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Reactor Vessel Matl Surveillance Program for Facility, Analysis of Capsule T.
	ML17326A522
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Site:	Cook
Issue date:	12/08/1977
From:	Norris E; SOUTHWEST RESEARCH INSTITUTE
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	02-4770, 2-4770, NUDOCS 8002270331
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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. 1 ANALYSIS OF CAPSULE T by E. B. Norris FINAL REPORT SwRI Project 02-4770 to American Electric Power Service Corporation 2 Broadway New York, New York 10004 J

December 8, 1977 Approved:

'4 $ }%$ $ . $

~ h $ i \ 'l ~ h 'I ~ P$ ~ i S. Lindholm, Director A',"}:=:"}lCP,Ii EL:.O'IHiC PG"'/ II SEl'}VLCC"- CORi.

Department of Materials Sciences

~c DATE 80 f'2 2V0 53 ( i

TABLE OF CONTENTS

~Pa e r

LIST OF TABLES iii LIST OF FIGURES

SUMMARY

OF RESULTS AND CONCLUSIONS BACKGROUND III. DESCRIPTION OF MATERIAL SURVEILLANCE PROGRAM IV. TESTING OF SPECIMENS FROM CAPSULE T 13 V. ANALYSIS OF RESULTS 35 VI. HEATUP,AND COOLDOWN LIMIT CURVES FOR NORMAL OPERATION OF DONALD C. COOK UNIT NO. 1 VII. REFERENCES 47 j

APPENDIX A .TENSILE TEST RECORDS A-1 APPENDIX B PROCEDURE FOR THE GENERATION OF ALLOWABLE B-1 PRESSURE-TEMPERATURE LIMIT CURVES FOR NUCLEAR POWER PLANT REACTOR VESSELS

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~ lb LIST OF TABLES Table ~Pa e Donald C. Cook Unit No. 1 Reactor Vessel Sur-veillance Materials Summary of Reactor Operations 16 Donald C. Cook Unit No. 1 Summary of Neutron Dosimetry Results 17 Donald C. Cook Unit No. 1 Capsule T IV Fast Neutron Spectrum and Iron Activation 19 Cross Sections for Capsule T Charpy V-Notch Impact Data 21 The Donald C. Cook Unit No. 1 Reactor Pressure Vessel Intermediate Shell Plate B4406-3 (Longitudinal Direction)

VI Charpy V-Notch Impact Data 22 The Donald C. Cook Unit No. 1 Reactor Pressure Vessel Intermediate Shell Plate B4406-3 (Transverse Direction)

VII Charpy V-Notch Impact Data 23 The Donald C. Cook Unit No. 1 Reactor Pressure Vessel Core Region Weld Metal VIII Charpy V-Notch Impact Data 24 The Donald C. Cook Unit No. 1 Reactor Pressure Vessel Core Region Weld Heat-Affected Zone Metal IX Charpy V-Notch Impact Data 25 A533 Grade B Class 1 Correlation Monitor Material Notch Toughness Properties of Capsule T Specimens 31 Donald C. Cook Unit No. 1 XI Tensile Properties of Surveillance Materials 32 Capsule T XII Projected Values of RTNDT for Donald C. Cook 40 Unit No. 1 for Up to 12 EFPY of Operation

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LZST OF TABLES (CONT'D.)

Table ~Pa e ZIZZ Projected Values of RTNDT for Donald C. Cook 41 Unit No. 1 for Up to 32 EFPY of Operation XZV Proposed Reactor Vessel Surveillance Capsule 42 Schedule Donald C. Cook Unit No. 1

LIST OF FIGURES

~Ft ure ~Pa e Arrangement of Surveillance Capsules in the Pressure Vessel 2 Vessel Material Surveillance Specimens 3 Arrangement of Specimens and Dosimeters in 12 Capsule T 4 , Charpy V-Notch Properties of Plate B4406-3 26 (Long.)

Donald C. Cook Unit No. 1 Surveillance Program Charpy V-Notch Properties of Plate B4406-3 27 (Trans.)

Donald C. Cook Unit No. 1 Surveillance Program Charpy V-Notch Properties of Core Region Meld 28 Metal Donald C. Cook Unit No. 1 Surveillance Program Charpy V-Notch Properties of Core Region HAZ 29 Material Donald C. Cook Unit No. 1 Surveillance Program Charpy V-Notch Properties of Correlation Monitor 30 Material Donald C. Cook Unit No. 1 Surveillance Program Dependence of Cv Shelf Energy on Neutron Fluence, 37 Donald C. Cook Unit No. 1 10 Effect of Neutron Fluence on RTNDT Shift, Donald 38 C. Cook Unit No. 1 Donald C. Cook Unit No. 1 Reactor Coolant Heatup 45 Limitations Applicable for Periods Up to 12 Effective Full Power Years 12 Donald C. Cook Unit No. 1 Reactor Coolant Cooldown 46 Limitations Applicable for Periods Up to 12 Effective Full Power Years

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I.

SUMMARY

OF RESULTS AND CONCLUSIONS The analysis of the first material surveillance capsule removed from the Donald C. Cook Unit No. 1 reactor pressure vessel led to the following conclusions:

(1) Based on a calculated neutron spectral distribution, Capsule T received a fast fluence of 1.80 x 101 neutrons/cm2 > 1 MeV.

(2) The surveillance specimens of the core beltline materials ex-perienced shifts in transition temperature of 75' to 130 F as a result of the above exposure.

(3) The weld metal and heat affected zone (HAZ) materials exhibited the largest shift in RTNDT. However, because the intermediate shell plate material has a high initial (unirradiated) RTNDT, it will control the heatup and cooldown limitations at least until the next surveillance- capsule is removed.

(4) The estimated maximum neutron fluence of 6.92 x 1017 neutrons/

cm > 1 MeV received by the vessel wall accrued in 1.27 full power years.

Therefore, the projected maximum neutron fluence after 32 effective full power years (EFPY) is 1.74 x 1019 neutrons/cm > 1 MeV. This estimate is based on a lead factor of 2.6 between Capsule T and the point of maximum pressure vessel flux.

(5) Based on Regulatory Guide 1.99 trend curves, the projected maxi-mum shift in ductile-brittle transition temperature of the Donald C. Cook Unit 1 vessel core beltline plates at the 1/4T and 3/4T positions after 12 EFPY of operation are 110 F and 50 F, respectively. These values were used as the bases for computing heatup and cooldown limit curves for up to 12 EFPY of operation.

(6) The maximum shifts in the transition temperature of the Donald C. Cook unit 1 vessel core beltline plates at the 1/4T and 3/4T positions after 32 EFPY of operation are pro)ected to be 180 F and 83 F, respectively.

(7) Since the weld metal and HAZ beltline materials are more sensi-tive to radiation embrittlement than the intermediate shell plate material, the operating limf.tations may come under control of the weld metal and HAZ material late in the 32 EFPY. design life of the plant.

(8) The Donald C. Cook Unit No.' vessel plates, weld metal and HAZ material located in the core beltline region are projected to retain suffi-cient toughness to meet the current requirements of 10CFR50 Appendix G throughout the design life of the unit.

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 ferritic materials, the allowable loadings depend on the reference stress intensity factor (KIR) curve indexed to the reference nil ductility temperature (RTNDT) presented in Appendix G, "Protection Against Non-ductile Failure,"

of Section III of the ASME Code.( ) Further, the materials in the beltline region of the reactor vessel must be monitored for radiation-induced changes in RTNDT per the requirements of Appendix H, "Reactor Vessel Material Surveil-lance Program Requirements," of 10CFR50.

The RTNDT is defined in paragraph NB-2331 of Section III of the ASME Code as the highest of the following temperatures:

(1) Drop-weight Nil Ductility Temperature (DW-NDT) per ASTM E 208; (2) 60 deg F below the 50 ft-lb Charpy V-notch (Cv) temperature; (3) 60 deg F below the 35 mil C temperature.

The RTNDT must be established for all materials, including weld metal and heat affected zone (HAZ) material as well as base plates and forgings, which com-prise the reactor coolant pressure boundary.

It is well established that ferritic 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).( ) Also, it has been established that tramp elements, particularly copper and

Superscript numbers refer to references at the end of the text.

phosphorous, affect the radiation embrittlement response of ferritic mate-rials.( ) The relationship between increase in RT~T and copper content is not defined completely. For example, Regulatory Guide 1.99, originally issued in July 1975, proposed an adjustment to RT~T proportional to the square root of the neutron fluence. westinghouse Electric Corporation, in their comments on the 1975 issue of Regulatory Guide 1.99( ), believed that the proposed relationship overestimates the shift at fluences greater than 1.9 x 1019 and underestimates the shift at fluences less than 1.9 x 10 On the other hand, Combustion Engineering, in their comments on the 1975 is-sue of Regulatory Guide 1.99 , suggested that the proposed relationship is overly conservative at fluences below 1019 neutrons per cm (E > 1 MeV) .

