ML100200239

Reactor Vessel Material Surveillance Program,Analysis of Capsule T Submitted in Support of Amend 1 to Application for Amend to License DPR-26.Vessel Matl Surveillance Capsule Was Tested & Evaluated

ML100200239
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Site:	Indian Point
Issue date:	06/30/1977
From:	Norris E Consolidated Edison Co of New York
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ML100200241	List: ML100200239 ML19270F004 ML19270F005
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1 REACTOR VESSEL MATERIAL SURVEILLANCE PROGRAM FOR INDIAN POINT UNIT NO. 2 ANALYSIS OF CAPSULE T by E. B. Norris FINAL REPORT SwRI Project'02-4531 to Consolidated Edison Company of New York, Inc.

4 Irving Place New York, New York 10003 June 30, 1977

~-'Al4>

' 0I 01 SOUTHWEST RESEARCH INSTITUTE SAN ANTONIO CORPUS CHRISTI HOUSTON

ERRATA SHEET (Dec., 1978)

"Reactor Vessel Material Surveillance Program for Indian Point Unit No. 2 Analysis of Capsule T" (June 30, 1977)

(1) pg. 18, TABLE III: The Measured activities of Cu, Ni and Co given in Colurn 3 are one Trder of magnitude too high (e.g., 5.28xi0 2 for Cu should be 5.28xi0 , etc.).

(2) pg. 30, TABLE IX: The Cv Upper Shelf Fnergy (ft-lb) for the Correlation Monitor (4) should be:

Unirradiated: 78 Irradiated: 68 E,ft-lbs: i0 E,% 13 (3) pg 42, paragraph(3):

105OF should be 110°F 270°F should be 280°F 130°F should be 140°F

SOUTHWEST RESEARCH INSTITUTE Post Office Drawer 28510, 8500 Culebra Road San Antonio, Texas 78284 REACTOR VESSEL MATERIAL SURVEILLANCE PROGRAM FOR INDIAN POINT UNIT NO. 2 ANALYSIS OF CAPSULE T by E. B. Norris FINAL REPORT, SwRI Project 02-4531 to Consolidated Edison Company of New York, Inc.

4 Irving Place New York, New York 10003 June 30, 1977 Approved:

U. S. Lindholm, Director Department of Materials Sciences 7 9O) 0 O

ABSTRACT The first vessel material surveillance capsule removed from the Indian Point Unit No. 2 nuclear power plant has been tested, and the re sults have been evaluated. Heatup and cooldown limit curves for normal operation have been developed for up to 5 effective full power years of ope ration.

TABLE OF CONTENTS Page LIST OF TABLES LIST OF FIGURES I.

SUMMARY

OF RESULTS AND CONCLUSIONS II. BACKGROUND Ill. DESCRIPTION OF MATERIAL SURVEILLANCE PROGRAM IV. TESTING OF SPECIMENS FROM CAPSULE T V. ANALYSIS OF RESULTS VI. HEATUP AND COOLDOWN LIMIT CURVES FOR NORMAL OPERATION OF INDIAN POINT UNIT NO. 2 VII. REFERENCES APPENDIX A - TENSILE TEST RECORDS APPENDIX B PROCEDURE FOR THE GENERATION OF ALLOWABLE PRESSURE-TEMPERATURE LIMIT CURVES FOR NUCLEAR POWER PLANT REACTOR VESSELS

LIST OF TABLES Table Page I Indian Point Unit No. 2 Reactor Vessel Surveil- 9 lance Materials II Summary of Reactor Operations 16 Indian Point Unit No. 2 III Summary of Neutron Dosimetry Results 18 Capsule T IV Fast Neutron Spectrum and Iron Activation 20 Cross Sections for Capsule T V Charpy V-Notch Impact Data 22 Indian Point Unit No. 2 Pressure Vessel Shell Plate B2002-1 VI Charpy V-Notch Impact Data 23 Indian Point Unit No. 2 Pressure Vessel Shell Plate B2002-2 VII Charpy V-Notch Impact Data 24 Indian Point Unit No. 2 Pressure Vessel Shell Plate BZ002-3 VIII Charpy V-Notch Impact Data 25 Correlation Monitor Material (Supplied by U. S. Steel)

IX Notch Toughness Properties of Capsule T 30 Specimens Indian Point Unit No. 2 X Tensile Properties of Surveillance Materials 31 Capsule T XI Projected Shifts in RTNDT for Indian Point 39 Unit No. 2 XII Proposed Reactor Vessel Surveillance 41 Capsule Schedule Indian Point Unit No. 2

