ML22048A003

SwRI-02-5928, Reactor Vessel Material Surveillance Program Analysis of Capsule T

ML22048A003
Person / Time
Site:	Cook
Issue date:	09/16/1981
From:	Morris E Southwest Research Institute
To:	Plant Licensing Branch III
Wall S
References
SwRI-02-5928
Download: ML22048A003 (103)

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{{#Wiki_filter:REACTOR VESSEL MATERIAL SURVEILLANCE PROGRAM FOR DONALD C. COOK UNIT NO.2 ANALYSIS OF CAPSULET C \\" -. 1_ by E. B. Norris r',;, '

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\\ i.:, !'j FiNAL REPORT 'Z \\t*) I,'- SwRI Project No. 02-5928 Bridgman, Michigan 49106 j \\;, September 16, 1981 \\ ; SOUTHWEST RESEARCH INSTITUTE , SAN ANTONIO HOUSTON

t ( SOUTHWEST RESEARCH INSTITUTE Post Office Drawer 28510, 6220 Culebra Road San Antonio, Texas 78284 REACTOR VESSEL MATERIAL SURVEILLANCE PROGRAM FOR DONALD C. COOK UNIT NO.2 ANAL VSIS OF CAPSULE T by E. B. Norris ~.: (".. ::". .. i ~.~ ........:...;.;.. :\\lil.:.... ~!! ~, .~ *. ~ for Indiana & Michigan Power Company Donald C. Cook Nuclear Plant Bridgman, Michigan 49106 September 16, 1981 Approved: AWkcIkL U. S. Lindholm, Dir"ector Department of Materials Sciences

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( ABSTRACT The first vessel material surveillance capsule removed from the Donald C. Cook Unit No.2 nuclear power plant has been tested, and the results have been evaluated. Heatup and cooldown limit curves for nor-mal operation have been developed for up to 12 and 32 effective full power years of operation. ii

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TABLE OF CONTENTS LIST OF FIGURES LIST OF TABLES I.

SUMMARY

OF RESULTS AND CONCLUSIONS II. BACKGROUND III. DESCRIPTION OF MATERIAL SURVEILLANCE 'PROGRAM IV. TESTING OF SPECIMENS FROM CAPSULE T A. Shipment, Opening, and Inspection of Capsule B. Neutron Dosimetry C. Mechanical Property Tests V. ANALYSIS OF RESULTS VI. HEATUP AND COOLDOWN LIMIT CURVES :fOR NORMAL OPERA-TION OF DONALD C. COOK UNIT NO. 2 VII. REFERENCES APPENDIX A - TENSILE TEST RECORDS AL'.I:'fu'lli)J.X.IS - J:'KU~~lJuR.E' FuR Inc; GEiiE:a.ATION 0;: ;,:::.r.C';;~L::::: PRESSURE-TEMPERATURE L:n1IT CURVES FOR NUCLEAR POWER PLANT REACTOR VESSELS iii Page iv v 1 3 7 13 13 14 22 33 ~~ 41 47 49 63

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r Figure 1 2 3 4 5 6 7 ( 8 9 10 11 12 13 LIST OF FIGURES Arrangement of Surveillance Capsules in the Pressure Vessel Vessel Material Surveillance Specimens Arrangement of Specimens and Dosimeters in C.;1psule T Radiation Response of Donald C. Cook Vessel Shell Plate C552l-2 (Longitudinal Orientation) Radiation Response of Donald C. Cook Vessel Shell Plate C552l-2 (Transverse Orientation) Radiation Response of Donald C. Cook Vessel Core Region Weld Metal Radiation Response of Donald C. Cook Vessel Core Region Weld HAZ Material Effect of Neutron Fluence on RTNDT Shift, Donald C. Cook Unit No. 2 Dependence of Cv Upper Shelf Energy on Neutron Fluence, Donald C. Cook Unit No. 2 Reactor Coolant System Pressure-Temperature Limits Versus 100°F/Hour &ate Criticality Limit and Hydrostatic Test Limit, 12 EFPY Reactor Coolant System Pressure-Temperature Limits Versus Coold~wn Rates, 12 EFPY Reactor Coolant System Pressure-Temperature Limits Versus 100°F/Hour Rate Criticality Limit and Hydrostatic Test Limit, 32 EFPY Reactor Coolant System Pressure-Temperature Limits Versus Cooldown Rates, 32 EFPY iv 8 11 12 27 28 29 30 35 38 43 44 45 46

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( Table I II III IV V VI VII LIST OF TABLES Donald C. Cook Unit No.2 Reactor Vessel Sur-veillance Materials [14J Summary of Reactor Operations, Donald C. Cook Unit No. 2 Results of Discrete Ordinates 3n Transport Analysis, Donald C. Cook Unit No.2, Capsule T Summary of Neutron Dosimetry Results, Donald C. Cook Unit No.2, Capsule T Charpy V-Notch Impact Test Results, D.C. Cook Unit No. 2 Vessel Intermediate Shell Plate C552l-2, Longitudinal Orientation Charpy V-Notch Impact Test Results, D.C. Cook Unit No. 2 Vessel Intermediate Shell Plate C552l-2, Transverse Orientation Charpy V-Notch Impact Test Results, D.C. Cook Unit No. 2 Vessel Core Region Weld Metal VIII Charpy V-Notch Impact Test Results, D.C. Cook Unit No. 2 Vessel Core Region Weld Heat-Affected Zone Material IX x XI XII Effect of Irradiation on CapsuleT. Surveillance Materials, Donald C. Cook Unit No.2 Tensile Properties of Surveillance Materials, Donald C. Cook Unit No. 2 Projected Values of RTNDT for Donald C. Cook Unit No. 2 Reactor Vessel Surveillance Capsule Removal Schedule [15J, Donald C. Cook Unit No.2 v 9 16 19 20 23 24 25 26 31 32 36 40

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SUMMARY

OF RESULTS AND CONCLUSIONS The analysis of the first material surveillance capsule removed from The Donald C. Cook Unit No. 2 reactor pressure vessel led to the following conclusions: (1) Based on a calculated neutron spectral distribution, Capsule T received a fast f1uence of 2.2 x 1018 neutrons/cm2 (E > 1 MeV). (2) The surveillance specimens of the core be1t1ine materials ex-perienc'ed shifts in transition temperature of 40 0 r to 80 0 r as a result of the above exposure. (3) The intermediate shell plate materials exhibited the largest shift in RTNDT and will control the heatup and cooldown limitations throughout the design lifetime of the pressure vessel. (4) The estimated maximum neutron fluence of 6.7 x 1017 neutronsl cm2 (E > 1 MeV) received by the vessel wall accrued in 1.08 effective full power years (EFPY). Therefore, the projected maximum neutron f1uence after 32 EFPY is 2.0 x 1019 neutrons/cm2 (E >.1 HeV). This estimate is based on a lead factor of 3.24 between the center of C~psu1e T and the point of maximum pressure vessel flux. (5) Based on Regulatory Guide 1.99 trend curves, the prOjected maximum shift in ductile-brittle transition temperature of the Donald C. Cook Unit 2 vessel core be1tline plates at the 1/4T and 3/4T positions after 12 EFPY of operation are l68°r and 113°r, respectively. These values were used as the bases for computing heatup and coo1down limit curves for up to 12 EFPY of operation.

(6) The maximum shifts in the transition temperature of the Donald C. Cook Unit 2 vessel core beltline plates at the 1/4T and 3/4T positions after 32 EFPY of operation are projected to be 240°F and 149°F, respectively. These values were used as the bases for computing heatup and cooldown limit curves for up to 32 EFPY of operation. (7) The Donald C. Cook Unit No. 2 vessel plates, weld metal, and HAZ material located in the core beltline region are projected to retain sufficient toughness to meet the current requirements of 10CFRSO Appendix G throughout the design life of the unit. 2 c . (

II. BACKGROUND The allowable loadings on nuclear pressure vessels are determined by applying the rules in Appendix G, "Fracture Toughness Requirements," of 10CFR50 [1]. In the case of pressure-retaining components made of fer-ritic materials, the allowable loadings depend on the reference stress intensity factor (KIR) curve indexed to the reference nil duc-tility temperature (RTNDT) presented in Appendix G, "Protection Against Non-ductile Failure," of Section III of the ASME Code (2]. Further,the materials in the beltline region of the reactor vessel must be monitored for radiation-induced changes in RTNOT per the requirements of Appendix H, "Reactor Vessel Ma terial Surveillance Program Requirements," of 10CF'R.S0 ~ The RTNDT is defined in paragraph NB-233l of Section III of the ASME Code as the highest of the following temperatures: (1) Drop-weight Nil Ductility Temperature (DW-NDT) per AS'l'M E 208 (3J; (2) 00 dE:;~ F ot:iuw temperature; .$0 i't-ll.,,...,.... ---.... .... cu."" V-notch (3) 60 deg F below the 35 mil Cv temperature. The RTNDT must be established for all materials, including weld metal ~~d heat-affected zone (HAZ) material as well as base plates and forgings, which comprise the reactor coolant pressure boundary. It is well established that fer-ritic materials undergo an increase in strength and hardness and a decrease in ductility and toughness when exposed to neutron fluences in excess of 1017 neutrons per cm2 (E > 1 MeV) [.4]. Also, it has been established that tramp elements, particularly 3

copper and phosphorus, affect the radiation embrittlement response of ferritic materials (5-61. The relationship between increase in RTNDT and copper content is defined in Regulatory Guide 1.99. Although this docu-ment is being revised by the NRC to reflect a more recent evaluation of neutron embrittlement data by the Metal Properties Council [7], estimates of shifts in RTNDT in this report are based on the current Revision 1 of Regulatory Guide 1.99 [8]. In general, the only ferritic pressure boundary materials in a nuclear plant which are expected to receive a fluence sufficient to affect RTNDT are those materials which are located in the core beltline region of the reactor pressure vessel. Therefore, material surveillance programs include specimens machined from the plate or forging material and weldments which are located in the core beltline region of high neutron flux density. ASTM E 185 (9J describes the recommended practice for monitoring and evaluating the radiation-induced changes occurring in the mechanical properties of pressure vessel beltline materials. Westinghouse has provided such a surveillance program for the Donald C'. Cook Unit No. 2 nuclear power plant. The encapsulated Cv specimens are located on the 0.0. surface of the thermal shield where the fast neutron flux density is about three times that at the adjacent vessel wall surface. Therefore, the increases (shifts) in transitiqn temperatures of the materials in the pressure vessel are generally less than the corresponding shifts observed in the surveillance specimens. However, because of azimuthal variations in neutron flux density, cap-sule fluences may lead or lag the maximum vessel fluence in a correspond-ing exposure period. The capsules also contain several dosimeter materials 4 ( ( (

( for experimentally determining the average neutron flux density at each capsule location during the exposure period. The Donald C. Cook Unit No. 2 material surveillance capsules also include tensile specimens as recommended by ASTM E 185. At the present t~e, irradiated tensile properties are used only to indicate that the materials tested continue to meet the requirements of the appropria-e material specification. In addition, the material surveillance capsules contain wedge opening loading (WOL) fracture mechanics specimens. Cur-rent technology limits the testing of these specimens at temperatures well below the minimum service temperatur~ to obtain valid fracture mechanics data per ASTM E 399 [lOJ, IIStandard Method of Test for Plane-Strain Fracture Toughness of Metallic Materials." However, recent work reported by ~~ger and Witt [llJ may lead to methods for evaluating high-toughness materials with small fracture mechanics specimens. Currently, the NRC suggests storing these specimens until an acceptable testing procedure has been defined. This report describes the results obtained from testing the contents of Capsule T. These data are analyzed to estimate the radiation-induced changes in the mechanical properties of the pressure vessel at the time of the refuelling outage as well as predicting the changes expected to occur at selected times in the future operation of the Donald C. Cook Unit No. 2 power plant. 5