There is also disagreement concerning the prediction of Cv upper shelf re-sponse to exposure to neutron irradiation.( ) After reviewing the comments and evaluating additional surveillance program data, the NRC issued a revision to Regulatory Guide 1.99 which raised the upper limit of the transition tem-perature adjustment curve. In this report, estimates of shifts in RTNDT are based on Revision 1 of Regulatory Guide 1.99 ), issued in April 1977.

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 region. of high neutron flux density. (10) describes in such a ASTM E 185 the current recommended practice for monitoring and evaluating the radiation-in-duced changes occurring in the mechanical properties of pressure vessel belt-line materials.

Westinghouse has provided such a surveillance program for the Donald C., Cook Unit No. 1 nuclear power plant; The encapsulated Cv specimens are located near the O.D. surface of the thermal shield at a point where the fast neutron flux density is about three times that at the adjacent vessel wall surface. Therefore, the increases (shifts) in transition temperatures of the materials in the pressure vessel are generally less than the corre-sponding shifts observed in the surveillance specimens. However, because of azimuthal variations in neutr'on flux density, capsule fluences may lead or lag the maximum vessel fluence in a corresponding exposure period. For example, Capsule T (removed during the 1977 refuelling outage) was exposed to a neutron fluence approximately 2.6 times that at the maximum exposure point on the vessel I.D., while Capsule X (scheduled for removal at a later date) is being exposed to a neutron flux about 60% of that at the point of maximum vessel exposure. The capsules. also contain several dosimeter mate-rials for experimentally determining the average neutron flux density at each capsule location during the exposure period.

The Donald C. Cook Unit No. 1 material surveillance capsules also in-clude tensile specimens as recommended by ASTM E 185. At the present time, irradiated tensile properties are used primarily to indicate that the mate-rials tested continue to meet the requirements of the appropriate material specification. In addition, the degree of radiation hardening indicated by the tensile yield strength is used to judge the credibility of the surveil-lance data.(7)

Wedge opening loading (WOL) fracture mechanics specimens, machined from plate material and weld metal, are also contained in the capsules. Current technology limits the testing of these specimens at temperatures well below

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the minimum service temperature to obtain valid fracture mechanics data per ASTM E 399~ ~, "Standard Method of Test for Plane-Strain Fracture Toughness of Metallic Materials." However, recent work reported by Mager and Mitt~1 ~

may lead to methods for evaluating high-toughness materials with small frac-ture 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 1977 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. 1 power plant.

III. DESCRIPTION OF MATERIAL SURVEILLANCE PROGRAM The Donald C. Cook Unit No. 1 material surveillance program is described in detail in WCAP 8047(13), dated March 1973. Eight materials 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 cap-sule is opposite the vertical center of the core. The neutron flux density at the Capsule T location leads the maximum flux density on the'vessel I.D.

by a factor of 2.6.( The capsules each contain Charpy V-notch, tensile and WOL specimens machined from the SA533 Gr B plate, weld metal and heat affected zone (HAZ) materials located at the core beltline plus Charpy V-notch specimens machined from a reference heat of steel utilized in a num-ber of Westinghouse surveillance programs.

The chemistries and heat treatments of the vessel surveillance mate-rials 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 HAZ specimens were machined from a stress-relieved weldment which joined sections of the inter-mediate shell course. HAZ specimens were obtained from the plate B4406-3 side of the weldment. The longitudinal base metal C specimens were oriented with their long axis parallel to the primary rolling direction and with V-notches perpendicular to the major plate surfaces. The transverse base metal C specimens were oriented with their long axis perpendicular to the primary rolling direction and with V-notches perpendicular to the major plate surfaces. Tensile specimens were machined with the longitudinal axis parallel to the plate rolling direction. The WOL specimens were machined

X (220')

270' (184') Y (320')

180'a S Z (356 )

(4')

V (176')

T (40) 0 90 u (140') Reactor Vessel Thermal Shield Core Barrel FIGURE 1 ~ ARRANGEMENT OF SURVEILLANCE CAPSULES RT THE PRESSURE VESSEL

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TABLE I D0NALD C. C0OK UNn No. 1 REACT0R VESSEL SURVEn.LANCE MATERZALS<>>)

Heat Treatment Histor Shell Plate Material:

Heated to 1600 F for 4 hours4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months , water quenched.

Tempered at 1225 F for 4 hours4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months , air cooled.

Stress relieved at 1150 F for 40 .hours, furnace cooled.

Weldment:

Stress relieved at 1150 F. for 40 hours, furnace cooled.

Correlation Monitor:

1675 F, 4 hours4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months , air cooled.

1650 F, 4 hours4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months , water quenched.

1225 F, 4 hours, furnace cooled 1150 F, 40 hours4.62963e-4 days 0.0111 hours 6.613757e-5 weeks 1.522e-5 months , furnace cooled to 600 F.

Chemical Com osition (Percent)

Material C Mn P S Si Ni Mo Cu Plate B4406-3 0.24 1.40 0.009 0.015 0.25 0.49 0.46 0.14 Weld Metal 0.26 1.33 0.023 0.014 0.18 0.74 0.44 0.27 Correlation Monitor 0.22 1.48 0.012 0.018 0.25 0.68 0.52 0.14

with the simulated crack perpendicular to both the primary rolling direction and to the major plate surfaces. All mechanical. test specimens, see Figure 2, were taken at least one plate thickness from the quenched edges of the plate material.

Capsule T contained 44 Charpy V-notch specimens (10 longitudinal and 10 transverse from the plate material, plus 8 each from weld metal, HAZ and the reference steel plate); 4 tensile specimens (2 plate and 2 weld metal);

and 4 WOL specimens (2 plate and 2 weld metal). The specimen numbering sys-tem and location within Capsule T is shown in Figure 3.

Capsule T also was reported to contain the following dosimeters for de-termining 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 1 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 capsule) was 1.75% Ag, 0.75% Sn and 97.5% Pb having a melting point of 590 F.

10

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.009 90~

.3I l.063 .3 5

.3 I4 l.053 .393

2. I25 2.I05 (a) Charpy V-notch Impact Specimen I.005 Gage length

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. 256 .255 256

.246 .245 .395 I6 493

.250 R

.I98 I.250'.26 l.495 .I9 I. 80 4.250 4.2 I 0 630 .790

.786

.395

.375

'ECTION A- A D

.37 (b) Tens ile Spec imen l.45 l.4P

.375 D.

I.I30

.380 I.I20 I.005 .765

~ 995

.745

.439 499 .437 I.005

.995

.04'73

.0463 D .SOI

.0667 .499

.0662

.0667 (c ) Wedge Opening Loading Specimen FIGURE 2. VESSEL MATERIALSURVEILLANCE SPECIMENS

fC,COI CO.CCS CLttfC tttL ~ ISLICIII CLtt ffI Itl CLI ff IIL ILICI fC>COI CO CIS II I III

~ IIL OOL llISILC Cllltt Clutt Clllt1 CIOItl Clllt1 CllltT CClltl CllltT CILItt Cllltt Cltltt W.LI I IO I~ 'll I-jl I SI

~ IL4 ~ I4 I.l~ I SI 1 SI ~ SI 4 SI ~ SI ill ILO I~ II I ~ I I II I ILL LI II ~ ~ 'll I I4 ISS lit SISS SS Y-SS I.ll I.lt ~ St W IL ~ SS ~ .Sl I SS I.ll I.IS 1- ~ I~ I.IL I-~ I I LS -II LI.I ~ .II

~ TOP ItICLLC~ ~IIIIIICIOC BOTTOM tllllILLOI I (IIIII'ITOIIIL IIICIIII)

~ ~ I OILS OCII.IIIICLI~ IOIC II tllIC ~ 'LIOI.S (LIILSI(III~ IIICII4I) IIL~ IC ILL ISILL COIIILLIIOISaallla FIGURE 3. ARRANGEMENT OF SPECIMENS AND DOSIMETERS IN CAPSULE T

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IV. TESTING OF SPECIMENS FROM CAPSULE T The capsule shipment, capsule opening, specimen testing and reporting of results were carried out in accordance with the Project Plan for Donald C. Cook Unit No. 1 Reactor Vessel Irradiation Surveillance Program. The SwRI Nuclear Projects Operating Procedures called out in this plan include:

(1) XI-MS-1, "Determination of Specific Activity of Neutron Radiation Detector Specimen."

(2) XI-MS-3, "Conducting Tension Tests on Metallic Materials."

(3) XI-MS-4, "Charpy Impact Tests on Metallic Materials."

(4) XIII-MS-1, "Opening Radiation Surveillance Capsules and Handling and Storing Specimens."

(5) XI-MS-5, "Conducting Wedge-Opening-Loading Tests on Metallic Materials."

i (6) XI-MS-6, "Determination of Specific Activity of Neutron Radiation Fission Monitor Detector Specimens."