LIST OF FIGURES Figure Page 1 Arrangement of Surveillance Capsules in the 8 Pressure Vessel 2 Vessel Material Surveillance Specimens 11 3 Arrangement of Specimens and Dosimeters 12 in Capsule T 4 Effect of Irradiation on Cv Impact Properties 26 of Indian Point Unit No. 2 Shell Plate B2002-1 5 Effect of Irradiation on C v Impact Properties 27 of Indian Point Unit No. 2 Shell Plate B2002-2 6 Effect of Irradiation on Cv Impact Properties 28 of Indian Point Unit No. 2 Shell Plate B2002-3 7 Effect of Irradiation on C v Impact Properties 29 of Indian Point Unit No. 2 Correlation Monitor Mate rial 8 Dependence of C v Shelf Energy on Neutron 36 Fluence, Indian Point Unit No. 2 9 Effect of Neutron Fluence on RTNDT Shift, 38 Indian Point Unit No. 2 10 Estimated Transverse Charpy V-Notch 44 Properties of Plate B2002-3 11 Indian Point Unit No. 2 Reactor Coolant Heatup 46 Limitations Applicable for Periods up to 5 Ef fective Full Power Years 12 Indian Point Unit No. 2 Reactor Coolant Cool- 47 down Limitations Applicable for Periods up to 5 Effective Full Power Years 13 Indian Point Unit No. 2 Inservice Leak Test 48 Limitations Applicable for Periods up to 5 Effective Full Power Years

I.

SUMMARY

OF RESULTS AND CONCLUSIONS The analysis ofthe first material surveillance capsule removed from the Indian Point Unit 2 reactor pressure vessel led to the following conclusions:

(1) Based on a calculated neutron spectral distribution, Cap sule T received a fast fluence of 2. 02 x 1018 neutrons/cm 2 > 1 MeV.

(2) The surveillance specimens of the three core beltline plates experienced shifts in transition temperature of 850F to 130 OF as a result of the above exposure.

(3) Plate B2002-3 exhibited the largest shift and will control the heatup and cooldown limitations.

(4) The estimated maximum neutron fluence of 6. 97 x 1017 neutrons/cm 2 > 1 MeV received by the vessel wall accrued in 1.42 full power years. Therefore, the projected maximum neutron fluence after 32 effective full power years (EFPY) is 1. 57 x 1019 neutrons/cm 2 > 1 MeV.

This estimate is based on a lead factor of 2. 9 between Capsule 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 Indian Point Unit 2 vessel core beltline plates at the 1/4T and 3/4T positions after 5 EFPY of operation are 110OF and 50°F, respectively. These values were used as the bases for computing heatup and cooldown limit curves for up to 5 EFPY of operation.

(6) The maximum shifts in the transition temperature of the Indian Point Unit 2 vessel core beltline plates at the 1/4T and 3/4T 0

positions after 32 EFPY of operation are projected to be 280 F and 140 °F, respectively.

(7) The Indian Point Unit 2 vessel plates located in the core beltline region are projected to retain sufficient toughness at the 1/4T and 3/4T positions to meet the current requirements of 10CFR50 Appen dix 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 tempera ture (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 RTNDT per the requirements of Appendix H, "Reactor Vessel Material Surveillance 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 3

ASTM E208;( )

(2) 60 deg F below the 50 ft-lb Charpy V-notch (Cv) temperature; (3) 60 deg F below the 35 mil Cv 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 comprise 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 Superscript numbers refer to references at the end of the text.

3

per cm 2 (E > 1 MeV). (4) to neutron fluences in excess of 1017 neutrons Also, it has been established that tramp elements, particularly copper and phosphorous, affect the radiation embrittlement response of ferritic materials. (5-7) The relationship between increase in RTNDT and copper content is not defined completely. For example, Regulatory Guide 1. 99, (7 ) , proposes an originally issued in July 1975, and revised in April 1977 adjustment to RTNDT proportional to the square root of the neutron fluence.

Westinghouse Electric Corporation, in their comments on the 1975 issue of Regulatory Guide 1. 99(8), believed that the proposed relationship overesti mates the shift at fluences greater than 1. 9 x 1019 and underestimates the shift at fluences less than 1. 9 x 1019. On the other hand, Combustion Engi neering, in their comments on the 1975 issue of Regulatory Guide 1. 99(9),

suggested that the proposed relationship is overly conservative at fluences 2 (E > 1 MeV). There is also disagreement con below 1019 neutrons per cm cerning the prediction of Cv upper shelf response to exposure to neutron ir radiation. (7-9) After reviewing the comments and evaluating additional sur veillance program data, the NRC issued a revision to Regulatory Guide 1. 99 which raised the upper limit of the transition temperature adjustment curve.

In this report, estimates of shifts in RTNDT are based on Regulatory Guide 1.99, Revision 1. (7)

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 Spec imens machined from the plate or forging material and weldments which are located in such a region of high neutron flux density. ASTM IE 185(10) de scribes the current 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 Indian Point Unit No. 2 nuclear power plant. The encapsulated Cv speci mens are located on the 0. D. 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 transition temperatures of the materials in the pressure vessel are generally less than the corresponding shifts observed in the surveillance specimens. However, because, of azi muthal variations in neutron flux density, capsule fluences may lead or lag the maximum vessel fluence in a corresponding exposure period. For ex ample, Capsule T (removed during the 1976 refuelling outage) was exposed to a neutron fluence approximately three times that at the maximum exposure point on the vessel 1. D. , while Capsule V (scheduled for removal at a later date) is receiving a neutron flux about equal to that at the point of maximum vessel exposure. The capsules also contain several. dosimeter materials for experimentally determining the average neutron flux density at each capsule location during the exposure period.