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( III. DESCRIPTION OF MATERIAL SURVEILLANCE PROGRAM The Donald C. Cook Unit No. 2 material sur.veillance program is described in detail in WCAP 8512 [12], dated November 1975. Eight mate-rials surveillance capsules were placed in the reactor vessel between the thermal shield and the vessel wall prior to startup~ see Figure 1. The vertical center of each capsule is opposite the vertical center of the core. The Capsule T neutron flux density was initially reported to lead the maximum flux density on the vessel I.D. by a factor of 2.9 [12]. How-ever,.in a letter to the American Electric Power Service Corporation [13], Westinghouse reported that, the 40° capsule lead factor had been changed to 3.7 as a result of refined calculational methods. The capsules each contain Charpy V-notch, tensile, and WOL specimens (c machined from the SA533 Gr B, Cl 1 plate, weld metal, and heat-affected zone (HAl) materials located at the core beltline. The chemistries and "h"'~f-r,.."''tIl'pnts of rne vesRel surveillance materials are summarized in Table I. All test specimens were machined from the test materials at the quarter-thickness' (1/4 T) location after performing a simulated postweld stress-relieving treatment. Weld and HAl specimens were machined from a stress-relieved weldment which joined sections of the intermediate and lower shell plates. HAl specimens were obtained from the plate C55Zl-Z side of the weldment. The longitudinal base metal Cv specimens were ori-ented with their long axis parallel to the primary rolling direction and with V-notches perpendicular to the major plate surfaces. The transverse base metal Cv specimens were oriented with their long axis perpendicular to the primary rolling direction and with V-notches perpendicular to the i

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AR...-q,ANGEMENT OF S1.j*RVEILu..~c:: CAPSULES !~ 1'EE PRESSURE VESSEL 3

TABLE I DONALD C. COOK UNIT NO. 2 REACTOR VESSEL SURVEILLANCE MATERIAl,S (14 J Hea': Treatment History Shell Plate Material: Heated to 1700 +/- 50°F for 4-1/2" hours, water quenched; Heated to 1600 +/- 50°F for 5 lloucs t water quenched; Tempered at 1250 +/- 50°F for 4-1/2 hours, air cooled; Stress relieved at 1150 +/- 25°F for 51-1/2 hours, furnace cooled. Weldment: Stress relieved at 1140 +/- 25°F for 9 hours, furnace cooled. Chemical Composition (Percent) \\ Material C Mn Sf Ni***. Mo Plate C552l-2 (a) j 0.21

1. 29 0.16

/0.58 0.50 I Plate C552l-2 (b) I 0.220 1.280 0.270 i 0.580 0.550. Metal (e) \\ Weld 0.08 1.42 0.36 ! 0.96 Metal (b) 0.110 0.440 / 0.540 Weld 1.330 \\0.970 / I (a) Lukeps Steel analysis (b) Westinghouse analysis (c) Chicago Br!~ge an~ Iron analysts Cu 0.14 0.110 0.05 0.055 Cr 0.072 0.07 0.068

major plate surfaces. Tensile specimens were machined with the longitudi-nal axis perpendicular to the plate primary rolling direction. The WOL specimens were machined with the simulated crack parallel to the primary rolling direction and perpendicular to the major plate surfaces. All mechanical test specimens, see Figure 2, were t~en at least one plate thickness ~rom the quenched edges of the plate material. Capsule T contained 44 Charpy V-notch specimens (8 longitudinal and 12 transverse from the plate material, plus 12 each from weld metal and HAZ material); 4 tensile specimens (2 plate and 2 weld metal); and 4 WOL specimens (either plate or weld metal). The specimen numbering system and location within Capsule T is shown in Figure 3. Capsule T also was reported to contain the following dosimeters for determining the neutron flux density: .( Target Element Form Quantity Iron Bare wire 5 Copper Bare wire 3 Nickel Bare wire 3 Cobalt (in aluminum) Bare wire 2 Cobalt (in aluminum) Cd shielded wire 2 Uranium-238 Cd shielded oxide 1 Neptunium-237 Cd shielded oxide I Two eutectic alloy thermal monitors had been inserted in holes in the steel spacers in Capsule T. One (located at the bottom) was 2.5% Ag and 97.5% Pb with a melting point of 579°F. The other (located at the top of the capsuie) was 1.75% Ag, 0.75% Sn, and 97.5% Pb having a melting ( 10

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( ( ( IV. TESTING OF SPECIMENS FROM CAPSULE T The capsule shipment, capsule opening, specimen testing, and ~e-porting of results were carried out in accordance with the Project Plan for Donald C. Cook Unit No. 2 Reactor Vessel Irradiation Surveillance Program. The SwRI Nuclear Projects Operating Procedures called out in this plan include: (1) XIII-MS-104-0, "Shipment of Westinghouse PWR Vessel Material Surveillance Capsule Using SwRI Cask and Equipment" (2) XI-MS-I01-0, "Determination of Specific Activity and Analysis of Radiation Detector Specimens" (3) XI-MS-I03-0, "Conducti?-g Tension Tests on Metallic Specimens" (4) XI-MS-104-0, "Charpy Impact Tests on Metallic Specimens" (5) XIII-MS-103-0, "Opening Radiation Surveillance Cap-sules and Handling and Storing Specimens" (6) XI-MS-5-0, "Conducting W'edge-Opening-Loading Tests un !~:a!li~ ~::.=!~l=" Copies of the above documents are on file at SwRI. A. Shipment, Opening, and Inspection of Capsule Southwest Research Institute utilized Procedure XIII-MS-104-0 for the shipment of Capsule T to the SwRI laboratories. SwRI personnel sev-ered the capsule from its extension tube, sectioned the extension tube into several lengths, supervised the loading of the capsule and extension tube mate~ials into the shipping cask, and transported the cask to San Antonio, Texas. 13

The capsule was opened and the conte~ts.identified and stored in Irdance with Procedure XIII-MS-103-0. The long seam welds were milled in a hot cell using a Bridgeport vertical milling machine. Before .ing the long seam weld beads, transverse saw cuts were made to remove capsule ends. After the long seam welds had been milled off, the top of the capsule shell was removed. The specimens and spacer blocks carefully removed and placed in indexed receptacles~ identifying each iule location. After the disassembly had been completed, each specimen carefully checked to insure agreement with the identification and ,tion as listed in WCAP 8512 [12]. No discrepancies were found. The thermal monitors and neutron dosimeter wires were removed from holes in the spacers. The thermal monitors, contained in quartz .s, were examined. No evidence of melting was observed~ thus indicating

the maximum temperature during exposure of Capsule T did not exceed F.

The top and bottom Cd-covered Co-Al neutron monitors were not in capsule, but all other neutron dosimeters were correctly accounted for. lack of Cd-covered Co-Al dosimeters had no significant impact on the 'eillance program because they were intended for the determination of thermal neutron flux and thermal flux burnup corrections are not needed 'WR programs. Neutron Dosimetry The gamma activities of the dosimeters were determined in accordance. Procedure XI-MS-I01-O using an 1T-5400 multichannel analyzer and a

Li) co~~ial detector system.

The calibration of the equipment was )mplished with 54Mn, 60Co, and 137Cs radioactivity standards obtained 1 the U.S. Department of Commerce ~ational Bureau of Stand~rds. All ~vities were corrected to the time-of removal (TOR) at reactor shutdown. 14 ( (

( ( ( The dosimeter wires were weighed ona Mettler microbalance, and the fission monitors were weighed on a Mettler digital balance after these materials had been deencapsulated. Infinitely dilute saturated activ-ities (ASAT) were calculated,for each of the dosimeters because ASAT is directly related to the product of the energy-dependent microscopic acti-vation cross section and the neU1ron flux density. The relationship be-tween ATOR and ASAT is given by: where: ASAT m-n r mal decay constant for the activation product, day-I; equivalent operating days at 3391 MwTh for op-erating period m; and decay time after operating period m, days

• An alternate expression which gives equivalent results is:

where: O'f)erating days; and average fraction of full power ~~~ing period. The Donald C. Cook Unit No. 2 operating history up to the 1979 refueling shutdown, which was used in the calculation of ATOa, is presented in Table II. The primary result desired from the dosimeter analysis is the total fast neutron fluence (> 1 MeV) which the surveillance specimens received. The average flux density at full power is given by: 15

I-' 0\\ Operating Dates Period Start Stop 1 03/22/18 04/19/18 04/20/18 05/02/78 2 05/03/18 05/19/18 OS/20/18 05/31/78 3 06/01/78 06/03/78 06/04/78 06/04/78 4 06/05/78 07/21/78 07/22/78 08/04/78 5 08/05/78 08/21/78 08/22/78 08/22/78 6 08/23/78 08/28/78 08/29/78 09/01/78 7 09/02/18 11/09/78 11/10/78 11/23/18 8 11/24/18 04/07/79 04/08/19 04/08/19 9 04/09/19' 05/19/79 OS/20/79 07/02/79 10 07/03/79 10/19/19 I TABLE II

SUMMARY

OF REACTOR OPERATIONS DONALD C. COOK UNIT NO. 2 Power Operat~ng Shutdown Generation Days: Days (MWDd 33, 23,969 ~7 J 14 27,264 1 1.2 3 I 5,510 1 47 : 103,423 14 17 i 51,223 1 6 12,543 4 69 206,496 14 135. 423,785 1 l.1 ~ 132,362 44 109 356 1918 TO,tal Cycle 1 - 1,343,493 (a) One equivalent operating day"" 3391 MHt (b) Equals 1.08 EFPY Equivalent Decay Time, Operating Days After Period (Tm) (a) (tm) 7.07 549 8.04 518 1.62 503 30.50 455 15.11 424 3.70 417 60.90 344 124.97 195 ' 39.03 153 105.25 0 396.l9(b)

( where: ASAT.. NO energy-dependent,neutron flux density, n/cm2-sec; saturated activity, dps/mg target element; spectrum-averaged activation cross section, cm2; and number of target atoms per mg

• The total neutron fluence is then equal to the product of the average neu-tron flux density and the equivalent reactor operating time at full power.

In Capsule T, all 12 weld metal and all but ewo of the 12 HAZ Charpy specimens were located in the specimen layer nearest to the vessel wall, and the vessel plate (longitudinal and transverse) Charpy specimens-were located in the specimen layer nearest to the core. Since there is a radial ( dependence of the fast neutron flux in the vessel, the neutron exposure re-ceived by the weld metal and HAZ Charpy specimens is expected to be lower program was planned to provide information on the raCial dependence of the fast flux since the copper and nickel threshold detectors were located on the radial centerlines of the Charpy specimen layers nearest to and farthest from the core, respectively, and the iron threshold detectors were located at the radial position corresponding to the interface be-tween the two Charpy specimen layers. Additional dosimetry included the fission monitors located at the radial centerline of the capsule and the bare cadium-shielded cobalt-aluminum monitors locate~ at the radial center-line of the Charpy specimen layer nearest the pressu=e vessel wall. l 17

A discrete ordinates Sn transport analysis for the Donald C. Cook Unit No. 2 reactor vessel was performed to determine the axial, radial, and azimuthal dependence of the fast neutron (E > 1.0 MeV) flux density and energy spectrum within the reactor vessel and surveillance capsules. These results were used to calculate the spectr~averaged cross-sections for the threshold detectors and the lead factors for use in relating neu-tron exposure of the pressure vessel to that of the surveillance capsule. The pertinent factors obtained from this transport analysis are summarized in Table III. The 3.24 center-of-capsule lead factor agrees well with the revised 3.7 average lead factor re~orted by Westinghouse [13]. The Capsule T dosimetry results are presented in Table IV. A sum-mary of the capsule and vessel I.D. fluxes calculated for full-power op-eration is as follows: Dosimeter Measured Capsule Flux Lead Peak Vessel Flux at I.D. Type cm-2.sec-l, E > 1 MeV Factor cm-2'sec-l, E > 1 MeV Copper 5.64 x 1010 3.38 1.67 x 1010 ~illj D _iii. Iron 5'.56 x 1010* 3.02 1.84 x 1010 Nickel 5.25 x 1010 2.71 1.94 x 1010

• If a fission-spectrum energy distribution is assumed at the capsule location, the cross"'section for the 54Fe(n,p)54Mn, raaction (E > l.0 MeV) would be 98.26 mb [4], and the resulting value for fast flux at the capsule location would be 4.88 x 1010 cm-2*sec-l
• This value is reported for reference only and has not been used in the analysis of results.