Copies of the above documents are on file at SwRI.

Southwest Research Institute utilized a procedure which had been pre-pared for the 1977 refuelling outage for the removal of Capsule T from the reactor vessel and the shipment of the capsule to the SwRI laboratories.

SwRI contracted with Todd Shipyards Nuclear Division to supply appropriate cutting tools and a licensed shipping cask. Todd personnel severed the cap-sule from its extension tube, sectioned the extension tube into three-foot lengths, supervised the loading of the capsule and extension tube materials into the shipping cask, and transported the cask to San Antonio.

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The capsule shell had been fabricated by making two long seam welds to join two half-shells together. The long seam welds were milled off on a Bridgeport vertical milling machine set up in one hot cell. Before mill-ing off the long seam weld beads, transverse saw cuts were made to remove the two capsule ends. After the long seam welds had been milled away, the top half of the capsule shell was removed. The specimens and spacer blocks were carefully removed and placed in an indexed receptacle so that capsule location was identifiable. After the disassembly had. been completed, the specimens were carefully checked for identification and location, as listed in WCAP 8047.(>>)

Each specimen was inspected for identification number, which was checked against the master list in WCAP 8047. No discrepancies were found.

The thermal monitors and dosimeter cfires were removed from the holes in the spacers. The thermal monitors, contained in quartz vials, were examined, and no evidence of melting was observed, thus indicating that the maximum temperature during exposure of Capsule T did not exceed 579 F.

The specific activities of the dosimeters were determined at SwRI with an NDC 2200 multichannel analyzer and an NaI(Th) 3 x 3 scintillation crystal. The calibration of the equipment was accomplished with appropri-ate standards and an interlaboratory cross check with two independent count-

'ing laboratories on Co-, 54Hn- and ~ Co-containing dosimeter wires. All activities were corrected to the time-of-removal (TOR) at reactor shutdown.

Infinitely dilute saturated activities (A8AT) were calculated for each of the dosimeters because ASAT is directly related to the product of the

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energy-dependent microscopic activation cross section and the neutron flux density. The relationship between ATOR and ASAT is given by:

ATOR (1-e -XTm m) (e

-Xtm)

E ASAT m~1 where: m = operating period; decay constant for the activation product, day 1; Tm equivalent operating days at 3250 MwTh for operat-ing period m; tm = decay time after operating period m, days.

The Donald C. Cook Unit No. 1 operating history up to the 1977 refuelling out-age is presented in Table II. The specific activity at time of removal (TOR) and the specific saturated activity calculated for each dosimeter are pre-sented in Table III.

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:

SAT m (2)

NOD where: energy-dependent neutron flux density, n/cm -sec; ASAT saturated activity, dps/mg target element; spectrum-averaged activation cross section, cm ;

NO number of target atoms per mg.

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

TABLE II

SUMMARY

OF REACTOR OPERATIONS DONALD C. COOK UNIT NO. 1 Power Equiualent Decay Time Operating Dates Operating Shutdown Generation Operating Days After Period Period Start ~DS S ~DS s ~DS S T )

2/2/75 2/14/75 13 2,194 0. 68 678 2/15/75 2/16/75 2/17/75 2/17/75 228 0. 07 675 2/18/75 2/20/75 2/21/75 3/18/75 26 29,604 9.11 646 3/19/75 4/3/75 16 4/4/75 6/24/75 82 200,616 61. l3 548 6/25/75 6/26/75 6/27/75 7/3/75 15,432 4. 75 539 7/4/75 7/22/75 19 7/23/75 10/ll/75 81 201,506 62.00 439 10/12/75 10/14/75 10/15/75 10/31/75 17 40,163 12.35 419 ll/1/75 11/14/75 14 11/15/75 1/1/76 48 116,552 35.86 357 1/2/76 1/4/76 1/5/76 4/12/76 256,178 78.82 255 4/13/76 5/9/76 27 10 5/10/76 7/1/76 53 143,868 44.27 175 7/2/76 7/5/76 7/6/76 9/10/76 67 205,682 63.29 104 9/11/76 9/18/76 12 9/19/76 11/20/76 63 196,520 60.47 33 11/21/76 11/21/76 11/22/76 12/23/76 32 92 754 28.54 0 Total, Cycle 1 1,501,297 461.94

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TABLE III

SUMMARY

OF NEUTRON DOSIMETRY RESULTS DONALD C. COOK UNIT NO. 1--CAPSULE T Monitor Activation- ATOR ASAT Identification Reaction (d s/m d s/m Fe- Top 54Fe(n,p)54Mn 193 x 103 3.34 x 103 Fe Top Mid. 1.69 x 103 2.94 x 103 Fe- Mid. 1.69 x 103 2.93 x 103 Fe Bot. Mid. 1.69 x 103 2.93 x 103 Fe Bot. 1.80 x 103 3.11 x 103.

Average 1.76 x 103 3.05 x 103 Cu Top Mid. 63Cu(n,a)60Co 5.14 x 101 3.43 x 102 Cu Mid.

ll 5.27 x 101 3.52 x 102 tf x 4.03 x 102 Cu Bot. Mid. 6.04 101 Ni Top Mid. 58Ni(n,p)58Co 3.83 x 104 4.46 x 104 If 4.38 x 104 Ni Mid. 3.77 x 104 II 3.95 x 104 4.59 x 104 Ni Bot. Mid.

Co - Top Co(n,p) Co 4.87 x 106 3.25 x 107 II x 1.22 x 107 Co(Cd) Top 1.83 106 If 5.03 x 106 3.36 x 107 Co Bot.

Co (Cd) Bo t. II 1.64 x 106 1.09 x 107 U-238 238U(n, f) 137C 1.20 x 103 N/A Np-237 237Np(n, f)137Cs 4.53 x 103 N/A 17

The neutron flux density was calculated from the 4Fe(n,p) 4Mn reac-tion because it has a high energy threshold and the energy response is well known. The energy spectrum for Capsule T was calculated with the DOT 3.5 two-dimensional discrete ordinates transport code with a 22-group neutron cross section library, a Pl expansion of the scattering matrix and an S8 order of angular quadrature. The normalized spectrum for Capsule T and the group-organized cross sections for the 54Fe(n,p)54Mn reaction derived from the ENDF/B-ZV library are given in Table IV. The value of o Fe is given by:

10 MeV aF (E)g(E)dE o (> 1 Mev) - 1'1 (3) 10

$ (E)dE

l. 00 where: VF Fe

(> 1 MeV) the calculated spectrum-averaged cross section for flux > 1 MeV, cm2 determined for the 54Fe(n,p)54Mn reaction.

The resulting value obtained for fast (> 1 MeV) neutron flux density at the Capsule T location was 4.50 x 101 neutrons/cm -sec. Since Donald C. Cook Unit No. 1 operated for an equivalent 461.94 full power days up to the 1977 refuelling outage, the total neutron fluence for Capsule T is equal to 1.80 x 1018 neutrons/cm 2 (E > 1 MeV) based on the calculated spectrum at the cap-sule location.

Assuming a fission-spectrum energy distribution at the capsule location, the cross-section for the 4Fe(n,p) 4Mn reaction (E > 1 MeV) would be 98.26 mb. (4) The resulting flux and fluence values would be 4.95 x 10 neu-trons/cm2-sec and 1.97 x 1018 neutrons/cm2, respectively.

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TABLE IV FAST NEUTRON SPECTRUM AND IRON ACTIVATION CROSS SECTIONS FOR CAPSULE T 54Fe(n,p)54Mn Energy Range Normalized Cross Section (MeV) Neutron Flux (barns)

8. 18 10. 0 0.0098 0.581 6.36 8.18 0.0254 9.577 4.96 6.36 0.0482 0.491 4.06 4.96 0.0471 0.354 3.01 4.06 0.0855 0.205 2.35 - 3.01 0.1400 0.099 1.83 2.35 0. 1752 0.023 1 11 1.83

~ 0.4689 0.0014 VF 0.108 barns Fe 19

The irradiated Charpy V-notch specimens were tested on a SATEC impact machine. The test temperatures were selected to develop the ductile-brittle transition and upper shelf regions. The unirradiated Charpy V-notch impact data reported by Westinghouse(13) and the data obtained by SwRI on the spec-imens contained in Capsule T are presented in Tables V through IX. The Charpy V-notch transition curves for the three plate materials and the cor-relation monitor material are presented in Figures 4 through S. The radia-,

tion-induced shift in transition temperatures for the vessel plates are in-dicated at 50 ft<<lb and 35 mil lateral expansion. A summary of the shifts in RTNDT and Cv upper shelf energies for each material are presented in Table X.