The Indian Point Unit No. 2 material surveillance capsules also in clude tensile specimens as recommended by ASTM E 185. At the present time, irradiated tensile properties are used only to indicate that the materials tested continue to meet the requirements of the appropriate material specification. In addition, some of the material surveillance capsules contain wedge opening loading (WOL) fracture mechanics specimens.

Current technology limits the testing of these specimens at temperatures well below the minimum service temperature to obtain valid fracture mechanics data per ASTM E 399( 11), "Standard Method of Test for Plane Strain Fracture Toughness of Metallic Materials." However, recent work reported by Mager and Witt('?-) 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 pro cedure 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 1976 refuelling outage as well as predicting the changes expected to oc-.

cur at selected times in the future operation of the Indian Point Unit No. 2 power plant.

III. DESCRIPTION OF MATERIAL SURVEILLANCE PROGRAM The Indian Point Unit No. 2 material surveillance program is de scribed in detail in WCAP 73Z3(13), dated May 1969. Eight materials surveillance capsules (five Type I and three Type I) 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 neutron flux density at the Capsule T location leads the maximum flux density on the vessel I. D. by a factor of 2.9. (14) The Type I capsules each contain Charpy V-notch, tensile and WOL specimens machined from the three SA533 Gr B plates located at the core beltline plus Charpy V-notch specimens machined from a reference heat of steel utilized in a number of Westinghouse surveillance programs.

The Type II capsules include specimens machined from weld metal and, HAZ material representative of those materials in the core beltline region of the vessel as well as base plate material. Capsule T, one of the Type I capsules, was removed during the 1976 refuelling outage.

The chemistries and heat treatments of the vessel surveillance ma terials contained in Capsule T are summarized in Table I. All test speci mens were machined from each of the materials at the quarter-thickness (1/4 T) location. The base metal Cv specimens were oriented with their long axis parallel to the primary rolling direction; the Cv notches were perpendicular to the major plate surfaces. Tensile specimens were

Z (Type I)

S (T ype II)

U (Type I) 00 W (Type I)

FIGURE 1.- ARRANGEMENT OF SURVEILLANCE CAPSULES IN THE PRESSURE VESSEL

TABLE I Indian Point Unit No. 2 Reactor Vessel Surveillance Materials( 13 )

Heat Treatment History Shell Plate Material:

1550°-1600°F, 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br />, water quenched 1225' +/- 25'F, 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br />, air cooled 11500 +/- 25 0 F, 40 hours4.62963e-4 days <br />0.0111 hours <br />6.613757e-5 weeks <br />1.522e-5 months <br />, furnace cooled to 600°F Weldment:

11500 +/- 25 0 F, 19.75 hours8.680556e-4 days <br />0.0208 hours <br />1.240079e-4 weeks <br />2.85375e-5 months <br />, furnace cooled to 600'F Correlation Monitor:

1650 °F, 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br />, water quenched to 300°F 1200°F, 6 hours6.944444e-5 days <br />0.00167 hours <br />9.920635e-6 weeks <br />2.283e-6 months <br />, air cooled Chemical Composition (Percent)

Material C Mn P S Si Ni Mo Cu (1 5)

Plate B2002-1 0.20 1.28 0.010 0.019 0.25 0.58 0.46 0.25 Plate B2002-2 0.22 1.30 0. 014 0. 018 0.22 0.46 0.50 0.14 Plate B2002-3 0.22 1.29 0.011 0. 020 0.25 0.57 0.46 0.14 Corr. Monitor 0.24 1.34 0.011 0. 023 0.23 (a) 0.51 (a)

Weld Metal (a) (a) (a) (a) (a) (a) (a) (a)

(a). Not reported.

machined from the three vessel plates with the longitudinal axis parallel to the plate rolling direction. The WOL specimens were machined with the simulated crack perpendicular to the primary rolling direction and to the 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 32 Charpy V-notch specimens (8 each from the three vessel plate materials plus 8 from the reference steel plate);

3 tensile specimens (1 from each plate); and 6 WOL specimens (Z from each plate). 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 Copper Bare wire 2 Nickel Bare wire 1 Cobalt (in aluminum) Bare wire 3 Cobalt (in aluminum) Cd shielded wire 3 In addition, slices were taken from five Cv specimens to serve as iron dosimeters.

Three eutectic alloy thermal monitors had been inserted in holes in the steel spacers in Capsule T. Two (located top and bottom) were 2. 5% Ag and 97. 5% Pb with a melting point of 579 *F. The third (located at the center of the capsule) was 1. 75% Ag, 0. 75% Sn and 97.5% Pb having a melting point of 590 "F.

.011R

.009 (a) Charpy V-notch Impact Specimen SECTION A-A (b) Tensile Specimen (c) Wedge Opening Loading Specimen FIGURE 2. VESSEL MATERIAL SURVEILLANCE SPECIMENS 11.

Co Ni Co (Cd) Cu Co (Cd)

To Vessel Bottom Vessel Wall Side

-4 2 - Plate B2002-2 3 - Plate B2002-3 R - Reference Material 1 - Plate BZ00Z-l FIGURE 3. ARRANGEMENT OF SPECIMENS AND DOSIMETERS IN CAPSULE T

IV. TESTING OF SPECIMENS FROM CAPSULE T The capsule shipment, capsule opening, specimen testing and report ing of results were carried out in accordance with the following SwRI Nuclear Project Operating Procedures:

(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.