The discrepancies in the peak vessel flux values determined from the sev-eral. dosimeter materials are attributed primarily to the uncertainties in ( ( (

( ( TABLE III RESULTS OF DISCRETE ORDINATES Sn TRANSPORT ANALYSIS DONALD C. COOK UNIT NO. 2 CAPSULE T A. Calculated Reaction Cross-Sections for Analysis of Fast Neutron Monitors (E > 1.0 MeV) Reaction 54Fe(n,p)54Mn, 58Ni(n,p)58Co 63Cu(n,a)60Co 238U(n,f) 237Np(n,f) C1 (barns) .0863 .115 .00093 .385 2.46 B. Calculated Capsule Lead Factors Position(a) Location within Capsule 211.7 em Center of core-side Charpy layer Center of capsule 212.2 em Center of two specimen layers 212.7 cm Center of vessel-side Charpy layer (a) Distance from center of core Lead Capsule neutron flux density, E > 1.0 MeV Factor (b) 3.38 3.02 2.71 (b) Maximum neutron flux density at vessel I.D., E > 1.0 MeV 19

N 0 TABLE IV

SUMMARY

OF NEUTRON DOSIMETRY RESULTS DONALD C. COOK UNIT NO.2, CAPSULE T Dosimeter Dosimeter Activation ATOR Position(a) Identification Reaction (dps/mg) 211. 7 em Cu (Top Middle) 63Cu(n,a)60Co 4.53 x 101 ~ Cu (Middle) 4.43 x 101 Cu (Bottom Middle) 4.51 x 101 211.9 em U-238 (Middle) r238U(n, f)137Cs 1.06 x 102 t Np-237 (Middle) 237Np(n,f)137Cs 8.22 x 102 54Fe (n, p) 54.10 103 212.2 em Fe (Top) 1.59 x I Fe (Top Middle) I 1.62 x 103 Fe (Middle) 1.57 x 103 }<'e (Bottom Middle) 1.'63 x 103 Fe (Bottom) 1.60 x 103 212.7 em Ni (Top Middle) 58Ni(n,p)58Co 3.59 x 104 Ni (Middle) l 3.50 x 104. Ni (Bottom Middle) 3.57 x 104 212.7 em Co (Top) ,)9Co(n~y)60Co 4.96 x 106 t Co (Bottom) t 6.31 x 106 (a) Distance from center of core (b) Calculated flux values subject to :!. 16.5% uncertainty (10) (e) Not applicable ~ /-.,. Capsule Flux, ASAT ~, E > 1.0 MeV(b) (dps/mg) cm-2*sec-l 3.47 x. 102 5.69 x 1010 3.39 x 102 5.56 x 1010 3.46 x 102 5.67 x 1010 Average 5.64 x 1010 4.30 x 103 7.36 x 1010 3.35 x 104 8.24 x 1010 Average.. 7.80 x 1010 2.97 x 103 5.50 x 1010 3.03 x 103 5.61 x 1010 2.94 x 103 5.45 x 1010 3.06 x 103 5.66 x 1010 3.00 x 103 5.56 x 1010 Average '" 5.56 x 1010 4.27 x 104 5.30 x 1010 4.16 x 104 5.17 x 1010 4.24 x 104 5.27 x 1010 Average 5.25 x 1010 3.80 x 107 (c) 4.83 x 107 (c)

( \\ the calculated spectra and in the reaction cross sections. Other neutronic factors contributing to the estimated +/- 16.5% uncertainty (10) in a ca1cu-1ated flux value are the determination of disintegration rates and the cal-culation of reaction rates (ASAX/NO). Averaging the results obtained from all Capsule X neutron dosimeters, the peak neutron flux incident of the I.D. surface of the pressure vessel during fuel cycle 1 is calculated to be 1.96 x 1010 em-2'sec-1, E > 1 MeV. The calculated full power neutron flux for the weld metal and HAZ Charpy specimen layer is given by: 1.96 x 1010 x 2.11 = 5.3 x 1010 Similarly, the calculated full power neutron flux for the vessel plate Charpy specimens, the tensile specimens, and the WOL specimens are given by: 1.96 x 1010 x 3.38 - 6.6 x 1010 (Plate Cv Specimens) 1.96 x 1010 x 3.02 - 5.9 x 1010 (Tensile Specimens) 1.96 x 1010 x 3.24 = 6.4 x 1010 (WOL Specimens) Sinc~ Donald C. Cook Unit No. 2 operated for 396.19 effective full power days up to the October 1979 refueling, the calculated capsule and vessel fluences to that time are as follows: Weld Metal and HAZ Charpy Specimens - l.8x 1018 n/cm2 Vessel Plate Charpy Specimens - 2.3 x 1018 n/cm2 Tensile Specimens - 2.0 x 1018 n/cm2 i-lOL Specimens - 2.2 x 1018 n/cm2 Pressure Vessel ID Surface - 6.7 x 1017 n/cm2 21

C. Mechanical Property Tests The irradiated Charpy V-notch specimens were tested on a SATEC im-pact machine in accordance with Procedure XI-MS-104-0. The test tempera-tures were selected to develop the ductile-brittle transition and upper shelf regions. The unirradiated Charpy V-notch impact data reported by Westinghouse [12J and the data obtained by SwRI on the specimens contained in Capsule T are presented in Tables V through VIII. The Charpy V-notch transition curves for the two plate materials, the weld metal, and the HAZ material are presented in Figures 4 through 7. The radiation-induced shift in transition temperatures are indicated at the 50 ft-lb, 30 ft-lb, and 35 mil lateral expansion levels. A summary of the shifts in RTNDT and Cv upper shelf energies for each material are presented in Table IX. Tensile tests were carried out in accordance with Procedure XI-MS-103-0 using a 50-kip capacity tester equipped with a strain gage exten-someter, load cell, and autographic recording equipment. Tensile tests on the plate material were run at 250°F and 550°F; those on4 the weld metal were run at 210°F and 550°F. The results, along with tensile data reported by Westinghouse on the unirradiated materials [12], are presented in Table X. The load-strain records are included in Appendix A. Testing of the WOL specimens was deferred at the request of Indiana & Michigan Electric Company. The specimens are in storage at the SwRI radi-ation laboratory. 22

( ( TABLE V CHARPY V-NOTCH IMPACT TEST RESULTS, D.C. COOK UNIT NO. 2 VESSEL INT~~IATE SHELL PLATE C5521-2, LONGITUDINAL ORIENTATION Specimen Condition No. Capsule T ML-40 ML-33 ML-39 ML-35 ML-36 ML-34 ML~37 ML-38 (al (b) (a) Unirradiated (12] (b) Not reported Test Temperature (deg F) 50 75 100 120 165 210 250 300 0 0 0 25 25 25 50 50 50 70 70 70 100 100 100 125 1.25 125 210 210 210 23 Fracture Lateral Fracture Energy Expansion Appearance (ft-lb) (mils) (% shear) 17 15 10 25 24 10 38.5 30 20 72.5 58 25 92 65 80 110 85 100 110 88 100 112 89 100 15 12 18 18 13 19 15 11 18 26 20 30 31 23 25 43 29 25 52 38 35 47 37 35 46 34, 35 65 47 "42 65 49 42 76 54 55 91 66 65 98 76 70 90 67 62 126 78 85 114 79 77 103 70 75 122 83 100 132 86 100 128 84 100

TABLE VI CHARPY V-NOTCH IMPACT TEST RtSULTS, D.C. COOK UNIT NO. 2 VESSEL INTERMEDIATE SHELL PLATE C5521-2, TRANSVERSE ORIENTATION Specimen Condition No. Capsule T MT-60 MT-49 MT-56 MT-57 MT-51 MT-58 MT-52 MT-59 MT-50 MT-53 MT-54 MT-55 (a) (b) I I I I, I i (a) Unirradiated [12J (b) Not reported Test Temperature (deg F) 20 75 75 100 120 120 165 165 210 250 300 300 -50 -50 -50 10 10 10 70 70 70 100 100 100 120 120 120 210 210 210 24 Fracture Lateral Fracture Energy Expansion Appearance (ft-lb) (mils) (% shear) 10 9 Nil 20.5 18 5 17 18 10 38.5 35 10 37 32 20 32 31 20 54 49 30 47 46 30 66.5 66 80 76 71 100 75.5 70 100 71 68 100 5.5 0 5 6 1 5 6 0 5 39 27 29 29 17

• 25 25 18 30 43 32 40 42 33 43 39 28 43 66 47 60 71.5 53 63 68 49 65 67.5 56 58 76 60 65 75 59 72 81 64 100 88 63 100 90 66 100

(

( ( ( TABLE VII CHARPY V-NOTCH IMPACT TEST RESULTS, D.C. COOK UNIT NO. 2 VESSEL CORE REGION WElD METAL Specimen Condition No. Capsule T MW-59 I MW-58 MW-57 I MW-60 MW-49 MW-56 MW-51 MW-52 MW-50 ~ MW-53 MW-54 MW-55 (a) (b) I I I I I ~ (a) (b) Unirradiated [12J Not reported Test Temperature (deg F) -25 20 50 60 75 75 120 165 210 250 300 300 -25 -25 -25 20 20 20 60 60 60 100 100 100 210 210. 210 300 300 300 25 Fracture La.teral Fracture Energy Expansion Appearance (ft-lb) (mils) (% shear) 12 11 5 21.5 20 15 23 21 10 35.5 32 30 58.5 48 75 46.5 44 9S 49 44 20 77 69 100 75 75 100 68 55 100 74 72 100 7S 63 100 20 20 48 22 17 30 31.5 25 40 32 24 38 35 28 47 33 27 50 58 48 74 47 37 65 39 29 50 74 63 95 56 47 85 65 53 95 72 68 100 70 63 100 77 64 100 72 66 98 79 71 100 81 70 100

TABLE VIII CHARPY V-NOTCH IMPACT TEST RESULTS, D.C. COOK UNIT NO. 2 VESSEL CORE REGION WELD HEAT-AFFECTED ZONE MATERIAL Specimen Condition No. Capsule T MH-59 I MH-58 I MH-60 MH-57 MH-56 MH-49 MH-51 MH-52 MH-50 MH-53 MH-54 MH-55 (a) (b) (a) Unirradiated [12] (b) Not reported. Test Temperature (deg F) -50 -25 0 20 50 75 120 165 210 250 300 300 -100 -100 -100 -50 -50 -SO -25 -25 -25 0 0 0 50 SO 50 125 125 125 210 210 210 26 Fracture Lateral Fracture Energy Expansion Appearance (ft-1b) (mils) (% shear) 10 11 20 22 20 30 38.5 32 50 43 32 60 97 68 90 92.5 60 90 125 80 95 104 66 100 62.5 56 100 110.5 69 100 112 72 100 110.5 70 100 21 12 30 5 1 12 14 8 29 34 16 35 23 21 27 70.5 39 53 89 52 65 70 43 60 90 52 60 95 59 70 76 52 65 130 75 100 84 55 90 67 48 85 136 76 100 95 66 95 104 75 99 82 71 90 147 77 100 113 80 100 86 71 100 ( ( (