Tensile tests were carried out in the SwRI hot cells using a Dillon 10,000-1b capacity tester equipped with a strain gage extensometer, load cell and autographic recording equipment. One each plate and weld metal tensile specimens was tested at room temperature (RT) and at 550 F. The results, along with tensile data reported by Westinghouse on the unirradi-ated materials(1 ), are presented in Table XI. The load-strain records are included in Appendix A.

Testing of the WOL specimens was deferred at the request of American Electric Power Service Corporation. The specimens are in storage at the SwRI radiation laboratory.

The Charpy V-notch results indicate that the HAZ is more sensitive to radiation embrittlement than the as-rolled and heat-created plate and about equal to that of the weld metal. This is surprising because the copper con-tent of HAZ is reported to be'uch lower than that of the weld metal.( 3) 20

TABLE V CHARPY V-NOTCH IMPACT DATA THE DONALD C. COOK UNIT NO. 1 REACTOR PRESSURE VESSEL INTERMEDIATE SHELL PLATE B4406-3 (LONGITUDINAL DIRECTION)

Test Impact Lateral Spec. Temp. Energy Shear Expansion Condition No. ( p) (ft-1b) (x) ~Mls Baseline (a) -40 10 13

-40

-40 ll 11.5 10 11 10 24.5 9 24 10 33 11 29 10 31.5 13 28 40 57 23 49 40 42 25 40 40 65 29 54 76 82 45 67 76 70 37 60 76 78 37 61 110 93.5 52 72 110 100 59 77 110 88 52 72 160 110 95 84 160 131.5 100 95 160 115.5 95 83 210 120 100 89 210 144 100 98 210 125 100 95 300 131.5 100 90 300 126 100 92 300 132 100 93 Capsule T A-44 10 10.5 1 10 A-45 40 29 5 24 A-49 82 38 20 31 A-50 110 46.5 35 38 A-41 135 62.5 25 53 A-47 160 84 55 58 A-42 185 99 95 80 A-48 210 105 95 83 A-43 250 110 100 89 A-46 300 105.5 100 89 (a) Not reported.

21

TABLE VI CHARPY V-NOTCH IMPACT DATA THE DONALD C. COOK UNIT NO. 1 REACTOR PRESSURE VESSEL INTERMEDIATE SHELL PLATE B4406-3 (TRANSVERSE DIRECTION)

Test Impact Lateral Spec. Tempt Energy Shear Expansion Condition Na. ~P) ~ft-1b) ~7.) ~milt Baseline (a) -40 11 12

-40 11.5 15

-40 14 15 10 28 14 28 10 23 9 22 10 30 9 26 40 40 18 36 40 41 23 35 40 37 18 34 76 83 27 56 76 43 27 44 76 50 32 46 76 50 27 44 110 84 48 71 110 54 37 51 110 68 41 57 160 97 90 80 160 77 90 71 210 90 100 75 210 95 100 79 210 97 100 79 300 100 100 83 300 94 100 75 300 101 100 85 Capsule T AT-44 1O 6 5 8 AT-45 40 25 5 23 AT-49 82 35 20 30 AT-50 110 37 30 35 AT-41 135 49.5 25 44 AT-47 160 57 40 47 AT-42 185 73.5 100 63 AT-48 210 87 100 73 AT-43 250 87 100 71 AT-46 300 89 lOO 83 (a) Not reported.

22

TABLE VII CHARPY V-NOTCH IMPACT DATA THE DONALD C. COOK UNIT NO. 1 REACTOR PRESSURE VESSEL CORE REGION WELD METAL Test Impact Lateral Spec. Temps Energy Shear Expansion Condition No. ('p) ~ft-1b1 (X) ~m11s Baseline (a) -140

-140 ll 21 10 19

-140 19 18

-100 23.5 18 22

-100 29 20 26

-100 20 11 18

-70 45.5 24 39

-70 51 42 47

-70 54 32 49

-40 63 47 52

-40 59 34 53

-40 69 47 60 10 83 73 69 10 84 71 72 10 92 75 75 76 114 99 88 76 107 100 87 76 107 100 88 210 110 100 90 210 112 100 87 210 111 100 93 Capsule T W-33'-35

>>40 24. 5 5 19 10 50 20 41 W-34 75 75.5 70 67 W-39 82 44 20 34 W-40 110 85 95 69 W-37 160 75 100 66 W-38 210 98 100 66 W-36 300 68.5 100 66 (a) Not reported.

23

TABLE VIII CHARPY V-NOTCH IMPACT DATA THE DONALD CD COOK UNIT NO. 1 REACTOR PRESSURE VESSEL CORE REGION MELD HEAT-AFFECTED ZONE METAL Test Impact Lateral Spec. Tempo Energy Shear Expansion Condition No. ~7) (ft-lb) ~(/ ~mals Baseline (a) -175 5.5

-175 7

-140 16 12

-140 22 18

-100 30 13 25

-100 33 14 28

-100 45 20 40

-70 52 21 39

-70 47 25 35

-70 27 14 21

-70 30 20 24

-40 54 55 53

-40 71 50 50

-40 47 43 45 10 97 90 83 10 89 43 67 10 82 69 64 76 112 100 86 76 '40 100 84 76 131 100 82 210 129 100 85 210 104 100 94 210 105 100 87 Capsule T H-33 -40 10 5 9 H-35 10 40.5 15 30 H-34 45 30.5 25 27 H-39 82 52.5 25 41 H-40 110 62.5 40 46 H-37 160 84 100 65 H-38 210 111.5 100 78 H-36 300 83 100 54 (a) Not reported.

TABLE IX CHARPY V-NOTCH EPACT DATA A533 GRADE B CLASS 1 CORRELATION MONITOR MATERIAL Test Impact Lateral Spec. Tempr Energy Shear Expansion Condition No. ~fe-1b) ~X ~mals)

Baseline (a) -50

-50

-20 6.5 9 6

-20 9 13 10

-20 6 13 9 10 12 23 15 10 14.5 23 14 10 13.5 23 14 40 22 33 23 40 36 29 32 40 35 29 32 85 58.5 43 51 85 41.5 41 42 85 52 42 45 110 82.5 58 60 110 85.5 67 71 110 63.5 55 54 160 108.5 84 72 160 81 85 69 160 109 87 79 210 117 98 84 210 115 98 88 210 121 100 87 300 125 100 87 300 117.5 100 83 300 127 100 84 Capsule T R-33 40 13.5 5 13 R-37 82 18.5 10 18 R-38 110 35 20 32 R-39 160 55.5 40 45 R-40 210 86.5 95 66 R-34 300 100 100 57 R-35 350 111 100 84 R-36 400 96.5 100 84 (a) Not reported.

25

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-200 -100 0 100 200 300 400 Temperature, deg F FIGURE 4 ~ CHARPY V-NOTCH PROPERTIES OF PLATE B4406-3 (LONG-)

DONALD C. COOK UNIT NO. 1 SURVEILLANCE PROGRAM 26

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-200 -100 0 100 200 300 400 Temperature, deg F FIGURE 5. CHARPY V-NOTCH PROPERTIES OF PLATE B4406-3 (TRANS.)

DONALD C. COOK UNIT NO. 1 SURVEILLANCE PROGRAM 27

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-200 -100 0 100 200 300 400 Temperature, deg F FIGURE 6. CHARPY V-NOTCH PROPERTIES OF CORE REGION WELD METAL DONALD C. COOK UNIT NO. 1 SURVEILLANCE PROGRAM 28

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-200 -100 0 100 200 300 400 Temperature, deg F FIGURE 7. CHARPY V-NOTCH PROPERTIES OF CORE REGION HAZ MATERIAL DONALD C. COOK UNIT NO. 1 SURVEILLANCE PROGRAM 29