(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 1976 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 appropri ate cutting tools and a licensed shipping cask. Todd personnel severed the capsule 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.

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 milling 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 7323. (13)

Each specimen was inspected for identification number, which was checked against the master list in WCAP 73Z3. No discrepancies were found. The thermal monitors and dosimeter wires 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 ex ceed 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 60 Go-, 54 Mn- and 58 Co-containing dosimeter wires. All activities were corrected to the time-of-removal (TOR) at reactor shutdown.

Infinitely dilute saturated activities (ASAT) were calculated for each of the dosimeters because ASAT is directly related to the product of the energy dependent microscopic activation cross section and the neutron flux density.

The relationship between ATOR and ASAT is given by:

AT OR m-n

( - XTm) (e - Xtm )

ASAT m1 where: X decay constant for the activation product, day-';

Tm equivalent operating days at 2758 MwTh for op erating period m; tm decay time after operating period m, days.

An alternate expression which gives equivalent results is:

m-n ATOR = Pm (-e -XT O)(e- -Xt m)

ASAT ml where: To operating days; Pm = average fraction of full power during operating period.

The Indian Point Unit No. 2 operating history up to the 1976 refuelling shut down is presented in Table II. The specific time of release and specific saturated activities for each dosimeter are presented in Table III.

The primary result desired from the dosimeter analysis is the total fast neutron fluence (> 1 MeV) which the surveillance specimens received.

TABLE II Summary of Reactor Operations Indian Point Unit No. 2 Shutdown Ope rating Decay Time Fraction of Full Period Date s Power in Period, Pm Days Days, Tm After Period, tm (M) Start Stop 1 08-15-73 08-24-73 949 0. 4377 08-Z5-73 08-25-73 0. 4532 935 2 08-26-73 09-07-73 13 09-08-73 09-20-73 0.3161 914 3 09-21-73 09-28-73 09-29-73 09-30-73 Z 0. 3088 900 4 10-01-73 10-12-73 10-13-73 01-25-74 105 791 0. 2412 5 01-26-74 01-29-74 01-30-74 03-21-74 51 0. 5438 712 6 03-22-74 04-18-74 04-19-74 04-28-74 10 697 0. 4962 7 04-29-74 05-03-74 05-04-74 05-04-74 1 690 0. 4743 8 05-05-74 05-10-74 05-11-74 05-12-74 z 687 0. 0730 9 05-13-74 05-13-74 05-14-74 05-20-74 7 655 0. 6653 10 05-21-74 06-14-74 06-15-74 06-16-74 2 617 0. 7691 11 06-17-74 07-22-74 07-23-74 07-23-74 1 613 0. 7593 12 07-24-74 07-26-74 07-27-74 08-05-74 10 571 0.6653 13 08-06-74 09-06-74 09-07-74 09-09-74 3 (Continued)

TABLE II (Cont'd.)

Period Dates Shutdown Ope rating Decay Time Fraction of Full (M) Start Stop Days Days, Tm After Period, t0 Power in Period, 14 09-10-74 09-30-74 547 0. 7429 10-01-74 10-11-74 15 10-12-74 11-09-74 507 0. 8657 11-10-74 11-10-74 16 11-11-74 12-06-74 480 0. 8306 12-07-74 12-07-74 17 12-08-74 01-01-75 454 0. 8495 01-02-75 01-04-75 18 01-05-75 01-05-75 450 0.5450 01-06-75 01-06-75 19 01-07-75 01-31-75 424 0.8810 02-01-75 02-02-75 20 02-03-75 02-28-75 396 0. 9408 03-01-75 04-03-75 21 04-04-75 05-02-75 333 0.7652 05-03-75 05-03-75 22 05-04-75 07-28-75 246 0.9114 07-29-75 08-10-75 23 08-11-75 09-12-75 200 0.7108 09-13-75 09-13-75 24 09-14-75 10-16-75 166 0. 7962 10-17-75 10-29-75 25 10-30-75 11-14-75 137 0. 7467 11-15-75 11-15-75 26 11-16-75 01-0.4-76 0.8427 01-05-76 01-05-76 27 01-06-76 01-29-76 0. 8703 01-30-76 02-04-76 28 02-05-76 03-30-76

0. 9122 Total Power Generation = 517. 749 Effective Full Power Days (includes 5. 749 EFPD accumulated before 08-15-73)

TABLE III Summary of Neutron Dosimetry Results Capsule T Monitor Activation Measured Activity Saturated Activity Identification Reaction (dps/mg) (dps /mg)

Rl(a) 4Fe(n, p)54Mn 1. 57 x 103 2. 72 x 103 R3(a) ft 1. 66 x 103 2.88 x 103 R4(a) I! 1. 59 x 103 2. 77 x 103 R6 (a) 1! 1. 59 x 103 2.77 x 103 R8(a) 1.49 x 10 3 2.-58 x 103 Avg. 1. 58 x 10 3 2.75 x 103 3.25 x 103 6 3 Cu(n.ca) 6 5.28x 10O Cu (Top) 0C o Z