160 120 80 40 o I I I I I I 1 I 1 I... .L

• 1 I...

i :/ i ! I f 1'- I I,I! f I 'A 1, 1-1 I I I I -..J I 1, t ! I I 'I 1 ' I I I I I I I I .1*1 I I , 1 I I t i I, I I I ,I 1 I HI-!,+-H-,f-' -i-: -+i"';;!-;!.,<fC.Y~+-+I-: M'"=",,,,: ATI +-r-+-H....... +-+++-+-H~: Unirrad~ated, WCAP 8512 ~~-t ~I-+-, +I-l-~'-l.., ~...;.; --H,,""+-+-H-+-+-i'-+-t-~-+-i-' -:-'-+-: +, -+-HH Caps ul e T, Uni t 2 I I I I I I I I I I I i ! 1.1 ~ I I I I -100 o 100 200 300 400 Cv Test Temperature, deg F 100 I , I I, I I I I I f I i i , aI I.JooO'T I i I ,I I ! i I I, I ! I i I I 75 I r j I i I t*1 ,, I A', 'Y .if I I ! ! l I i j 1, t I I [ I i I A,_ I "! I I I ! I I 1 ' i 1 Ii Ii , I 11 I I I I I I I I I I I I , I I I 50 1 I i i i I I I I, !I! I j I ( f ! i j! ii, I I I I Ii I I I j I I i , I i I I I t j _~i 1 I! i I

j.

i i I I J I II. I I 1 I ,VI J i J I I I I I I I 25 J I ! i j i I I ! III I 1 1/. ~ I i It i I I _I 1 ', i I 1 I ! /1"', I '/ I I , i If,....; , ! I I I i III 114" Ii rIA-.' i ~....j,.., -+1-, "";-'-j +-........A-&-i-'-",-'-, "':',-'-, "",-+-, -i-f-"';'-'~-:-! "',~, +-,H Unirradiated, WCAP 8?12 i----t j....;.---+-1-1-_..... '-'.;....' +-1'-+1...,.'-,--i-;-:"'..J1-;-+--.,-'... '-:--'-H-, '-;,-+-, ~ Capsule T, Unit 2 i I*', I I ! I I I o -100 I I I o I 100 200 300 400 FIGTJRE 4. Cv Test Tempera.ture, deg F RADIATION RESPONSE OF DONALD C. COOK VESSEL SHELL PLATE C5521-2 (LONGITUDINAL ORIENTATION) 27

h f

y ~j f '"'/ \\ '. \\ \\

~ I.., 00 ~ ~ c: J;:J u 100 1 ' 75 I ~_ L I , I 50 I, 1 [I'" I

i.

I 25 ! i I I I I

i.

H-++H++-i74-+-i-++H7f-+-if-++H+-H-++-H++-'...i.'-l.....L LL I I I I I I I I i i,--.'....... 0..+ rt~~~~~~~~~~++~~~++~~Unirradiated, WCAP 8512 rt~~~++~~r+~~~++~~~++,~:~~Capsule T, Unit 2 I, ILl I 1 If, I 1 I 1 1 I. I a t:Cj:i:tj:t:ti:tjc!:tj:!:tJ:t:tj:t:tj:t,j':i:tj:~tj:tj':+/-::':l:t::+/-:t'J: 3tj:+/-:ti:i,:t, j:~.jlt:::j:+/-:t -100 a 75 I I f I f I I f I-I

I 50 f

I I 1 25 1 ' 100 200 300 400 Cv Test Temperature, deg F , i I, I f ! 1 I ia-; I , I I I I I I , I 1 ! i I ! I f I I , I 1 II J I I d 1 I I I 1""" I ! V I "":II/!'I .Il I I I I f 'L fA!); i I' .II 'I II! I! at Jl I I I I f~ I I I i I , I j I I 1 I I - A ..... ~... .1 I I I '/1 I I ./1 I,, 1 I i ~ I I i ! !..If I 'I! I l , i ~ I I 'A'l"i i' I' I I" r,! t I I i i 1 i I ii' I I* f I 'I I i I I 1 f I I i I 1 i I I I I, I i I j I ! 1 ! I I j i f i H-+++-l-+..I...';-O-70~t;--!-+~-:-' '+-+-:-+++-+'-+1 +-r-+...... ~~' 1 i I I I : i I I ! I i I ~1...,1i-+~-+--!--;f_~7'9-+i-+~.".l-t", -t-'ii-+-H-+-ri-+-;"-:--+-~ Unirr adia ted, t-lCAP 8512 :--"-~4 I!.A' ~ f 1 I , I I I, ~-+, ...L'-+-+I-i-..A'-r!;-+-+-='~,-+, -+;--,..;.; -,~, -+-+, -;-i-~-i-""':"-; Cap sule T, Uni t 2' a '., -100 FIGURE j..;f I I 1 I I I I t ,Q I I a 100 200 300 Cv Test Temperature, deg'F RADIATION 'RESPONSE OF DONALD C. COOK VESSEL SHELL ~TE C5521-2 (~~SVERSE ORIENTATION) 28 , I 1 400 ( (

,.Q r-i I..., ~ oc Ioi Q) C ~ U

rtl I

I I r-I I I I 75 t,.. ~....... : I I IJ ! I I I , I I.... 1, f7 i I ..... ai if , i l."j ~tT" 50 I I' I I. I, I 1/ !A I 1 Y I I I IV, I I I ~ !~ ... 1:1 I, I I i, i'" Go' i I 1 I 'I 25 V ~ i 1 I I a -100 100 73 I 50 i 25 I"y ,;"r rw I , I I I I I I I , I I 1 I I I I 1'"' I I ,...., I iii i I I I a -100 J t j i 1 , \\ FIGURE 6. a a iA.... I I I i I I I I ! 1 I I I I I I ,I I I ! I I, ; Unir-radiated, WCAP 8512, ! ! I I , I Capsule T, Unit 2 I I I I 1 I, I 1, I, I 1 I 1 1 I I i,,, I I I, I, , i,, IT 100 200 300 400 Cv Test Temperature, deg F I Ii* I I I I I I I .Y Y, : I I I/. i7 V ...,(f I j}i if I If' I I ,11 iI' A I i i

.*.AA, til i

I I I t i' I 1 I I I 100 I I I \\ i I r i t I ( I \\, i I I, 1 I I ! 1 ""!" I I i I I I ! I I I iii ..J.: I I i I: ...,~_. I ~ i ':-'1iii~ '1'1 .-y, r ! If! I i .f'" I I ' I 1

II

'I I i I I I I I i f I I I I I I I I I I I 200 I I .... I I I' I, I,,, I i , I , I I ! I I I I i I I , i I I ! I 1 i \\ 'I I ! i I i I I , i I Unirradiated, \\.]CAP 8512~.....-4 Capsule T, Unit 2 300 400 Cv Test Temperature. deg F RADIATION RESPONSE OF DONALD C. COOK VESSEL CORE REGION I-lELD METAL 29

120 80 40 I I I I I J I I I 1"\\ I I I I I I I I I C / Y I j I I / I I I I '.J, I i, I I I I , 1 j I I I I I I. I '/ I I I I / , I I '!:I~ 'I"" I ' ft' I ; i I q;J./~' ~/_':"'---~-;'.!II!--+-..;,....--~---+--..;,....-..,....,.--e-Unin:adiated, WCAP 8512 _-.....:...-t ,/ Capsule T, Unit 2 , I' I a -100 a 100 200 300 400 Cv Test Temperature, deg F 100 T, --------~--------~--------~--------~~--~--~----~ 75 50 25 a i I I,

I I

I I I I I I I, I, ~ I I ~ , I ~-,~ .nr - I' I L r '.J. ' I I I 11 I I a / $-.~~/-.;LL---+----------+-----~-~Unin:adiated, WCAP 8512----~ ~v~~/~----~------~----+-----------~Capsule T, Unit 2 -100 a FIGURE T. 100 200 300 Cv Test Temperature, deg F RADIATION RESPONSE OF DONALD C. COOK VESSEL CORE REGION WELD HAZ MATERIAL ~ 30 I 400 (

TABLE IX EFFECT OF IRRADIATION ON CAPSULE T SURVEILLANCE MATERIALS DONALD C. COOK UNIT NO. -2 \\... r - '. ') /1 t I ~_. Criterion (1) Transition Temperatur~ Shift @50 ft-lb @ 30 ft-lb @ 35 mil lIRTNDT(4) Cv Upper Shelf Drop 1"',1 Weld Metal(2) p.;>.q. 300~! 40°F 25°~ 40°/ 2 ft-Lb (3%) t'), \\ \\ (1) Refer to Figures 4-7. (2) Fluence g 1.8 x 1018 n/em2, E > 1 ~eV (3) Fluence g 2.3 x 1018 n/em2, E > 1 ~ev HAZ Material(2) p.30 50°F SO°F 50°F Trans. Plate C5521-2 (3) p.).B 90°F 80°F 60°F 50°F ':;:'.C' ( p,,(;, 80°F 22 ft-:-lb .... 12 ft-lb (18%)

-;j) h 'j )

(1lI%) ,F".-" I \\

' I U

\\ \\ () /\\. ': (I.) Maximum transHion temper'ature shift at 30 ft-lb or 35 mil Long. Plate C5521-2(3) f..;l. 7 55°F 55°F 50°F 55°F 16 ft-lb (12%) I {.., Cl. \\ 1 "-\\' I. I

Test Condition Material Capsule T Plate C5521-2 (Transverse) i Weld Metal w t N (a) Plate C5521-2 (Transverse) I Weld Metal I (a) Unirradiated (12) TABLE X TENSILE PROPERTIES OF SURVEILLANCE MATERIALS DONALD C. COOK UNIT NO. 2 Fracture Fracture Spec. Temp. 0.2% YS UTS Load Stress No. (OF) (kai) (ksi) (lb) (kst) MT-lO 250 58.1 89.9 3100 139.6 MT-9 550 66.5 88.3 3150 141.9 MW-10 210 11.0 93.7 2900 113.6 MW-9 550 67.7 90.5 3180 179.7 Room 67.4 81.3 3200 161. 2 Room 65.4 85.9 2950 156.4 300 58.8 78.6 2650 146.1 300 60.5 79.5 2675 157.6 550 57.5 83.0 3225 142.1 550 58.9 83.1 3150 145.6 Room 75.7 93.2, 2850 173.4

Room, 76.9.