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-200 -100 0 100 200 300 400 Temperature de@ F FIGURE 8. CHARPY V-NOTCH PROPERTIES OF CORRELATION MONITOR MATERIAL DONALD C. COOK UNIT NO. 1 SURVEILLANCE PROGRAM 30
TABLE X NOTCH TOUGHNESS PROPERTIES OF CAPSULE T SPECIMENS DONALD C. COOK UNIT NO. 1 Plate B4406-3 Weld Weld Correlation
~(Lan .) (Trans.) Metal HAZ Monitor 50 ft-lb C Tem . (de F)
Irradiated 150(a) 140 60 70 145 Unirradiated 75(a) 65 -70 -60 75 AT 75( ) 130 130 70 35 mil C Tem . (de F)
Irradiated 135 (b) 110 50 55 125 Unirradiated 60(b) 40 -80 -75 60 AT 75(b) 130 130 65 C U er Shelf Ener ft-lb)
Unirradiated 130 94 110 120 120 Irradiated 108 84 80 93 102 hE, ft-lbs 22 10 30 27 18 AE, 16.9 10. 6 27.3 22.5 15 (a) Energy transition at 77 ft-lb.
(b) Lateral expansion transition at 54 mil.
31
TABLE XI TENSII E PROPERTIES OP SURVEILLANCE MATERIALS CAPSULE T Test 0.2X Yield Tensile Total Reduction Condition Specimen Ident.
Temp.
('P) ~si Strength Strength
~sf ~I Elongation in Area
(%)
Baseline B4406-3 Room 68,650 90,650 27.7 70.4 (Long.) Room 68,250 90,250 27.4 69.6 300 61,350 82,650 23.4 69.4 300 61,200 82,300 22.6 69.7 600 58,000 87,000 26.0 65.1 600 58,550 87,400 25.4 67.0 Capsule T A-1 Room 72,700 99,800 24. 3 65.7 A-2 550 66,700 93,000 20. 2 64.3 Baseline B4406-3 Room 68,700 90,300 26.6 65.8 (Trans.) Room 67,600 89,450 25.6 65.0 300 61,000 82,800 23.0 65.0 300 60,900 81,900 23.3 64.6 600 58,300 86,000 24.8 58.8 600 55,900 86,600 24.7 58.6 Baseline Veld Metal Room 66,900 81,500 28.7 73.2 Room 67,350 82,250 25.0 65.3 300 '9,700 74,600 24.0 72.9 300 59,800 74,500 23.3 71.8 600 57,200 79,400 23.4 65.2 600 56,300 78,500 23.6 63.4 Capsule T W-9 Room 86,100 103,400 23.6 65.0 M-10 550 75,800 95,300 19.3 60.8 32
~
~
The tensile properties of the weld metal appeared to be the most af-fected by the radiation exposure in Capsule T as expected from .the reported copper contents.
33
' ~
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 10CFR50. This requires a projection of the measured reduction in C 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. 1 was calculated for Capsule T with the DOT 3.5 discrete ordi-nates transport code. The lead factor for Capsule T reported by Westinghouse is 2.6 for the vessel I.D. surface.( ) This was supported by the SwRI DOT 3.5 analysis. The DOT 3.5 analysis also predicted that the fast flux at the 1/4T and 3/4T positions in the 8-5/8-in. pressure vessel wall would be 49% and 7.8%,
respectively, of that at the vessel I.D. These figures are in good agreement with fluence attenuation determinations of 46% and 10% for an 8-in. steel plate by the Naval Research Laboratory.( ) However, currently the NRC pre-fers to use more conservative figures of 60% and 15%, respectively, for the attenuation of fast neutron flux at the 1/4T and 3/4T positions in an 8-in.
vessel wall. (16) This conservatism allows for the increased fraction of neutrons which might accrue in the 0.1 to 1.0 MeV range in deep penetra-tion situations. For the 8-5/8-in. wall thickness of the D.C. Cook Unit No. 1 vessel, the attenuations become 57% and 12.5% for the 1/4T and 3/4T positions, respectively.
A method for estimating the reduction in Cv upper shelf energy as a function of neutron fluence is given in Regulatory Guide 1.99, Revision
.(7))
1.( The results from Capsule T are compared to a portion of Figure 2 of Regulatory Guide'.99, Revision 1, in Figure 9. The embrittlement response of the weld metal, reported to contain 0.27% Cu( ), is in good agreement with the prediction of Regulatory Guide 1.99, Revision 1. However, the plate is less sensitive and the HAZ is more sensitive than predicted for the 0.14% copper content. The behavior of the HAZ specimens may reflect some copper pickup in the HAZ from the weld deposit or the placement of the notch unusually close to the fusion line. Using the dashed curve drawn through the data point for the weld metal, it is predicted that the weld metal Cv shelf energy will reach 50 ft-lbs at a fluence of about 2.1 x 10 (E > 1 MeV). This corresponds to approximately 38 effective full power years (EFPY) of operation at the vessel I.D., in excess of the 32 EFPY design life of the plant. The plate and HAZ materials are projected to require even larger fluences to reach the 50 ft-lb shelf level. These projections will be reex-amined after the next surveillance capsule has been removed.
A similar approach can be taken to estimate the increase in RTHDT as a function of reactor power generation. Figure 10 compares the Donald'. Cook Unit No. 1 surveillance data on the three surveillance materials to selected portions of Figure 1 of Regulatory Guide 1.99, Revision 1. The results
60 s I Is sr :lli I!!! s Isl s ' I ~ Is l sl I'I s I s I ~
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2 x 1017 4 6 8 1018 2 4 6 8 1019 2 4 Neutron Fluence, n/cm (E > 1 MeV)
FIGURE 9. DEPENDENCE OF Cv SHELF ENERGY ON NEUTRON FLUENCE, DONALD C. COOK UNIT NO. 1
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FIGURE 1 EFFECT OF NEUTRON FLUENCE ON RTNDT SIIIFT>> DONALD C COOk UNIT NO'
indicate that the measured shift in RTNDT of the weld metal is in agreement with that predicted by Regulatory Guide 1.99, Revision 1, but that the mea-sured shifts in RTNDT for the plate and HAZ materials are underpredicted by the guide.
The predicted shifts in RTNDT for the Donald C. Cook Unit No. 1 reac-tor pressure vessel obtained from Figure 10 are summarized in Tables XII and XIII. The values predicted at the 1/4T and 3/4T after 12 EFPY (Table XII) are used to develop heatup and cooldown limit curves to meet the require-ments of Appendix G to Section III of the ASME Code, as described in Section VI of this report. These projections for Cv shelf energy reductions and RTNDT shifts, and the resulting heatup and cooldown limit curves, are based on extrapolations from one data point representing the most sensitive material.
After a second capsule has been removed and tested, one will be able to inter-polate between two data points.
The Donald C. Cook Unit No. 1 reactor vessel surveillance program sched-ule proposed by Westinghouse~ ~ is summarized in Table XIV. It has been or-ganized to satisfy Appendix H of 10CFR50 as closely as possible. There are seven additional capsules in the vessel, all of which contain base plate, weld metal and HAZ specimens. There is no reason to consider changing the proposed capsule removal schedule at this time.
39
TABLE XII PROJECTED VALUES OF RTNDT FOR DONALD C. COOK UNIT NO. 1 FOR UP TO 12 EFPY OF OPERATION Calculated Fluence RT (de F)
Location Material (n/cd E > 1 MeV) Initial Shift 12 EFPY(a )
Vessel I.D.
~ ~ Inter. Shell Plate 6.55 x 1018 45(b) 145 190 Weld Metal -52(b) 245 193 HAZ -60(c) 245 185 Vessel 1/4T Inter. Shell Plate 3. 73 x 1018 45(b) 110 155 Weld Metal 52(b) 185 133 WZ -60(c) 185 125 Vessel 3/4T Inter. Shell Plate 45(b) 50 95 Weld Metal 52(b) 87 35 MZ -60(c): 87 27 (a) 1 EFPY 1,186,250 M&t.
(b) Reference 18.
(c) References 13 and 18.
TABLE XIII PROJECTED VALUES OF RTNDT FOR DONALD C. COOK UNIT NO. 1 FOR UP TO 32 EFPY OF OPERATION Calculated Fluence R DT (de F)
Location Material (n/cm2 E > 1 MeV) Initial Shift 32 EFPY(a )
Inter. Shell Plate 45(b) 240 285 Meld Metal -52(b) 320 268 HAZ -60(c) 320 260 Vessel 1/4T Inter. Shell Plate '1.0 x 1019 45(b) 180 225 lfeld Metal -52(b) 285 233 HAZ -60(c) 285 225 Vessel 3/4T Inter. Shell Plate 2.2 x 1018 45(b) 83 128 Meld Metal -52(b) 142 90 HAZ 60(c) 142 82 (a) 1 EFPY = 1,186,250 MMDt.
(b) Reference 18.
(c) References 13 and 18.
TABLE XIV PROPOSED REACTOR VESSEL SURVEILLANCE CAPSULE SCHEDULE DONALD C. COOK UNIT NO. 1 Capsule Lead Identification Factor Removal Time 2.6 Removed and tested at end of first core cycle 2.6 10 Years (postirradiation test) 0.6 10 Years (reinsert in Capsule T location) 0.6 10 Years (reinsert in Capsule X location) 2.6 20 Years (postirradiation test) 0.6 20 Years (reinsert in Capsule U location) 2.6 30 Years (postirradiation test) 0.6 30 Years (reinsert in Capsule Y location)
~ ~
VI. HEATUP AND COOLDOMN LIMIT CURVES FOR NORMAL OPERATION OF DONALD C. COOK UNIT NO. 1 Donald C. Cook Unit No. 1 is a 3250 Mwt pressurized water reactor oper-ated by American Electric Power Service Corporation. The unit has been pro-vided with a reactor vessel material surveillance program as required by 10CFR50, Appendix H.
The first surveillance capsule (Capsule T) was removed during the 1977 refuelling outage. This capsule was tested by Southwest Research Institute, the results being described in the earlier sections of this report. In sum-mary, these results indicate that:
(1) The RTNDT of the surveillance materials in Capsule T increased a maximum of 130 F as a result of exposure to a neutron fluence of 1.80 x 10 neutrons/cm2 (E > 1 MeV).
(2) Based on a ratio of 2.6 between the fast neutron flux at the Capsule T location and the maximum incident on the vessel wall, the vessel wall fluence at the I.D. was 6.92 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 be 185 F at the 1/4T and 87 F at the 3/4T vessel wall locations, as controlled by the weld metal and HAZ materials.
(4) The intermediate shell plate material, although less sensitive to radiation embrittlement than the weld and HAZ materials, is projected to control the limiting RTNDT for a considerable length of time because of a much higher initial (unirradiated) RTNDT of 45 F.(
43
~ ~
The Unit No. 1 heatup and cooldown limit curves for 12 EFPY have been computed on the basis of (4) above because it is anticipated that the RTNDT of the primary pressure boundary materials will be highest for the plate ma-terial at least through that time period (see Table XII). The procedures employed by SwRI are described in Appendix B.
The following pressure vessel constants were employed as input data in this analysis:
Vessel Inner Radius, ri 86.50 in., including cladding Vessel Outer Radius, ro 95.34 in.
Operating Pressure, Po 2235 psig Initial Temperature, To 70 F Final Temperature, Tf 550'F Effective Coolant Flow Rate, Q ~ 135.6 x 10 ibm/hr Effective Flow Area, A 26.72 ft2 Effective Hydraulic Diameter, D ~ 15.05 in.
Heatup curves were computed for a heatup rate of 60 F/hr. Since lower rates tend to raise the curve in the central region (see Appendix B), these curves apply to all heating rates up to 60 F/hr. Cooldown curves were com-puted for cooldown rates of 0 F/hr (steady state), 20 F/hr, 40 F/hr, 60 F/hr, and 100 F/hr. The 20 F/hr curve would apply to cooldown rates up to 20 F/hr; the 40 F/hr curve would apply to rates from 20 F to 40 F/hr; the 60 F/hr curve would apply to rates from 40 F to 60 F/hr; the 100 F/hr curve would apply to rates from 60 F/hr to 100 F.hr.
The Unit No. 1 heatup and cooldown curves for up to 12 EFPY are given in Figures ll and 12.
44
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60 100 150 200 250 300 350 400 Indicated Temperature, deg F FIGURE ll. DONALD C. COOK UNIT NO. 1 REACTOR COOLANT HEATUP LIMITATIONS APPLICABLE FOR PERIODS UP TO 12 EFFECTIVE FULL POfKR YEARS
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60 100 150 200 250 300 350 400 Indicated Temperature, deg F FIGURE 12. DONALD C. COOK UNIT NO. 1 REACTOR COOLANT COOLDOWN LIHITATIONS APPLICABLE FOR PERIODS UP TO 12 EFFECTIVE FULL POWER YEARS
VII. REFERENCES
1. Title 10, Code of Federal Regulations, Part 50, "Licensing of Produc-tion and Utilization Facilities."