4.53 x i0 2.79 x 103 1t Cu (Bottom)

Ni (Center) 58Ni(n, p)58Co 3.69 x 105 4.68 x 105 8

Co (Top) 5 9 Co(n, y) 6 0 Co 5.00 10 7 3. 08 x 10 107 8 Co-Cd (Top) !I 2.36 1.46 x 108 107 1.0 Co (Center) ft 5.66 3.49 x 107 Co-Cd (Center) f! 2.24 107 1.38 x 108 Co (Bottom) !I 5.14 3.16 x 108 107 Co-Cd (Bottom) 2.31 1.42 x 108 (a) Charpy specimen from which sample was removed.

The average flux density at full power is given by:

ASAT NO&

where: CP = energy-dependent neutron flux density, n/cm2-sec; ASAT saturated activity, dps/mg target element; T = spectrum-averaged activation cross section, cm 2 ;

No = 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.

The neutron flux density was calculated from the 54 Fe(n, p) 54 Mn reaction 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 PI expansion of the scattering ma trix and an S 8 order of angular quadrature. The normalized spectrum for Capsule T and the group-organized cross sections for the 5 4 Fe(n, p) 5 4 Mn reaction derived from the ENDF/B-IV library( 1 6 ) are given in Table IV.

The value of rFe is given by:

10 MeV T Fe (E) P (E) dE

(> 1 MeV)

TFe 1 10 MeV Co (E) dE 1.00 where: rFe (> 1 MeV) the calculated spectrum-averaged cross section for flux> 1 MeV, cm Z determined for the 5 4 Fe(n, p)5 4 Mn reaction.

19

TABLE IV Fast Neutron Spectrum and Iron Activation Cross Sections for Capsule T 54 Fe(n, p) 5 4 Mn Normalized Cross Section Energy Range Neutron Flux (barns)

(MeV)

0. 0105 0.581 8.18 - 10.0

0. 0284 0. 577 6.36 - 8.18

0. 0535 0.491 4.96 - 6.36

0. 0535 0.354 4.06 - 4.96

0. 0925 0.205 3.01 - 4.06

0. 1451 0. 099 2.35 - 3.01

0. 1705 0.023 1.83 - 2.35

0. 4459 0.0014 1.11 - 1.83

'Fe = 0. 097 barns

The resulting value obtained for fast (> 1 MeV) neutron flux density at the Capsule T location was 4.51 x 1010 neutrons/cm,-sec. Since Indian Point Unit No. 2 operated for an equivalent of 518 full power days up to the 1976 refuelling outage, the total neutron fluence for Capsule T is equal to 2. 02 x 1018 neutrons/cm 2 (E > 1 MeV).

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 specimens contained in Capsule T are presented in Tables V through VIII. The Charpy V-notch transition curves for the three plate materials and the correlation monitor material are presented in Figures 4 through 7.

The radiation-induced shift in transition temperatures for the vessel plates are indicated at 77 ft-lb and 54 mil lateral expansion because the specimens are longitudinally oriented. A summary of the shifts in RTNDT and C v up per shelf energies for each material are presented in Table IX.

Tensile tests were carried out in the SwRI hot cells using a Dillon 10, 000-lb capacity tester equipped with a strain gage extensometer, load cell and autographi& recordin'ge " uipment. 'Allof the t e sp cimns-"....

were tested at 550 F!. The resiilts, along wijh tensile data reportedi by Westinghouse on theufirradited materiale(1 3), are pre~ented in Table X.

The load-strain records are included in Appendix A.

TABLE V Charpy V-Notch Impact Data Indian Point Unit No. 2 Pressure Vessel Shell Plate B2002-1 Lateral Spec. Temp. Energy Shear Expansion Condition No. (-F) (ft-lbs) (mils)

M)

Baseline (a) -40 10.0 10 8

-40 9. 0 10 6

-40 8.0 10 7

-z0 19.0 15 15

-20 14.5 15 16

-0 8.5 15 8 0 21.5 25 18 0 33.5 25 28 0 34.0 25. 29 10 40.5 30 32 10 36.0 25 28 60 35.5 25 27 60 68.5 45 54 60 62. 0 45 49 60 50.5 40 40 88.0 65 70 110 56

69. 5 60 110 77.0 60 61 160 116.5 98 84 160 122. 0 98 86 160 I12.0 95 85 Z10 111.0 100 83 210 121.5 100 88 210 119.0 100 83 Capsule T 1-6 40 27. 0 5 23 1-1 78 38.5 10 33 1-5 120 53.5 15 44 1-7 140 52. 0 20 47 1-3 160 69.0 60 56 1-8 180 69.5 60 61 1-2 Z10 95. 5 95 81 300 99.5 100 81 1-4 (a) Not reported.