91.3. 2950 178.8 300. 70.7. 88.0. 2900 111.0 300' 11.0 ' 85.3. 2875 119.0 550* 70.0. 81.2. 3160 157.2 550* 68.2. 87.8. 3050 166.0 Uniform Total Elongation Elongation R.A. (%). (%) (%) 9.8 19.1 54.4 9.0 18.0 54.4 9.0 21.4 66.0 8.4 19.1 64.0 13.4 23.4 59.6 15.0 27.1 61.1 13.0 22.6 63.1 . dup"- 10.6 19.8 65.4 11.5 19.0 53.8 12.7 20.5 56.0 13.9 25.7 66.8 12.2 ' 22.6* 66.6 10.1. 20.1. 66.0 10.3' 21.2. 61.5. 10.1 19.2. 59.6. 9.3. 20.2. 62.8

V. ANALYSIS OF RESULTS The analysis of data obtained from surveillance program specimens has the following goals: (1) Estimate the period of time over which the properties of the vessel beltline materials will meet the fracture toughness requirements of Appendix G of 10CFRSO. This requires a' projection of the measured reduc-tion in Cv upper shelf energy to the vessel wall using knowledge of the energy and.spatial distribution of the neutron flux and the dependence of Cv upper shelf energy on the neutron fluence. (2) Develop heatup and cooldown curves to describe the operational limitations for selected periods of time. This requires a projection of the measured shift in RTNDT to the vessel wall using knowledge of the dependence of the shift in RTNDT on the neutron fluence and the energy and spatial dis-tribution of the neutron flux. The energy and spatial distribution of the neutron flux for Donald C. Cook Unit No. 2 was calculated for Capsule T with a discrete ordinates transport code~ This analysis predicted that the lead factor (ratio of fast flux at the capsule location to the maximum pressure vessel flux) for Capsule Twas 3.24 at the capsule centerline, 3.38 for the core-side Charpy layer, and 2.71 for the vessel-side Charpy layer (see Table III). This analysis also pr~dicted that the fast flux at the 1/4T and 3/41' positions in the 8.5-in. pressure vessel wall would be 54% and 10%, respectively. of that at the vessel I.D. However, in this report the projection of Cap-sule T results to the pressure vessel wall utilizes the more conservative attenuation figures of 60% and 15% for the 1/4T and 3/4T positions to allow 33

for the increased fraction of neutrons which might accrue in the 0.1 to 1.0 MeV range in deep penetration situations. A method for estimating the increase in RTNDT as a function of neu-tron fluence and chemistry is given in Regulatory Guide 1.99, Revision 1 [8]. However, the Guide also permits the extrapolation of credible surveillance data by constructing response curves through the data points and parallel to the Guide trend curves, as shown in Figure 8. The Donald C. Cook Unit No. 2 intermediate shell surveillance plates are projected to control the adjusted value of RTNDT through the 32 EFPY design life of the pressure vessel because (1) the unirradiated values of RTNDT for both intermediate shell plates are higher than those of all other reactor vessel materials (14], and (2) the intermediate shell plates are more sensitive than the other core belt1ine surveillance materials are to irradiation embrittlement. Intermediate shell plates C5521-2 (RTNDT.. ( 38°F, % P.. 0.013, and % Cu = 0.14) and C5556-2 (RTNDT.. 58°F, % P.. 0.014, and % Cu.. 0.15) are equivalent surveillance materials in accc~danc~ Tr.i:h ASTM E 185-73, Annex A1 [9]. Westinghouse selected plate C5521-2 for the surveillance program because it had the lower Charpy V-notch upper shelf energy [14] as directed by Figure A1 of ASTM E 185-73. However, the ves-se1 RTNDT projections are based on an unirradiated RTNDT of 58°F reported for plate C5556-2 because it is more conservative to do so. A summary of the projected values of RTNDT for 12 and 32 EFPY of operation of Donald C. Cook Unit No. 2 is plresented in Table VI. 34 (

"~ 600 00 400 III A I 'WI ~ 111 -.~ - -.. ~ I I I iiI II 1 r Ih, I' I', I .-- _. -f-III ~~

J I

Reg. Guide 1.99 J Upper Lim it ~ -."""' l-I tU I-i ~.... III If 200 III H I<" . Y ~. - - ~ .~ I III (J ~ III I-i III 4-1 ~ 100 ~ I--"" ~. ~ t::: t=;r... 1--"""'" i--". f;-i'" ~ I V + I I ,/ I 4-1 I I I 0 l-I t L') m V1 B 60 UI

J....,

'd ~ 40 'd III OJ (J

• rl

'd ,..... ~ h I f!" I---' I I i--J ~If ff Code: .14% Cu ,... -""" ~.. Re' ! Guid~ 1.99

Plate,.

~ 0 Plate C552l-2, Trans. ,/' ~... I T:' 0 Plate C5521-2, Long. ./ ~ I HAZ Material III I*. ~ 20 r-.. -.-_.. i l~ p ... - -. Weld Metal i 2 x 1017 lOIS 1019 Neucron Fluence, n/cm2 (E > 1 MeV) HGURE S. EFF'ECT OF NEUTRON FLUENCE ON RTNDT SHIFT, DONALD C. COOK UNIT NO. 2

TABLE XI PROJECTED VALUES OF RTNDT FOR DONALD C. COOK UNIT NO. 2 Initial EFPY P.V. Material Location RTNDT(a) F1uence(b) tlRTNDT Adj. RTNDT 12 Plate e55~:l: 2 I.D. 58°F 7.4 x 1018 144 202

f. 2.cf

~ C!.-6 5G(' ~:1. 1/4T 58°p 4.4 x 1018 110 168 da.:ta..- 3/4T 58°p 1.1 x 1018 55 113 HAZMa~ 12 LD. 20 0 p 7.4 x 1018 101 121 I 1/4T 20 0 p 4.4 x 1018 78 98 3/4T 20 0 P 1.1 x 1018 40 60 12 Weld Metal I.D. -35°p 7.4 x 1018 81 46 . ~ 1/4T -35°p 4.4 x 1018 63 28 3/4T -35°p 1.1 x 1018 31 -4 32 Plate C5521-! I.D. 58°p 2.0 x 1019 235 293 ( ?3'1 ~C5~ 1/4T 58°p 1.2 x 1019 182 240 3/4T 58°p 3.0 x 1018 91 149 f2TfoJ"DT . ~- -.-._......... 32 HAl Material I.D. 20 0 p 2.0 x 1019 165 185 ~ 1/4T 20 0 p

1. 2 x 1019 129 149 3/4T 20°F 3.0 x 1018 64 84 32 Weld Metal I.D.

-35°p 2.0 x 1019 133 98 ~ 1/4T -35°F

1. 2 x 1019 103 68 3/4T

-35°P 3.0 x 1018 52 17 (a) Reference 14 (b) Neutrons I cm2, E > 1 MeV ( 36

A method for estimating the reduction in Cv upper shelf energy as a function of neutron f1uence is a1s~ given in Regulatory Guide 1.99, Revi-sion 1 [8]. The results from Capsule 1 are compared to a portion of Fig-ure 2 of Regulatory Guide 1.99, Revision 1, in Figure 9. The embritt1ement responses of the pressure vessel plate and RAZ surveillance materials are in good agreement with the prediction of Regulatory Guide 1.99, Revision 1, for 0.15% Cu material. On the other hand, the shelf energy response of the weld metal is well below the corresponding design curve. Referring to the conservative NRC design curves from Regulatory Guide 1.99 shown in Figure 9, the projected Cv shelf energies of the vessel mate- ~ia1s are as follows: Plate C5521-2 (Unirradiated Cy Shelf = 86 ft-1b) 32 EFPY at I.D. -- 62 ft-1b 32 EFPY at 1.41 -- 64 ft-1b 32 EFPY at 3/4T -- 70 ft-lb Pla.t!'! C5'5'56-2 CUnirradiated Cv~e1f =..90 ft::llU. 32 EFPY at LD. 65 ft-1b 32 EFPY at 1/41 67 ft-lb 32 EFPY at 3/41 74 ft-lb Weld Meta1-{Unirradiated Cy Shelf = 97 ft-lb) 32 EFPY at I.D. 75 ft-lb 32 EFPY at 1/41 78 ft-lb 32 EFPY at 3/41 83 ft-lb RAZ Material (Unirradiated Cy Shelf = 109 ft-lb) 32 EFPY at LD. 78 ft-lb 32 EFPY at 1/41 82 ft-lb 32 EFPY at 3/41 89 ft-lb 37

60 ~ ~ C1I tJ ~ 20 C1I p.. ~ ,t>-. co H C1I Jj 'H 10 M C1I ..c: w til 00 Q ',..l' C1I 6 (/) III C1I ~ tJ C1I A 4 2 J I ' ;I!

l.l.

. i* ~m" T

I' I

r.:-. J , -I, I, I ! Ii " i L, Reg. Guide 1.99 I Upper Limit ... -. c-I- ~ i=- i-- Reg. Guide 1.99 0._- .15% Cu Plate ~ ~ .05% Weld 1 f' Cu ~ I' i I ~' I I I 1+ 'I I I I I I Code: 0 Plate CS521-2, Trana. 0 Plate C5521- ~, Lons*

• HAZ Material

~leld Metal 2 x 1017 Neutron Fluence, n/cm2 (E > 1 MeV) FIGURE 9. DEPENDENCE OF Cv UPPER SHELF ENERGY ON NEUTRON FLUENCE. DONALD C. COOK UNIT NO.2 ~,

( These projections indicate that the core beltline materials in the Donald C. Cook Unit No. 2 pressure vessel will retain adequate shelf toughness throughout the 32 EFPY design lifetime. The current Donald C. Cook Unit No. 2 reactor vessel surveillance program removal schedule, revised to conform to ASTM E 185-99 [16J, is summarized in Table XII. There are seven capsules remaining in the ves-sel, of which three are standbys. 39

TABLE XII REACTOR VESSEL SURVEILLANCE CAPSULE REMOVAL SCHEDULE [14] DONALD C. COOK UNIT NO. 2 WOL Removal Equivalent Vessel Capsule Material Time F1uenee T Weld Metal (a.) 4 EFPY at I.D. Y Weld Metal 3 EFPY 11 EFPY at I.D. X Trans. Plate S EFPY E.O.L. at 1/4 T U Weld Metal 9 EFPY E.O.L. at I.D. S Trans. Plate 32 EFPY E.O.L. at 1.0. V Trans. Plate Standby W Trans. Plate Standby Z Weld Metal Standby (a) Removed at 1.08 EFPY 40 ( (

VI. BEATUP AND COOLDOWN LIMIT CURVES FORNOBMAL OPERATION OF DONALD C. COOK UNIT NO. 2 Donald C. Cook Unit No. 2 is a 3391 Mwt pressurized water reactor operated by American Electric Power Service Corporation. The unit has been provided with a reactor vessel material surveillance program as re-quired by 10CFRSO, Appendix H. Surveillance capsule T was removed at the first refueling during the 1979 outage. This capsule was tested as described in earlier sec-tions of this report. In summary, these test results indicate that: (1) The RTNDT of the surveillance materials in Capsule T in-creased a maximum of 80 0 p as a result of exposure to a neutron fluence of 2.2 x 1018 neutrons/cm2 (E > 1 MeV}. .,r-. (2) Based on a ratio of 3.24.bet:ween the fast neutron flux at the radial center of Capsule T and/the maximum incident on the vessel (.lll, the vessel wall fluence at the I.D. was 6.7 x 1017 neutrons/cm2 (E > 1 MeV) at the time of removal of Capsule T. (3) The maximum shift in RTNDT after 12 effective full power years (EFPY) of operation was predicted to be168°P at the 1/4T and 1l3°P at the 3/4T vessel wall locations, as controlled by the L~ter-mediate shell plate material. (4) The maximum shift in RTNDT after 32 EFPY of operation was predicted to be 240 0 p at the 1/4T and 149°F at the 3/4T vessel wall 10-cations, as controlled by the intermediate shell plate material. 41

The Unit No. 2 heatup and cooldown limit curves for 12 EFPY and 32 EFPY have been computed on the bases of (3) and (4) above. The procedures employed by SwRI are described in Appendix B. The following pressure vessel constants were.!!mployed as input data in this analysis: Vessel Inner Radius, ri 86.50 in., including cladding Vessel Outer Radius, ro = 95.2 in. Operating Pressure, Po 2235 psig . Initial Temperature, To 70 0 p Final Temperature, Tf. 550 0 p Effective Coo lan t Flow Rate, Q 134.6 x 106 lbm/hr Effective Flow Area, A 26.72 ft 2 Effective Hydraulic Diameter, D.. 15.05 in. Heatup curves were computed for a heatup rate of 100oP/hr. Since lower ( rates tend to raise the curve in the central region (see Appendix B), these curves apply to all h~ating rates up to 60oP/hr~ Cooldown curves were com-puted for cooldown rates of OOP/hr (stea4y state), 20oP/hr, 40oP/hr, 60°F/hr, and lOQoP/hr. The 20°F/hr curve.would apply to cooldown rates up to 20°F/hr; the 40 oP/hr curve would apply to rates up to 4QoP/hr; the 60°F/hr curve would apply to rates up to 60oP/hr; the lOQoF/hr curve would apply to rates up to lOQoP/hr. The Unit No. 2 heatup and cooldown curves for up' to 12 EFPY are given in Figures 10 and 11; those for up to 32 EFPY are given in Figures 12 and 13. ( 42

G en n.. ~

l en U)

~ n.. ~ en .~ .p. § w ~ ~ 2600 . 2'400

2200, 2000 1800 1600 1400 1200 1000 800 600 400 200 60 REACTOR COOLANT SYSTEM HEATUP LIMITATIONS AP~I PLICABLE FOR FIRST 12 EFFECTIVE FULL POWER YEARS.