2. ASME Boiler and Pressure Vessel Code, Section III, "Nuclear Power Plant Components," 1974 Edition.

3. ASTM E 208-69, "Standard Method for Conducting Drop-Weight Test to De-termine Nil-Ductility Transition Temperature of Ferritic Steels," 1975 Annual Book of ASTM Standards.
Steele, L. E., and Serpan, C. Z., Jr., "Analysis of Reactor Vessel Radiation Effects Surveillance Programs," ASTM STP 481, December 1970.
5. Steele, L. E., "Neutron Irradiation Embrittlement of Reactor Pressure Vessel Steels," International Atomic Energy Agency, Technical Reports Series No. 163, 1975.

6. ASME Boiler and Pressure Vessel Code, Section XI, "Rules for Inservice Inspection of Nuclear Power Plant Components," 1974 Edition.

7. Regulatory Guide 1.99, Revision 1, Office of Standards Development, U.S. Nuclear Regulatory Commission, April 1977.

8. Comments on Regulatory Guide 1.99, Westinghouse Electric from NRC Public Document Room, Washington, D.C.
Corporation,'btained
9. Position on Regulatory Guide 1.99, Combustion Engineering Power Sys-tems, Obtained from NRC Public Document Room, Washington, D.C.

10. ASTM E 185-73, "Standard Recommended Practice for Surveillance Tests for Nuclear Reactor Vessels," 1975 Annual Book of ASTM Standards.

11. ASTM E 399-74, "Standard Method of Test for Plane-Strain Fracture Toughness of Metallic Materials," 1975 Annual Book of ASTM Standards.

12. Witt, F. J., and Mager, T. R., "A Procedure for Determining Bounding Values of Fracture Toughness KIc at Any Temperature," ORNL-TM-3894, October 1972.

13. "American Electric Power Service Corporation Donald C. Cook Unit No. 1 Reactor Vessel Radiation Surveillance Program," WCAP-8047, March 1973.

14. ENDF/B-IV, Dosimetry Tape 412, Mat No. 6417 (26-Fe-54), July 1974.

15. Loss, F. J., Hawthorne, J. R., Serpan, C. Z., Jr., and Puzak, P. P.,
"Analysis of Radiation-Induced Embrittlement Gradients on Fracture Characteristics of Thick-Walled Pressure Vessel Steels," NRL Report 7209, March 1, 1971.
47
16. Telecon, E. B. Norris to Ken Hogue (NRC Staff) January 19, 1977.

17. Hazleton, W. S., Anderson, S. L., and Yanichko, S. E., "Basis for Heatup and Cooldown Limit Curves," WCAP-7924, July 1972.