TABLE VI Charpy V-Notch Impact Data Indian Point Unit No. 2 Pressure Vessel Shell Plate B2002-2 Lateral Spec. Temp. Energy Shear Expansion Condition No. -4F) (ft-lbs) M.2L (mils)

Baseline (a) -40 6.5 10 4

-40 8.0 10 6

-40 7.5 10 6

-20 18.0 10 18

-20 9.0 10 7 10.0 10 7 10 20.0 25 19 10 42.5 30 34 10 14.0 25 14 30 37. 5 30 32 30 41.0 35 35 30 49.5 30 41 60 73.5 40 58 60 49.0 40 39 6o 59.5 40 50 110 80.5 65 60 110 76.0 60 60 110 91.0 65 70 160 108.0 85 84 160 120.5 95 85 160 118.0 95 84 210 112.0 100 80 210 120..5 100 83 210 115.0 100 82 Capsule T 2-1 78 10.0 5 9 2-6 100 43. 0 10 44 2-5 120 22.5 5 18 2-7 140 63.0 35 49 2-3 160 65.0 70 57 2-8 180 97.5 80 53 2-2 210 103.0 100 85 2-4 300 88.0 100 72 (a) Not reported.

TABLE VII Charpy V-Notch Impact Data Indian Point Unit No. 2 Pressure Vessel Shell Plate B2002-3 Lateral Spec. Temp. Energy Shear Expansion Condition No. (OF)[ (ft-lbs) _LL (mils)

Baseline -40 6.5 10

-40 8.0 10

-40 6.0 10

-20 25. 5 2o

-20 14.0 15

-20 11.0 15 10 41. 5 25 10 17. 0 25 10 37. 5 25 30 34.0 35 30 45.5 35 30 42. 5 35 60 54.5 40 60 51.5 40 60 41.0 40 110 71. 0 60 110 79.5 70 110 83.5 70 160 116.5 99 160 110.0 95 160 95.5 90 210 109.0 90 210 113.5 100 210 113. 0 100 Caps ule T 3-1 78 20.0 5 3-3 160 34.5 Z0 3-8 180 40.5 10 3-2 210 53.0 35 3-7 260 88. 0 100 3-4 300 88.0 100 3-5 350 92.5 100 3-6 400 89.5 100 (a) Not reported.

TABLE VIII Charpy V-Notch Impact Data Correlation Monitor Material (Supplied by U. S.

Steel)

Late ral Spec. Temp. Energy Shear Expansion Condition No. (ft-lbs) (%) (mils)

-150 Baseline (a) -150 12.5 10 10

-150 10.5 15 11

-100 35.0 25 29

-100 9. 0 20 9

-100 18. 0 30 19

-80 13.0 20 12

-80 32.5 20 27

-80 26.0 20 23

-40 34. 0 30 30

-40 35.5 35 31

-40 48.0 35 40 10 78.5 60 64 10 74.0 60 60 10 81. 0 70 68 60 102.5 80 78 60 102.0 85 82 60 100.0. 85 80 110 112.5 99 88 110 108.5 90 87 110 108.5 98 88 160 115.5 100 90 160 113.0 100 92 160 120.0 100 93 210 121.0 100 92 210 123.5 100 91 210 117.5 100 92 Capsule T 40 9.5 nil 9 78 20.5 5 18 100 31.0 10 27 120 35.5 15 32 140 79.5. 100 66 160 49. O 70 41 180 52.5 60 40 210 68.0 100 60 (a) Not reported.

120 80 40 0

-100 0 100 200 300 400 500 Temperature, Deg F 120 0

- 100 0 100 200 300 400 500 Temperature, Deg F FIGURE 4. EFFECT OF IRRADIATION ON Cv IMPACT PROPERTIES OF INDIAN POINT UNIT NO. 2 SHELL PLATE B2002-1

IZO 40 0

00i 0 100 200 300 400 500 Temperature, Deg F 120

-100 0 100 200 300 400 500 Temperature, Deg F FIGURE 5. EFFECT OF IRRADIATION ON Cv IMPACT PROPERTIES OF INDIAN POINT UNIT NO. Z SHELL PLATE B2002-2

120 40 0

-100 0 100 200 300 400 500 Temperature, Deg F 1 2 0. . ... .. ....

0L""

-100 0 100 200 300 400 500 Temperature, Deg F FIGURE 6. EFFECT OF IRRADIATION ON Cv IMPACT PROPERTIES OF INDIAN POINT UNIT NO. 2 SHELL PLATE B2002-3

-100 0 100 200 300 400 500 Temperature, Deg F H-+

-100 0 100 200 300 400 500 Temperature, Deg F FIGURE 7. EFFECT OF IRRADIATION ON Cv IMPACT PROPERTIES OF INDIAN POINT UNIT NO. 2 CORRELATION MONITOR MATERIAL

TABLE IX Notch Toughness Properties of Capsule T Specimens Indian Point Unit No. 2 Plate Plate Plate Correlation 2( 2 ) (3) (4 )

B200Z-l ( ) B2002- B200Z-3 Monitor 77 ft-lb Cv Temp. (deg F)

Irradiated 178 158 240 165 Unir radiated 88 83 110 75 AT 90 75 130 90 54 mil C v Temp. (deg F)

Irradiated 157 152 225 135 Unirradiated 77 67 95 55 AT 80 85 130 80 Cy Upper Shelf Energy (ft-lb)