(MARGINS OF 60 PSIG AND 10°F ARE IN-. ClUDED FOR POSS I BLE I NSTRltw1ENT ERROR) 1111111 1111111111111 I LEAK TEST LIMIT, MATERIAL PROPERTY BASIS BASE I1:TAl CU = 0.14% ACCEPTABLE INITIAlRTNDT = 58°F OPERATION 12 EFPY RTNDT'(1/4T) = 168°F UNACCEPTABLE , (3/4T) = 113 OF OPERATION n1111111111' 111111 111111111111, mmtllllll~m IlIIIWlm: 1m PRESSURE-TEMPERATURE CR I TI CALI TV. lIMIT FOR H~ATUP RATES LIMIT UP TO 100oF/-iR ,100 150 200 250 300 350 400 AVERAGE REACTOR COOLANT SYSTEM TEfilERAME (OF) FIGURE 10. REACTOR COOLANT SYS'IEM PRESSURE-TEMPERATURE LIMITS VERSUS 100°F/HOUR RATE CRITICALITY LIMIT AND HYDROSTATIC TEST LIMIT, 12 EFPY 450

(!) (J) 0... W gj (J) U) ~ 0... ~ ~ U) .~ ~ .~ 8 u ~ ~ 0:: 2600 2400 2200 2000 1800 1600 1400 1200 1000 800 600 400 200 60 REACTOR COO~T SYSTEM COOLDOWN LIMITATIONS APPLI CARLE FOR FIRST 12 EFFECTI VE FULL PO'IF.R YEA~S. (MARGINS OF 60 PSIG ~ID 10°F ARE IN-CLUDED fOR POSSIBLE INSTRUMENT ERROR.) ""IIIIIIIIIlIllllllIllIIllIIllllllllllllllllIllllIIllIIIIIII MATERIAL PROPERTY BASIS BASE METAL CU = 0.14% INITIAL RTNDT = 58°P 12 EFPY RTrIDT (1/4T) = 168°P UNACCEPTABLE. (3/4T) = 113°F OPERATION MI~~II@1m PRESSURE-TEMPERATURE ACCEPTABLE LIMITS........ OPERATION COOLOO~~ RATE of HR Q~I 20 4~.. 60 100" v 100 150 200 250 300 350 400 AVERAGE REACTOR COOLANT SYSTEM TEMPERATURE (OF) HGlJRE 11. ~ REACTOR COOLANT SYSTEM PRESSURE-TEMPERATURE LIMITS VERSUS COOLDOWN RATES, ] 2 EI~PY 450

(!) (/) 0- ~

l U)

(/) ~ 0-ffi (/) .j:.' ~ VI S 2600 REACTOR COOLNH SYSTEM HEATtJP LIMITATIONS AP-* 2400 PLICABLE FOR FIRST 32 EFFECTIVE FULL P<:MER YEARS. (MARGINS.OF 60 PSIG AND 10°F ARE IN-2200 2000 CLUIED Fffi POSSIBLE INSTRtJv1ENT ERROR) LEAK TEST LI HI T U~lli~~ MATER I AL PROPERTY BAS I S 1800 BASE METAL CU = 0.14% 1600 INITIAL RTNDT = 50°F 32 EFPY RTNDT (l/4T)== ~40oF (3/4T) ~ lli9°F 1400 1200 UNACCEPTABLE ACCEPTABLE OPERATION OPERATION HUJI m IIIIU][llllllll 1000 PRESSUHE-TEMPERATURE ~CRITICALITY LHlIT FOR tiEATUP RATES LIMIT 800 UP TO lOOoF/HR 600 400 200 60 100 150 200 250 300 350 400 450 AVERAGE REACTOR COOLANT SYSTEM TEMPERATURE (OF) FIGURE 12. REACTOR COOLANT SY',TEM PRESSURE-TEMPERATURE LIMITS VERSUS 100 0 F /HOUR RATE CRITICALITY LIMIT,\\ND HYDROSTATIC TEST LIMIT. 32 EFPY

<.!) U) 0- ~

l U)

U) ~ 0- ~ U) -t-' ~ Q'\\ 8

• u

~ ~ 2600 '".11 REACTOR COOLANT SYSTEM COOLDOWN liMITATIONS 2400 APPLICABLE FOR fIRST 32 EF.FECTIVE FULL PGlER YEARS. (MARGINS OF 60 PSIG AND 10°F ARE IN-2200 2000 CLUDED FOR POSSIBLE INSTRUMENT ERROR.) 111111 1111 111111111~m""llllllllllllllllllli 1IIIIIIIIIIIItII MATERIAL PROPERTY BASIS 1,800 BASE METAL CU = 0.14% 1600 INITIAL R~T = 58°F 32 EFPY R NDT (1/4T) = 240°F UNACCEPTABLE (3/4T) = 149°F 1400 1200 III OPERATION 1000 PRESSURE-TEMPERATURE ACCEPTABLE LIMITS . OPERATIOO 800 COOLlXl'lN 600 RATE °F/HR 400 28\\ ~~ 100 200 60 100 150 200 250 300 350 400 AVERAGE REACTOR COOLANT SYSTEM TEMPERATURE (OF) FIGURE 13. REACTOR COOLANT SYSTEM PRESSURE-TEMPERATURE LIMITS VERSUS COOLDOWN RATES, 32 EFPY 450

1.

2.

3.

4.

5.

6.

7.

• 8.

9.

10.

11.

12.

VII. REFERENCES Title 10, Code of Federal Regulations, Part 50, "Licensing of Production and Utilization Facilities." ASME Boiler and Pressure Vessel Code, Section III, "Nuclear Power Plant Components," 1974 Edition. ASTM E 208-69, "Standard Method for Conducting Drop-Weight Test to Determine Nil-Ductility Transition Temperature of Ferritic Steels," 1975 Annual Book of ASTM Standards. Steele, L.E., and Serpan, C. z., Jr., "Analysis of Reactor Ves-sel Radiation Effects Surveillance Programs," ASTM STP 481, December 1970. Steele, L. E., "Neutron Irradiation Embrittlement of Reactor Pressure Vessel Steels," International Atomic Energy Agency, Technical Reports Series No. 163, 1975. ASME Boiler and Pressure Vessel Code, Section XI, "Rules for Inservice Inspection of Nuclear Power Plant Components," 1974 Edition. "Prediction of Shift in the Ductile-Brittle Transition Tempera-tute of LWR Pressure Vessel Materials," prepared by the Metal Properties Council, Subcommittee 6 on Nuclear Materials, July 1, 1980. Regulatory Guide 1.99, Revision 1, Office of Standards Development, U. S. Nuc.lear Regulatory Commission. April 1977. ASTM E 185-73, "Standard Recommended Practice for Surveillance Tests for Nuclear Reactor Vessels," 1975 Annual Book of ASTM Standards. ASTM E 399-74, "Standard Method to Test for Plane-Strain Fracture Toughness of Metallic Materials," 1975 Annual Book of ASTM Standards. Witt, F. J. and Mager, T. R., "A Procedure for Determining Bounding Values of Fracture Toughness ~c at Any Temperature," ORNL-TM-3894, October 1972. "Donald C. Cook Unit No. 2 Reactor Vessel Radiation Surveillance Program," WCAP-85l2, November 1975. 47

ADDENDUM TO FINAL REPORT ON "REACTOR VESSEL MATERIAL SURVEILLANCE PROGRAM rOR DONALD C. COOK UNIT NO.2, ANALYSIS OF CAPSULE Tn Add the following references as page 48:

13.

Let~er, F. Noon of Wes~inghouse to J. R. Jensen of the American Electric Power Service Corporation, Document AEP-80-528, March 19, 1980.

14.

Donald C. Cook Unit No. 2 Technical Specifications.

15.

US NRC Standard Review Plan, NUREG-75/08], Section 5.3.2, Pres-sure-Temperature Limits,,,, November 24, 1975.

16.

ASTM E 185-79, "Standard Practice for Conducting Surveillance 1'es~s for Light-Wa~er Cooled Nuclear Power Reactor Vessels," 1979 Annual Book of ASTM Standards. '0. 48

>... :.~" ... "'"~ ( (

( APPENDIX A TENSILE rEST RECORDS 49

( Southwest Rese~rch Institi. Department of Materials Sciences TENSILE TEST DATA SHEET Test No. T-_-,,/ Est. U. T. S. Spe~. No. L?t ('.(/_/0 Initial G. L. psi /.tJ in. Project No. 0:;- ~9;2 { :;el Machine No. S-o Ie Temperature 1-.2 It::) of Initial Dia.

• *25eJ in.

Date $'- 'Z _ pO Strain Ra.te Initial Thickne s s Initial Width Top Temperature ______ OF Bottom Temperature _____ oF Final Gage Length /, 6 If' Final Dia.meter .. /1./ t, 7 Final Area _ 0 / ~ Z.1.7.7/ U.T.S. = tJ? MOl.Xirn:.lm Load Initial Ar ea = in. in. 2

m.

in. in. Initial Area

• 04/~

./ Maximum Load...z.:ci' 00(:/.

o. Z% Offset Load J.7lJ 4' 0.020/0 Offset Load lb Upper Yield Point ____ lb 71 " -

= o. Z% Offset Load =_ ~/C.f'I. Initial Area ~ 1'"S'%" O. Z% Y. s. 0.02% Y. S. Upper Y. S. % Elongation O. 02% Offset Load = = Initial Area

Upper Yield Point

Initial Area _____ psi _____ psi

Final G. L. - Initial G. L. x 100

Initial G. L. % R.A. Initial Area - Final Area ~ = Initial Area x 100 = 6 $' 0/, Signature: ___________________ _ b~./ I-u-f ~1.x-fSUtif/brf / //~J'tJ (

~ ~ ~ oJ. III * ~ H.. t \\)

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- --_.. ---._-................... ~~l;..l-r;\\J,:l;e .Department of Ma.terials Sciences TENSILE TEST DATA SHEET istrO, T-ec. J.'fo. /11 W-7 Est. U. T. S. psi Initial G. L. ? 0 in. Initial Dia. < 2 S"' CJ in. Project No. 0;' -fr'1:1 '1'. 00 / Machine No. a 0 K o mperature f" 55'0 OF 'ain Rate -----

---

~--

Date .. ~- ~/ - PO Initial Thickne s s in. Initial Width _____ in. c Top Temperature,..!5:S-r' Bottom Temperature b $/ Final Gage Length. /. I Cj / Final Diameter , /.5 a in. in. ,OJr'P1 ~. 2 Final Area ______ ~~f::.r_I-- in. 2)J

u. T. S. =

0.2% Y. S. Maximum Load Initial Area =

o. *2,"0 Offset Load

= = Initial Ar ea 0.020/0 Y. S. = O. 02% Offset Load = Initial Area Upper Y. S. = Upper Yield Point = Initial Area II ** / Maximum L~ad..,....,.4- ~ I,n~,.lb O. 20/0 OOset Load 332. P ~b jjJJ 0.02% Offset Load k Ib Upper Yield Point ____ Ib _____ psi _____ psi (If 1 Final G. L. - Initial G. L. x 100 = -{O E ongation = Initial G. L. ~. { q I rio / % % R. A. = Initial Area - Final Area Initial Area ~1\\ XI00~~% t/.