18. Donald C. Cook Unit No. 1 Technical Specifications, as of November 30, 1977.
48
APP END IX A TENSILE TEST RECORDS
Southwest Research Institute Department of Materials Sciences TENSILE TEST DATA SHEET Test No. T- .. l Est. U. T.S. PS1 Spec. No. -1 Initial G. L. r 41Z1 ~ Machine No. Temperature I4 'F Initial Dia.. I in. Date 77 ts J Strain Rate, < 2 tzpi> Inisial Thickness in. Initial Area Initial Width in. Top Temperature Maximum Load 40 lb Bottom Temperature 'F 0. 2'%ffset Load 88 9 D lb Final Gage Length p 4T ine 0.02% Offset Load lb Final Diameter /~~I in. Upper Yield Point Final Area ine 2 r Maximum Load Initial Area psi cjoy P 2 Init1al Area 2- ~g psi
0. 02% Offset Load 0 02/ Y S Initial Area PS1 UPPer Y S .. Upper Yield Point I tial Area PS1 Final G. L. - Initial x 100=
Initial Area - Final Area 1 p @~7 Initial Area Signature: A-2
-0;0 rZi9ahJ A-3
Southwest Research Institute Department of Materials Sciences TENSILE TEST DATA SHEET Test No. T- . Z Est. U. T. S. psi Spec. No. Initial G. L. .O in. Tem per afore ~P'F Initial Dia. . g C 'n. Strain Rate . C'~/Wrr/ Initial Thickness in. Initial Area . +H / Initial VTidth in. I Tap Temperature 'F Maximum Load S~ 7S lb Bottom Temperature 0 2%%uo Offset Load 5 2.=.~~ lb Final Gage Length 0.02%%utf Offset Load lb Final Diameter . l+J ln ~ Upper Yield Point lb Final Area . o '72 2 r Initial Area
0. 2% Offset Load Initial Area
~ 02%%u'ff et Load 0 02%%u Y S Initial Area psl U er Yield Point PPer .-
I tlal Area Final G. L. - Initial x 100'= ~ ~'
%%utl Elongation %%uo Initial Area - Final Area 100 Initial Area tt
~ ~ ) '0' a
0 g~< A-5
Southwest Research Institute Department of Materials Sciences TENSILE TEST DATA SHEET Test No. T- Est. U. T. S. psi Spec. No. Initial G. L. dd in. Machine No. )> / J~~ Temperature >+ 'F Initial Dia. Date Initial T hie kne s s in. Initial Are a '~8 7 Initial Width in. Top Temperature oF Maximum Load 5.> G lb 0.2% Offset Load ~~n,~ > lb Final Gage Length 111 ~ 0.02% Offset Load lb ~ s Final Diamete" in. Upper Yield Point lb Final Area sP/ 74 +m. 2 Maximum Load 0.2'ls Offset Load Initial Area g~ gg.
= 2/o Offset Load 0 02$ Y S Initial Area ps1 pp, ' er Yield Point ~tel Area p81 u Fin G. L. - Initial G. L. % Elongation Initial G, L.
R A Initial Area - Final Area
% Initial Area Signature:
A-6
'A-7 1
~ ~
Southwest Research Institute Department of Materials Sciences TENSILE TEST DATA SHEET Test No. T- Est. U. T.S. psi Project No. 6<-a>>n -of" / Spec. No. Initial G. L. Machine No. Temperature 5 ft < 'F Initial Dia. Date Strain Rate Initial Thickness Initial Area . OHg'7 Initial Width 1ne Top Temperature 5 ~l ~ 'F Maximum Load + C~~'0 lb Bottom Tempe ratur e o 84 0. 2% Offset Load ~?~. 5 ib 0.02% Offset Load lb in. Upper Yield Point Final Area Maximum Load Initial Area 0.2%%uo Offset Load Initial Area p 02 Y S
0. 02 % 0 ffs e t Lo a d ps1
%%u Initxal Area U er Yield Point Initial Area ps'inal G. L. - Initial G L x lpp 0 E OIlgation ~.+ l G L // 7' Initial Area - Final Area Initial Area Signature:
A-8 b,t,
A-9 ~ 1 )1 APPENDIX B PROCEDURE FOR THE GENERATION OF ALLOWABLE PRESSURE-TEMPERATURE LIMIT CURVES FOR NUCLEAR POWER PLANT REACTOR VESSELS
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 cooldown operations, and core operation of reactor pressure vessels ~ The safety margins employed in these procedures equal or exceed those recommended in the ASME Boiler and Pressure Vessel Code, Section III, Appendix G, "Protection Against Nonductile Failure. " B. Back round The basic parameter used to determine safe vessel operational conditions is the stress intensity factor, KZ, which is a function of the stress state and flaw configuration. The KI corresponding to membrane tension is given by KI 'm'm where Mm is the membrane stress correction factor for the postulated flaw and o.m the membrane stress. Likewise, KI corresponding to bend-ing is given by KIb 'b 0'b (2) where Mb is the bending stress correction factor and o.b is the bending stress. For vessel section thickness of 4 to 12 inches, the maximum B-2
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 l. 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. 1). These curves are a function of the stress ratio parameter r/r, where o. (Py is the material yield strength which is, taken to be 50,000 psi. The bending correction factor is defined as 2l3 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 XI, "Rules for Inservice Inspection of Nu-clear Power Plant Components," Article A-3000. The Code specifies the minimum KI that can cause failure as a func-tion of material temperature, T, and its reference nil ductility temperature, RTNDT. This minimum KI is defined as the reference stress intensity fac-tor, KIR, and is given by KIR = 26777. + 1223. exp 0. 014493(T - RT NDT + 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-Tem erature Relationshi s
1. Inservice Leak and H drostatic Test During performance of inservice leak and hydrostatic tests, the reference stress intensity factor, KIR, must always be greater than B-3
3.8 MEh<8RAHQ I( m M ~ra 1.0 Im m Mb x 0.7
< 2/3hlm, 0.5 Mb O. I 3.2 3.0 E
2.~i 2.2 2.0 1.6 1.2 1.0 1.0 1.2 I A 1.6 1,0 2.0 2.2 2.~i 2.6 2.8 3.0 3. 2 3A 3.6 3.8 4.0 FIGURE 1. STRESS CORRECTION FACTOR
I70 I'R 26 777) V'IIERE l30 'EFEAFHCE STRESS INTENSITY FACTOR I20 TEhIPERATURE AT VIHICH I'IR IS PERhIITTED,'F II 0 RTHPT 'EFERFHCE HIL-DUCTILITY Lg tco TEMPERATURE SO 70 60 50 40 IO 0
-240 -200 -IGO -I20 -eO -40 0 40 80 .I20 IGO 200 240 TEIAPERATUAE RELATIYE TO ATHP,(T-ATHPT), FAHREIIHEI D GREES FIGURE 2. REFERENCE STRESS INTENSITY FACTOR B-5
l. 5 times the KZ caused. by pressure, thus

l. 5 Kl'p KZR (4) or
'5 Mm <m ~ K1R (5)
For a cylinder with inner radius ri and outer radius ro, the stress distribution due to internal pressure is given by With 1/4T flaws possible at both inner and outer radial locations, i. e., at rl/4 = ri 4 1/4(ro- ri) and r3/< r j+ 3/4(ro- ri), the maximum stress will occur at the inner flaw location, thus I r j.2 r + (1/4ro+ 4 3/4ri) o. max =P o ro2 - ri2 (1/4ro p 3/4r i)2 With the operation pressure known, i. e., Po, we deter-mine the minimum coolant temperature that will satisfy Equation (4) by e valuating KlR = '5Mm<max and determine the corresponding coolant temperature, T, from Equa-tion (3) for the given RT~~DT at the 1/4T location. For this calculation, Equation (3) takes the form I-*I- 6 ..6 .I [-666 ']. S-6
The inservice curves are generated for an operating pres-sure range of 96 Po to
~ l. 14 Po, where Po is the design operating pres sur e.
2. Heatu and C ooldown 0 e rations At all times during heatup and cooldown operations, the ref-erence stress intensity factor, K1R, must always be greater than the sum of 2 times the Klp caused by pressure and the Klt caused by thermal gra-dients, thus