Unirradiated 117 116 113 116 Irradiated 99.5 103 88 68 AE, ft-lbs 17.5 13 25 48 11 (1) 0. 16% Copper (2) 0. 17% Copper (3) 0.25% Copper (4) Energy transition at 50 ft-lb Lateral expansion transition at 35 mil

TABLE X Tensile Properties of Surveillance Materials Capsule T Test 0.2% Yield Tensile Total Reduction Specimen Temp. Strength Strength Elongation in Area Condition Ident. (OF) (psi) (psi) (%0) (7%)

Baseline B2002- 1 Room 68,500 89,000 25.1 67.8 Room 65,850 87,800 25.3 67.4 200 61,550 79,900 24. 1 68.6 200 67,950 89,400 23. 8 67.6 400 57, 900 79, 900 23.1 64.7 400 59,800 82, 20O 22. 2 67.8 600 56,750 80,550 21.9 64.3 600 57,750 85, 700 22.9 64.2 1-1 Caps ule T 550 71,100 96,200 20.4 55.5 Baseline BZOO2 -2 Room 62,350 83,800 27.1 70.0 Room 66,750 90, 500 28.2 69.6 200 63,650 84,450 24.8 70.5 200 63,200 83,800 25.5 67.3 400 53,800 77,900 23.1 68.5 400 52,650 73, 150 22.4 67.6 600 53,500 78,800 22.7 64.4 600 54,700 81,450 24. 7 64.4 Capsule T 2-1 550 58,600 85,900 20.7 64.8 Baseline B2002-3 Room 65,650 87,300 27.6 67.3 Room 65, 000 87,350 24.8 66.7 200 67,800 88,900 23.4 68.6 200 67,700 89, 150 22.1 64.9 400 57,950 79,550 22. 3 68.7 400 55,350 77,100 23.2 64.9 600 57,750 83,850 24. 9 68.2 600 58,350 86,500 24. 9 64.7 Capsule T 3-1 550 63, 100 89,400 20. 2 63. 1

Testing of the WOL specimens was deferred at the request of Consolidated Edison Company. The specimens are in storage at the SwRI radiation laboratory.

The Charpy V-notch results indicate that Plate B2002-3 is more sensitive to radiation embrittlement than Plates B2002-1 and B2002-2, in opposition to the expectation that Plate B2002-1 would be the most (15) sensitive because it was reported to have the highest copper content.

Check analyses for copper content were carried out on seven broken Charpy V-notch specimens, using an X-ray fluorescence technique meeting the requirements of ASTM E 322(17), with the following results:

Plate No. Specimen No.  % Copper B2002-1 1-2 0.17 B2002-1 1-3 0.15 B2002-2 2-2 0. 18 B2002-2 2-3 0. 17 B2002-3 3-2 0.27 B2002-3 3-3 0.23 Correlation Monitor R-2 0.25 These results correlate with the Charpy V-notch impact data.

The tensile properties of Plate B2002-1 appeared to be the most affected by the radiation exposure in Capsule T. Check analyses for copper on three tested tensile specimens gave the following results:

Plate No. Specimen No. %oCopper B2002-l 1-1 0.z1 B2002-2 2-1 0.13 B2002-3 3-1 0.09

These results correlate with the observed tensile test data. It is not known why the copper content of the test specimens is different from the reported copper content of the reactor vessel plates.

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 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.

(Z) 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 distri bution of the neutron flux.

The energy and spatial distribution of the neutron flux for Indian Point Unit No. 2 was calculated for Capsule T with the DOT 3. 5 discrete ordinates transport code. The lead factor for Capsule T reported by West inghouse is 2. 9 for the vessel I. D. surface. This was supported by the SwRI DOT 3.5 analysis. This analysis also predicted that the fast flux at the l/4 T and 3/4 T positions in the pressure vessel wall would be 50% and 8. 5%, re spectively, 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. (18) However, currently the NRC

prefers 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. (19) This conservatism allows 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 reduction in Cv upper shelf energy as a function of neutron fluence is given in Regulatory Guide 1. 99, Revision 1. (7)

However, it has been suggested( 9 ) that a square root of fluence dependence fits existing data better. The results from Capsule T are compared to a portion of Figure 2 of Regulatory Guide 1. 99, Revision 1, in Figure 8.

The embrittlement response of Plate B2002-3, the most sensitive of the three plates, is in good agreement with the prediction of Regulatory Guide 1.99, Revision 1.

The projection of the Cv shelf energy of base Plate B2002-3 is com plicated by the fact that the surveillance specimens are all oriented in the strong" direction and the 50 ft-lb lower limit of 10CFR50 Appendix G ap plies to "weak" direction properties. In a method established by Westing house (20)

, the estimated upper shelf energy in the "weak" direction is taken to be 65% of that in the "strong" direction. Therefore, the unirradiated Cv shelf energy of Plate B2002-3 is estimated to be 73.5 ft-lbs, and this mate rial could sustain a reduction in shelf energy of 32% before reaching 50 ft-lbs.