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. (" Southwest Research Instil Department of Materials Sciences TENSILE TEST DATA SHEET Test No. T------ (~ Project No. e~_f5"f ~ 7-00 / Est. U. T. S. _____ psi Spec. No. ~ 1,5 - 10 Initial G. L. _~::..:..' _0 __ in. Machine No. $" to I~ Initial Dia. ,*:2 yf in. Date ~- 2-90 Strain Rate Initial Thickness in. Initial Area Initial Width _____ m. 7 Top Temperature ______ of Ma.xim.um Load 4-52? q, _ JlS' ISJ// Bottom Temperature _____ of Final Gage Length t,\\ <3 \\ in. Final Diameter

• 1 ~S? 7 in.

Final Area ( t.:/'::;' 21fp 2

m.

U.T.S. = 0.2% Y. S.

o. 02% Y. s.

Upper Y. S. Moucimum Load = = Initial Area

o. 2% Offs et Load Initial Ar ea

o. 02% Offset Load

Initial Area

Upper Yield Point

Initial Area 0.2% Offset Load.q".f'~ lb 0.020/0 Offset Load Ib Upper Yield Point ____ Ib _____ psi

---

psi

/ at 1 Final G. L. - Initial G. L. x 100 = / a.. /',ffi 1710 ,0 E ongation, = Initial G. L. /. _..- ,I / = Initial Area - Final Area x 100 = ~// I~- at 0/0 R. A. Initial Area Zi --r '" "ft)/o dfJ Signature: (

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ure r~ $' <!) of' Initial Dia.

.24-9 in. Date ___ r:..-,-_C)_o_-........ J'O __ :;/' te Initial T hickne s s in. Initial Width in. Initial Area.Q%~ I. '/ ( 'P Temperature D s I Ittom Temperature 5 LiP etal Gage Length /. / eo J Maximum Load 4-'3too (lIb in. 1:// 0.2'0 Offset Load 3'2.LtD IJ; Ib /' 0.02% Offset Load Ib nal Diameter _-,-, "-/"~::......IIf ____ in. Upper Yield Point ____ lb

1al Area _--:.'_o=--'Z._~_y_~,g,;..,... __ in. 2

~% Y. s. )2% Y. S. ( 7 Ma.xim.:.lm Load ~ u_ Initial Area = 0 e, :3 ~1M'i r:s: /

0.20/0 Offse-e Load

/~ S . t<. Initial Area ~~. ~ Sf

0.02% Offset Load

Initial Area _____ psi = Upper Yield Point =' psi Initial Area ~longation = Final G. L. - Initial G. L..... 100 = Initial O. L.

t. A. =

x 100 = Initial Area Initial Area - Final Area ~JI -{R~ ~~ ~~t!&z{ tPj ~(;7/~/tJ

II ~ ~ ~:~ " ,~ '.r:: ' '" ";I~. ~, i ~.. <=; ~ ("'\\ \\:Q ~ --=- -.- -- ~ .j ~ ...l-" ~ V, ~ ~ u~ ~I,. ~*I Qil N\\ ~ Cf'1-. I, -=-...

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(

( APPENDIX B PROCEDURE FOR THE GENERATION OF ALLOWABLE PRESSURE-'I'EMPERA'l'URE LIMIT CURVES FOR NUCLEAR POWER PLANT REACTOR VESSELS 63

PROCEDURE FOR THE GENERATION OF ALLOWABLE PRESSURE-TEMPERATURE LIMIT CURVES FOR NUCLEAR POWER PLANT REACTOR VESSELS A. Introduction The following is a description of the basis for the generation of pressure-temperature limit curves for inservice leak and hydrostatic tests. heatup and cooldownoperations, and core operation of reactor pressure vessels. The safety margins employed in these procedures equal or eXgeed those recommended in the ASME Boiler and Pressure Vessel Code, Section m, Appendix G, "Protection Agamst N.onductile Failure. " B.

Background

The basic parameter used to dete rmine safe ves sel operational conditions is the str,ess intensity factor, Kr, which is a function,of 'the stress state and flaw configuration. The Kr corresponding to membrane tens ion is given by where Mm is the membrane stress correction factor for the postulated' flaw and CJ"m the membrane stress. Likewise, Kr corresponding to bend-ing is given by ( ( (1) where Mb is the bending stres s correction factor and crb is the bending C stress. For vessel section thickness of 4 to 12 inches, the maximum

c postulated surface flaw, which is assumed to be normal to the direction of maximum stress, has a depth of 0.25 of the section thickness and a length of 1.50 times the section thickness. Curves for Mm versus the square root of the vessel wall thickness for the postulated flaw are given in Figure 1 as taken from the Pressure Vessel Code (ref. Figure G-2114. l)c These curves az:e a function of the stress rati~ parameter (1"1(11' where (11 is the material yield strength which is taken to be 50,000 psi. The bending correction factor is defined as 2/3 Mm and is therefore determined from Figure 1 as well. The basis for these curves is given in ASME Boiler and Pressure Vessel Code, Section Xl. "Rules for Inservice Inspection of Nu-clear Power Plant Components, /I Article A-3000. ( The Code specifies the minimum KI that can cause failure as a fUllc~ tion of material temperature, T. and its reference nil ductility temperaturep .tor, KIR.. and is given by Km = 26777. + 1223. exp [0. 014493(T - RTNDT + 160) ] (3 ) where all. temperatures are in degrees Fahrenheit. A plot of this expression is given in Figure 2 taken from the Code (ref. Figure G-2010.1). C. Pressure-Temperature Relationshi~s

1.

Inservice Leak and Hydrostatic Test During performance of inservice leak and hydrostatic tests, the reference stress intensity factor, KLlt. must always be greater than

3.8 3.6 3.4 3.2 3.0 2.8 2.G e

s 2.4 2.2 2.0

.i.S i--""';---;----i---iI_-+--';'--+--+--4--.- -.. ---- --+1--4 LG I 1.4 1.2 1.0 1.0 1.2 1.4 1.6 I.a 2.0 2.2 2.4 2.6 2.8 3.0 3.2 3.43.6 3.a 4.. 0 V THICKNESS (IN.) FIGURE 1. STRESS CORRECTION FACTOR (

( 110 IGO I --L. -J [I 150 140 130 120 /10 ~'CO ._ 90 ~ 60 ~ ~ 10 f-- (!~ IR - 26.117) ~1.223 eO.OI-4493 (T-fRT NOT -1601) 7 \\'mERE 7 !(lR

• nC:FERENCE STRESS INTENSITY FACTOR 7

T

• TEMPERATURE AT WHICH KIR

~I-IS PERMITTED, *F / RTNDT

• REFERENCE

'~IL-OUC11LITY I TEMPERATURE. I / / GO SO 40 30 i--- _.- / ./ / -~ 20 ( I 10 o I 1_ -240 -200 -IGO -120 -eo -40 0

• .fo eo

.120 IGO 200 240 TEMPERATURE RELATIVE TO RiNOT.(T-RT~DT)' FAMF:E:lHEIT DEGREES FIGURE 2. REFERENCE STRESS INTENSITY FACTOR

1. 5 times the KI caused by pressure, thus (4 )

or (5) For a cylinder with inner radius ri and outer radius r 0' the stress distribution due to internal pres:sure is given by (6 ) With l/4T flaws possible at both inner and outer radial locations, i. e., ( will occur at the inner flaw location, thus (7) With the operation pressure known, i. e., POI we deter-mine the minimum coolant temperature that will satisfy Equation (4) by evaluating. t8 ) and determine the corresponding coolant temperature, T, from Equa-tion (3) for the given RTNDT. at the 1/ 4T location. For this calculation, Equation (3) takes the form T = RT (l/4T) - 160 +68 9988 lnfKIR - 26777. ] NDT L. 1223. (9 )

The inservice curves are generated for an operating pres-e ilre range of.96 Po to 1.14 Po. where Po is the design operating pressure.

z.

Heatup and Cooldown Operations At all times during heatup and cooldown operations, the ref-erence stress intensity factor. Km., must always be greater than the swn of Z times the KIp caused by pressure and the KIt caused by thermal gra-dients, thus z.o KIp + 1. 0 KIt < Km. (10) or Z. 0 Mm a'max = Km - Kit (ll ) ("'here 11' is the maximum allowable stress due to internal pressure, max and KIt is the equivalent linear stress intensity factor produced by the thermal gradients. To obtain the equivalent linear stress intensity fac-tor due to therma~ gradients requires a detailed thermal stress analysis. The details of the required analysis are given in Section D. During heatup the radial stres s distributions due to internal pressure and thermal gradients are shown schematically in Figure 3a. Assuming a possible flaw at the 1/4T l'ocation, we see from Figure'3a that the thermal stress tends to alleviate the pressure stress at this point in the vessel wall and, therefore, the steady state pressure stress (Wauld re pre sent the maxiznum s tres s condition at the 1/4 T location. At*

OUTER RAD IUS 3/4T + 1/4T INNER RAD IUS Pressure stress distribution Therll1a-t-stress distribution ( a) Heatup OUTER RAD IUS 3/4T - + 1I4T INNER RADIUS Pressure stress distribution Thermal stress distribution ( b) Cooldown Figure 3. Heatup and Cooldewn Stress Distribution l -c

the 3/4T flaw location. the pressure stress ~d thermal stress add and. therefore. the combination for a given heatup rate represents the maxi-mum stress at the 3/4T location. The maximum overall stress between the l/4T and 3/4T location then determines the maximum allowable reac-tor pressure at the given coolant temperature. The heatup pressure-temperature curves are thus generated by calculating the maximum steady state pressure based on a possible flaw \\at the 1/4T location from (lZ) where Mm is determined from the curves in Figure land Km is obtained from Equation (3) using the coolant temperature and RTNDT at the l/4T location. Here we may note that Mm must be iterated for since it is a function of the final stress ratio to yield strength <<(1'I(1'y)' At the 3/4T location, the maximum pressure is determined from Equation (11) as (13 ) where Km is obtained' from Equation (2) using the material temperature and RTNDT at the 31 4T location and KIt is determined from the analys is procedure outlined in Section D. Mm is determined from Figure 1.

The minimum of these maximum allowable pressures at ( the given coolant temperature determines the maximum operation pressure. Each heatup rate of interest must be analyzed on an individ-ual basis. The cooldown analysis proceeds in a similar fashion as that described for heatup with the following exceptions: We note from Figure 3b that during cooldown the 1/ 4T location always controls the maximum stress since the therma.l gradient produces tensile stresses a.t the 1/4T location. Thus the steady state pressure is the same as that given in Equation (12). For each cooldown rate, the maximum pressure is evalu-ated at the 1/ 4T location from ( . (14) where Km is obtained from Equation (3) using the material temperature and RTNDT at the 1/4T location. KIt is determined from the thermal analysis described in Section D. It is of interest to note that during cooldown the material temperature will lag the coolant temperature and, therefore, the steady state pressure, which is evaluated at tp,e coolant temperature, will ini-tially yield the lower maximum allowable pressure. When the thermal gradients increase, the stresses do likewise, and, finally, the transient analysis governs the maximum allowable pressure. f;lence a point-by-point (

om(ariSOn must be made between the maximum allowable pressures pro-luce.... by steady state analyses and transient thermal analysis to determine tte: minimum of the maximum allowable pressures.