2. 0 Klp + l. 0 Klt < KZR (10) or 2 0 Mm 0 max K1R - KZt where o max is the maximum allowable stress due to internal pressure, and KZt is the equivalent linear stress intensity factor produced by the thermal gradients. To obtain the equivalent linear stress intensity fac-tor due to thermal gradients requires a detailed thermal stress analysis.
The details of the required analysis are given in Section D. During heatup the radial stress distributions due to internal pressure and thermal gradients are shown schematically in Figure 3a. Assuming a possible flaw at the 1/4T location, 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 would represent the maximum stress condition at the 1/4T location. At
OUTER RAD IUS 3/4T Z/4T INNER RAD IUS Pressure stress distribution Thermal stress distribution ( a ) Heatup OUTER RAD IUS 3/4T 1/4T INNER RADIUS Pressure stress distribution Thermal stress distribution ( b ) Cooldown Figure 3. Heatup and Cooldown Stress Distribution B-8
the 3/4T flaw location, the pressure stress and 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 1/4T 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 K1R max( rj (12) ro + (1 /4ro 0 3/4r;) 2Mm roZ - rj (1/4ro+3/4rj)2 where Mm is determined from the curves in Figure 1 and K1R is obtained from Equation (3) using the coolant temperature and RTNDT at the 1/4T location. Here we may note that Mm must be iterated for since it is a function of the final stress ratio to yield strength (0./ay). At the 3/4T location, the maximum pressure is determined from Equation (ll) as KZR - Ku P (3/4T) (13) rj r oZ + (1/4r j + 3/41 o) 2M roZ r.Z (1/4ri+ 3/4ro)2 where K1R is obtained from Equation (2) using the material temperature and RTNDT at the 3/4T location and Klt is determined from the analysis procedure outlined in Section D. Mm is determined from Figure 1, B-9
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 bas is. 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 thermal gradient produces tensile stresses at the 1/4T location. Thus the steady state pressure is the same as that given in Equation (12). For each coo)down rate, the maximum pressure is evalu-ated at the 1/4T location from max( (14) 2M ri ro~ + (3/4ri 0 1/4r o) r - r ~ (3/4ri+ 1/4r ) where KIR 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 the 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. Hence a point-by-point
comparison must be made between the maximum allowable pressures pro-duced by steady state analyses and transient thermal analysis to determine the minimum of the maximum allowable pressures.
3. Core 0 eration At all times that the reactor core is critical, the temperature must be higher than that required for inservice hydrostatic testing, and in addition, the pressure-temperature relationship shall provide at least a 40'F margin over that required for heatup and cooldown operations. Thus the pressure-temperature limit curves for core operation may be constructed directly from the inservice leak and. hydrostatic test and heatup analysis results.
D. Thermal Stress Anal sis The equivalent linear stress due to thermal gradients is obtained from a detailed thermal analysis of the vessel., The temperature distribu-tion in the vessel wall is governed by the partial differential equation PcT< - K[(1/r) T + T .1 = o (15) subject to initial condition T(r,0) = T and boundary conditions
-KTr(ri, t) = hLTc(t) - T(ri t)
I (17)
and Tr(roit) = 0 (18) whe re Tc = To+ Rt. (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 various time steps along the heatup or cooldown cycle. The finite difference equations for N radial points, at distance 6r apart, across the vessel are: for 1<n<N T =Ll-htK 2(2 )JT
+
QtK (g )Z L
~
(1+ gr )Tn+1.+ Tn-1J (2o) (21) B-12
andfor n = N t+()t N [ pc(()r)Z J N pr())r)2 N-1 (22) For stability in the finite difference operation, we must choose ht for a given hr such that both ()t K pc(kr)22(2+ Zr )c r1 1 (23) and
~(1+
ht K pc(hr) (Ih,r rl )+ pc(hr) C 1 (24) are satisfied. These conditions assure us that heat will not flow in the 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'F increment in coolant temperature and then updating their values for the next 5'F increment. The dependence of (E/pc) called the thermal diffusivity and E, the thermal conductivity, can be determined from the ASME Boiler and Pressure Ves-sel Code, Section III, Appendix I- Stress Tables. A linear regression analysis of the tabular values resulted in the following expressions: K(T) = 38. 211 - 0. 01673 ~ T (BTU/HR-FT-'F) (25) B-13
and k(T) "-(K/pc) = 0. 6942 - 0. 000432 ~ T (FT /HR) (26) where T is in degrees Fahrenheit. The heat transfer coefficient 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~ h(T) = 170(1+10 ~T - 10 ~T ) v /D (27) where v is in ft/sec, D in inches, the temperature is in 'F, and h is in Btu/hr-ft - 'F. The values for the heat-transfer coefficient given by this relationship are in good agreement with those obtained from the Dittus-Boelter equation for temperatures up to 600'F. The mean velocity of the coolant, v, is generally given in terms of the effective coolant flow rate Q (Lbm/hr) and effective flow area A (ft ). Given the relationship p(T) = 62. 93 - 0. 48 x 10 2 <'- 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 = O/(3600 > p (T) ~ A) (29) Glasstone, S., Princi les of Nuclear Reactor Engineerin, D. Van Nostrand Co., Inc., New Jersey, pp. 667-668, 1960.
The thermal stress distribution is calculated from r2+ ri2 ro aT(r,t) = t [ jri T(r,t)rdr-T(r,t)+ 3 3 ( 0 3 1 3)jC T(r,t)rdrj (30) where a is the coefficient of thermal expansion (in/in 'F), E is Young's modulus, and v is Poisson's ratio. This expression can be obtained from Theor of Elasticit by Timoshenko and Goodier, pp. 408-409, when 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 n and E are evaluated as a function of the average temperature across the vessel T =
~(3 jri T(r)rdr (31)
The dependence of the coefficient of thermal expansion on temperature is taken to be a(T) = 5.76 x 10-6+ 4.4 x 10-9 4 T (32) and the dependence of Young's modulus on temperature is taken to be E(T) = 27.9142 + 2.5782 x 10 ~" T - 6.5723 x 10 6 4 T (33) as obtained from regression analysis of tabular values given in Section III, 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 B-15
r~ ~ stress distribution is given by ro M(t) = b f a T (r, t) rdr (34) where b is
unit depth of the vessel. Here we note that the moment is a function of time, i. e., coolant temperature via Tc = To + Rt. For a lin-ear stress distribution we have that P
=
Mc
~max I (35')
where 0 ax is the maximum outer fiber stress, c the distance from the neutral axis, taken to be (ro - ri)/2, and I the section area moment of inertia which is given by bh b(ro - r;)3 12 12 (36) Combining these expressions results in the equivalent linear stress due to thermal gradients ro rrttax rbt (r.-r ) TJJ 't'T (r') r~ (37) 1i The thermal stress intensity factor KIt is then defined as KIt = Mb 0 bt (38) where Mb is determined 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 stress calculations during a cooldown analysis since the thermal gradients
produce compressive stresses at the vessel outer radius. This sign change must then be reflected in the Klt calculation for the cooldown analys is. 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, Tavg, by (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 heatup cycle. Similar temperature and thermal stress distributions are developed during cool-down. The trends are nearly identical as those shown in Figure 4 when the inner and outer vessel locations are reversed with the I/4T location becoming the critical point. E. Exam le Calculations The following example is based on a reactor vessel with the follow-ing characteristics: Inner Radius 82. 00 in. (r ) Outer Radius 90 00 in. (r ) Operating Pressure 2250 psig (Po)
OUTER WALL 1.0
/ /
0.8 0.6 // 0.4
/ /
0.2 -1.0 1.0 -1.0 1.0 INNER WALL Norma lized temperature Normalized stress distribution ( 4T/h,Tma) distribution ( o/ omax ) Figure 4. Typical Normalized Temperature and Stress Distribution During Heatup
Initial Temperature 70'F (To) Final Temperature 550'F Effective Coolant Flow Rate 100 x 10 Lbm/hr (Q) Effective Flow Area 20. 00 ft2 (A) Effective Hydraulic Diameter = 10. 00 in. (D) RTNDT (1/4T) 200OF RTNDT (3/4T) 140'F In the thermal stress analysis 21 radial points were used in the finite difference scheme. Going from 70'F to the final temperature of 550'F, approximately 12, 000 time (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, KIR, as a function of temperature indexed to RTNDT (1/4T). For the steady state analysis, KIR is converted directly to allowable pressure via Equation 12. During the heatup and cooldown thermal analyses the material tem-perature at the 1/4T and 3/4T and thermal stress intensity factors Kzt are required to compute allowable pressure via Equations (13) and (14). The material temperatures versus coolant temperature during the 100'F/hr heatup and cooldown analyses are given in Figure 6. These temperatures allow computation of the corresponding reference stress intensity factors, KIR (3/4T) and KIR (1/4T). Figure 7 gives the corresponding thermal stress intensity factor at the 3/4T and 1/4T locations as a function of coolant tempe rature.
200 RTNDT(1i4T) - 200 160 F
~ 120 80 hC I
otV 40 50 150 200 250 300 350 400 TEMPERATURE ( F ) Figure 5. Reference Stress Intensity Factor as a Function of Temperature Indexed to RTNDT(1/4T )
400 100'F/HR HEATUP i 3/4T Location i 100'F /HR COOLDOWN 1/4T Location ( ) 300 200 100 50 100 150 200 250 300 350 COOLANT TEMPERATURE ('F ) Figure 6. Vessel Temperature at 1/4T and 3/4T Locations as a Function of Coolant Temperature
10 cu hC 6 100'F/HR HEATUP (3/4T Location 100'F/HR COOLDOWN ( 1/4 Location i
)
50 10Q 150 200 250 3QQ 350 COOLANT TEMPERATURE ('F ) Figure 7. Thermal Stress Intensity Factor at 3/4T and 1/4T Locations as a Function of Coolant Temperature
Figures 8 and 9 demonstrate the construction of the allowable com-posite pressure and temperature curves for the 100'F/hr heatup and cool-down rates. The composite curves represent the lower bound of the thermal and steady state curves with the addition of margins of +10'F and -60 psig for possible instrumentation error. Figure 8 also shows the leak test 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 providing for a 40'F 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. B -23
LEAK TEST LIIIIIIT 2400 2000 COMPOS ITE CURVE 100'F/HR HEATUP ( Margins of +10 F and -60 psig for instrument error ) 1600 I STEADY STATE CR I TI CALITY 1200 LIMIT 800 HEATUP 400 50 100 150 200 250 300 350 400 INDICATED TEMPERATURE ( F ) Figure 8. Pressure- Temperature Curves for 100 F/Hr Heatup
2400 2000 COMPOSITE CURVE -100 F/HR COOLDOWN ( Margins of +10 F and
-60 psig for instrument error )
1600 CXI PJ 1200 CD COO LDOWN Ch 800 STEADY STATE 400 50 100 150 200 250 300 350 INDICATED TEMPERATURE ('F ) Figure 9. Pressure-Temperature Curves for 100'F/Hr Cooldown
ADDENDUM TO FINAL REPORT ON "REACTOR VESSEL MATERIAL SURVEILLANCE PROGRAM FOR DONALD C. COOK UNIT NO. 1, ANALYSIS OF CAPSULE T" Plate B4406-3 Held Held Correlation ,30 ft-1b C Tem . '(de T) ~(lan .) (Ttana.) Metal Mtt Monitor Irradiated 65 90 ~. -10 '0 105 Unirradiated. 5 20 -90 -100 45 AT 60 70 80 120 60 Monitor Height Identification ~(m ) Fe Top 18.2 Fe Top Mid. 15.3 Fe Mid. 17.2 Fe Bot. Mid. 16.6 Fe Bot. 16.4 Cu *- Top Mid. 64.9 Cu - Mid. 62.9 Cu Bot. Mid. 70.9 Ni - Top Mid. 22.9 Ni Mid. 25.5 Ni Bot. Mid. 24.5 Co Top 9.3 Co(Cd) Top 8.7 Co-- Bot. 9.5 Co(Cd) Bot. 7.7 U-238 12.0(a) NP-237 20.0(a) (a) As reported in WCAP-8047.
i ADDENDUM NO. 2 TO FINAL REPORT ON "REACTOR VESSEL MATERIAL SURVEILLANCE PROGRAM FOR DONALD C. COOK UNIT NO. 1, ANALYSIS OF CAPSULE T" Additional Tensile Test Data Specimen Fracture Load ~ Fracture Stress Uniform Elongation< > No. si %%u4 64,700 188,600 5.00 63,250 177,000 2.45 W9 87,600 250,000 4.56 757800 193,700 2.87 (a) Using method of change in cross-sectional area of unnecked portion of specimen per ASTM E 184-62.}}

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