 1: Pressure Test P-T Curve (Curve A) for 54 EFPY with Bias CF on Fluence [81 1,200 1 I I t T7TI1~ 2ziz

J, I J.- I I I 1,100 -+- V F-ý- -+- ---,F-I i ý i 1 1 t 1 1 1 1 1 i  ;

I ~*II 1,000 I-.--.

-i-4 900

0. 800 700 r-... .... .-

0 . z z z z,~t+ .i.........

U) 600

--... I i .t.. - - . . . ----' . .. ..... ....... ... .. -

T . . [r 500 7 -- -Bt H d LU 400

!i I i' f i I1 CL 300 200

. *. ,- - - r, Vessel Upper - - ,-4=

100 -*--:-~~~ ~ ................

~ ~ ~ ~ ~ ~ ~ ~ oto Head..............

0 0 50 100 150 200 250 MINIMUM REACTOR VESSEL METAL TEMPERATURE (°F)

PNPS Pressure Test Curve (Curve A), 54 EFPY (NOTE: The fluence used on the beitline curve is increasedby a calculatedbias on fluence of 1.78.)

File No.: PNPS-27Q-301 Page 20 of 22 Revision: 1 Contains References to Proprietary Information F0306-O1RO

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Figure 2: Calculated Hydrotest Temperature and 1/4t ART versus Fluence 250 200 150 LL E 100 50' 0

0 1E+18 2E+18 3E+18 4E+18 5E+18 6E+18 1/4t Fluence (n/cmA2)

File No.: PNPS-27Q-301 Page 21 of 22 Revision: 1 Contains References to Proprietary Information F0306-01 RO

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Figure 3. Predicted Decrease in Upper Shelf Energy as a Function of Pct. Copper and Fluence (from Reg. Guide 1.99, Rev. 2 [71) 100 8

7 6

a)

U a) 1~

a.

L.

a)'

C w

.4-10 a) 8 C,)

C a) 6 In II 5 a)

U a) 4

• 92 3 4 5 6 75a9 1.OE+17 1.6E+18 1.OE+19 1.OE+20 2

FLUENCE, n/cm (E > 1MeV)

File No.: PNPS-27Q-301 Page 22 of 22 Revision: 1 Contains References to Proprietary Information F0306-OIRO