3.

Core Operation At all times that the reactor core is critical, ~e temperature lust be higher than that required for inservice h.ydrostatic testing~ and in ddition, the pressure-temperature relationship shall provide at least a o of margin over that required for heatup and cooldown operations. Thus te pressure-temperature limit curves for core operation may be constructed Lrectly from the inservice leak and hydrostatic test and heatup analysis ~sults. Thermal Stress Analysis The equivalent linear stress Que to thermal gra.dients is obtained m in, the vessel wa.ll is governed by the partial differential equation ( 15) .bject to initial condition T(r,O) = To ' (16) ,dboundary conditions ( 17)

and (18 ) where (19 ) P is the material density, c the material specific heat, K the heat conduc-tivity of the material, h the heat transfer coefficient between the water coolant and vessel material. R the heating rate, To the initial coolant temperature, T(r, t) the temperature distribution in the vessel, r the spatial coordinate, and t the temporal coordinate. A finite difference solution procedure is* employed to solve for the ( radial temperature distribution at v.arious time steps along the heatup or cooldown cycle. The finite difference equations for N radial points, at distance 60r apart, across the vessel are: for 1 < n < N t+6ot r 60tK 60r ] Tn = Ll - pc(Ar)2 (2+;;) T~ 60tK [6or t t l + pc(Ar)2 ( 1 +;; ) Tn + 1 + Tn - 1 J ' (20 ) for n = 1 Tt+At = [1 _6otK (1 +~) _ 60th l...J Tt 1 pc(Ar)2 r1 pc{Ar) 1 60tK [ 60r t 60rh tJ + pc (6.r )2 (1 +;r) T 2 + K T c (21 ) (

c and for n = N For stability in the finite difference operation, we must choose dt for a given dr such that both dtK dr < (.4. }Z (2 +-) _ 1 pc Qr r 1 and dtK /).r /).th z{l+-)+ ~1 pc(/).r) r 1 pc CAr) are satisfied. These conditions assure us that heat will not flow in the C direction of increasing temperature, which, of course, would violate the second law of thermodynamics. Since a large variation in coolant temperature is considered. the dependence of (K/ pc), K, and h on temperature is included in the analysis by treating these as constants only during every 5 OF increment in coolant temperature and then updating their values for. the next 5 OF increment. (22 ) The dependence of (K/ pc) called the thermal diffusivity and K, the thermal conductivitYI can be determined from the ASME Boiler and Pressure Ves~ sel Code. Section m, Appendix I - Stress Tables. A linear regression analysis of the tabular values res.ulted in the following expres sions: K(T) = 38.211 - 0.01673

• T (BTU!HR-FT- "F)

(25)

and ( k{T) = (K/pc) = 0.6942 - 0.000432

• T (FTZ/HR)

(26 ) where T is in degrees Fahrenheit. The heat transfer coefiic ient is calculated based on forced con-vection under turbulent flow conditions. The variables involved are the mean velocity of the fluid coolant, the equivalent (hydraulic) diameter of the coolant channel, and the density, heat capacity, viscosity. and thermal conductivity of the coolant. For water coolant, allowance for the variations in physical properties with temperature may be made by writing* (27) where v is in ft/ sec, D in inches, the temperature is in of, and h is in ( Btul hr-ft2 - of. The values for the heat-transfer coefficient given by this relationship are in good agreement with those obtained -from the Dittus-Boelter equation for temperature s up to 600 of. The mean velocity of the coolant, v, is generally given in terms of the efiective coolant flow rate 0 (Lbm/hr) and effective flow area A (£t2). Given the relationship peT) = 62.93 - 0.48 x 10-Z

• T - 0.46 x 10-4
• T2 (28 )

for the density of water as a function of temperature, the mean velocity of the coolant is obtained from v = 0/(3600

• o(T)
• A).

(29)

• Glasstone, S., Principles of Nuclear Reactor Engineering.. D. Van C

Nostrand Co., Inc., New Jersey. pp. 667-668, 1960.

c The thermal stress distribution is calculated from

a. E [ 1 I r 1

rZ + r i Z r r 0 ] I7'T(r, t) = r:v 7 T(r, t) rdr

• T(r, t) + ;"z (r Z _ r.Z ) J T(r, t) rdr r'

0 1 r' 1 1 (30 ) where a. is the coefficient of thermal expansion (in/in -F), E is Young's modulus, and \\J is Poisson!s ratio. This expression can be obtained from. Theory of Elasticity by Timoshenko and Goodier, pp. 408*409, wh.en im* ' posing a zero radial stress condition at the cylinder inner and outer radius. Poisson's ratio is taken to be constant at a value of 0.3 while a. and E are evaluated as a function of the average temperature across the vessel c 2 ro Tavg = (r 2. r'Z) I T{r)rdr. (31) o 1 r' 1 The dependence of the coefficient of thermal expansion on temperature is taken to be a.(T) = 5.76 x 10-6 + 4.4% 10-9

• T (32 )

and the dependence of Young's mod,ulus on temperature is taken to be E(T) = 27.9142 + 2.5782 % 10-4

• T - 6.5723 x 10-6
• TZ (33)'

as obtained from regression analysis of tabular values given in Section m. Appendix I of the ASME Boiler and Pressure Vessel Code. " The resulting stress distribution given by Equation (30) is not linear; however, an equivalent linear stress distribution is determined ( from the resulting moment. The moment produced by the nonlinear

stres s distribution is given by ro M(t) = b f a'T (r, t) rdr ro 1 where b is a unit depth of the vessel. Here we note that the moment is a function of time, i. e., coolant temperature via T c = To + Rt. For a lin-ear stres s distribution we have that Mc a'max = T where a'max is the maximum outer fiber stress, c the distance from the neutral axis, taken to be (ro - ri)/Z, and I the section area moment of inertia which is given by Combining these expressions results in the equivalent linear stress due to thermal gradients The thermal stress intensity factor KIt is then defined as where Mb is deterII?-ined from the curves given in Figure 1 wherein Mb = 2/3 Mm " It is of interest to note that a sign change occurs in the stres s calculations during a cooldown analysis since the thermal gradients (34) (35 ) ( (36 ) (37) (38 ) (

produce compressive stresses at the vessel outer radius. This sign ( change must then be reflected in the Kit calculation for the cooldown analysis. Normalized temperature and thermal stress distributions during a typical reactor heatup are given in Figure 4. The radial temperature is shown normalized with respect to the average temperature, T avg' by T_-_T..:a;:.v.:.,jg=--_ T = (T - Tavg}max (39 ) The thermal stress and equivalent linearized stress, as calculated by Equations (30) and (37), are normalized with. respect to the maximum thermal stress. Here we note that the actual thermal stress at the 3/4T ( location is considerably less than the maximum. equivalent linear stress which yields additional safety margins during the beatup cycle. Similar down. The trends are nearly identical as those shown in Figure 4 wh.en the inner and outer vesseIlocations are reversed with. the 1/4T location becoming the critical point. E. Exam pIe Calc ulations The following example is based on a reactor ves sel with the follow c ing characte ristics: Inner Radius = 82.00 in. Oute r Ra dius = 90.00 in. ( Operating Pressure = 2250 psig

OUTER WALL 1.0 0.8 0.6 0.4 0.2 0 -1.0 0 1.0 -1.0 INNER WALL Normalized temperatu re distribution ( 6. T {6. Tmax ) 0 1.0 Normalized stress distribu'tion ( a lama~' ) Figure 4. Typical Normalized Temperature and Stress Distribution During Heatup (

( Initial Temperature = 70°F (To) Final Temperature = 550°F (Tf) Effective Coolant Flow Rate = 100 x 106 Lbm/hr (Q) Effective Flow Area = ZOo 00 ftl (A) Effective Hydraulic Diameter = 10.00 in. (D) RTNDT (1/4T) = ZOO°F RTNDT (3/4T) 140°F In the thermal stress analysis Z1 radial points were Ilsed in the finite difference scheme. Going from 70°F to the final temperature of 550 of. approximately 1Z, 000 tiIne (temperature via T = To + Rt) steps were required in the thermal analysis for the 100°F /hr heatup rate. The ( results of the computation are shown in Figures 5 through 9. Figure 5 gives the reference stress intensity factor, Km, as a analysis, Km is converted directly to allowable pressure via Equation lZ. During the heatup and cooldown thermal analyses the material tem-perature at the 1/4T and 3/4T and thermal stress intensity factors KIt are required to compute allowable pressure via Equations (13) and (14). The material temperatures versus coolant temperature during the 100 OF /hr heatup and cooldown analyses are given in Figure 6. These temperatures allow computation of the correspollding reference stress intensity factors, KIR (3/4T) and KIR (l/4T). Figure 7 gives the corresponding thermal ( stress intensity factor at the 3/4T and 1/4T locations as a function of coolant tempe rature.

, 200r-------.--------.------~------~~------------------------ 160 ~ 120 ~ 40 o~------~------~--,----~--------~------~------~------~ 50 100 150 200 250 300 350 400 TEMPERATURE (OF) Figure 5. Reference Stress I ntensity Factor as a Function of Temperature I ndexed to RTNDT ( 1I4T )

400 -- 100 of 1 HR HEA TU f" ( 3/4 T location ) -- 100°F IHR COOLDOWN ( 1I4T location) 300 LI-0 - L.LJ a::: t-<< O!: 200 w a..

f w

.....J W Vl (/) W > 100 o~ __ ~ __ ~ ______ ~ ______ ~ ________ ~ ______ ~ ______ ~ ____ ~ 50 100 150 200 250 300 350 COOLANT TEMPERAllJRE (OF) Figure 6. Vessel Temperature at 1I4T and 3/4T locations as a Function of Coolant Temperature

~. V>> ~ -...... t-t

w:::

12r-------~------_r------~--------~------~------~----~ 10 8 6 ---. 100 of IHR HEATUP ( 3/4T location) 4 --- 100 of IHR COOlDOWN ( 114 location )

2 o

5~0--~--~lOO~-----1~~~----~~~----~2~50------~300~----~35~O----~ COOLANT TEMPERATURE (OF) ~re 7. Thermal Stress Intensity factor at 3t!{ and 1I4T locations as a Function of Coolan~m~erl~ure

( Figures 8 and 9 demonstrate the construction of the allowable com-posite pressure and temperature curves for the lOO°F/b.r b.eatup and cool-down rates. The composite curves represent the lower bound of the thermal and steady state curves with the addition of margins of +lO°F and -60 psig for possible instrumentatiol... error. Figure 8 also shows the leak te'st limit, corrected for instrument error, as obtained from Equation (9). The limit points are at the operating pressure 2250 psig and at 2475 psig which cor-responds to 1. 1 times the operating pressure. The criticality limit is also shown in Figure 8 and is constructed by provicting for a 40 OF" margin over that required for heatup and cooldown and by requiring that the minimum temperature be greater than that required by the leak test limit. (

2400 . lEAK TEST lIMI T 400 50 100 150 200 250 300 350 INDICATED TEMPERATURE (OF) figure 8. Pressur~emperature Curves for 100°F IHr Heatup

2000 .. 01 COMroSITE CURVE -100°F JHRCOOlDOWN VI c.. (Margins()f + 10°f and -60 psig fo(instrument error) LtJ 1600 a:::

l Vl Vl LtJ

.0:= a.. 0 1200 ~ "'U 0 z 800 ~ 50 100 150 200

• 250 300 350 INDICATED TEMPERATURE" (OF) figure. 9. Pressure -Temperature Curves for 100°f IHr Cool~own

( . (

/ \\ PC~~ol Wcl<:/ol J.lcX~DI v

p. s+/ A(P I 31 V

AJIIl '. tVL,(L42~f-tUl-' ) t.u - 38}}