Using the dashed curve drawn through the data point for Plate B2002-3, it is predicted that the Cv shelf energy of Plate B2002-3 will reach 50 ft-lbs at a fluence of about 9. 5 x 1018 (E> 1 MeV). This corresponds to approximately 35

+

U

+ F :H HIM

+Hfl77 !IIMFH+Hll _Lft* I C7 c I ITI iff$a

-+4+++H- 1 it (U

r)

+ 3

1. P T- 111

__Zj Plat

.. 41" P4444---:

+ ..... .........

_4+ 111111111 1111i di I :

4+R1H-H+/-H+/-ELH+/-h1 ffiffl 2z 17 2x10 1 8 1018 4 6 8 1019 4 4 6 2

Neutron Fluence, n/cm (E> 1 MeV)

FIGURE 8. DEPENDENCE OF C v SHELF ENERGY ON NEUTRON FLUENCE, INDIAN POINT UNIT NO. 2

19 effective full power years (EFPY) of operation at the vessel I. D. and 32 EFPY (design life) at the vessel 1/4T position. This projection will be refined as additional surveillance capsules are removed.

A similar approach can be taken to estimate the increase in RTNDT as a function of reactor power generation. Again, a method for estimating shifts in RTNDT is given in Regulatory Guide 1. 99,* Revision 1. Figure 9 compares the Indian Point Unit No. 2 surveillance data on the three plate materials to selected portions of Figure I of Regulatory Guide 1. 99, Re vision 1. The results indicate that the measured shifts in RTNDT are higher than predicted by Regulatory Guide 1. 99, Revision 1.

The predicted shifts in RTNDT for the Indian Point Unit No. 2 reac tor pressure vessel obtained from Figure 9 are summarized in Table XI.

The values predicted at the 1/4T and 3/4T after 5 EFPY are used to de velop heatup and cooldown limit curves to meet the requirements of Ap pendix G to Section III of the ASME Code.

These projections for Cv shelf energy reductions and RTNDT shifts are based on one data point each for the three vessel plates but, more im portantly, no data on the weld metal. Even if it is assumed that the weld metal is more sensitive than Plate B2002-3 to radiation embrittlement, it will require a number of years of operation for the weld metal to become controlling because the initial RTNDT of the weld metal is F(1 3 ),

nearly 100°F below that of Plate B2002-3.

600 400 44

')

0) 2001 4)

U 0

4) iiiiiiiiiiii iiiiiiiiiiiii iiiiiiiiii I 111 i I . lI[II 1 11 1lllt1l1l I II II lIl it1 i 4) 4)

1001 I III

.1 li~ I i l I i i LI~f[{~T1 ~

U ttlflhll II 0

oo 4)

LU; . .

17 1018 8 1019 2x 10 4 6 8 4 6 4 6 Neutron Fluence, n/cm 2 (E > 1 MeV)

FIGURE 9. EFFECT OF NEUTRON FLUENCE'ON RTNDT SHIFT, INDIAN POINT UNIT NO. 2

TABLE XI Projected Shifts in RTNDT for Indian Point Unit No. 2 5 EFPY(a) 32 EFPY Location Fluence(b) ARTNDT(C) Fluence ARTNDT Pressure Vessel I. D. 2.5 x 1018 145 OF 1.6 x 1019 310OF Pressure Vessel 1/4T 1.5 x 1018 11 0 OF 9.6 x 1018 280°F Pressure Vessel 3/4T 3.7 x 1017 50°F 2.4 x 1018 140°F (a) 1 EFPY = 1, 006, 700 MWDt (b) Neutrons/cm z , E > 1 MeV (c) ARTNDT is maximum for Plate B2002-3. Refer to Figure 9.

It is anticipated that the reliability of the trend curves will be im proved as more surveillance data becomes available and a better under standing of the factors affecting radiation embrittlement has been achieved. As an example of the latter, Mr. E. C. Biemiller of Combus tion Engineering, in a paper given at the ASTM E 10 Symposium on Effects of Radiation on Structural Materials in St. Louis, May 4-6, 1976, indicated that a parameter of (% Ni + % Si) + (% Mo + % Cr + % Mn)4may explain the variation in radiation embrittlement observed in ferritic materials of nom inally the same copper content. Also, the Metal Properties Council is de veloping new radiation damage curves that will be based on more data than those currently in use.

The Indian Point Unit No. 2 reactor vessel surveillance program schedule proposed by Westinghouse (13 ) is summarized in Table XII. It has been organized to satisfy Appendix H of 10CFR50 as closely as possible.

There are seven additional capsules in the vessel, three of which contain weld metal specimens. There is no reason to consider changing the basic concept of the removal schedule until a capsule containing weld metal has been removed and tested. However, consideration may be given to the removal of Capsule Y instead of Capsule S at the time of the second region replacement because Capsule Y contains Charpy V-notch specimens from Plate B2002-3 and Capsule S does not.

TABLE XII Proposed Reactor Vessel Surveillance Capsule Schedule Indian Point Unit No. 2 Capsule Caps ule Material Expos ure Ident. Type(a) Content(b) Time S II. 1, W, H Second Region Replacement T I 1, 2, 3 Removed 1976 U I 1, 2, 3 Spare V 2, W, H 10 Years I5 W 1, 2, 3 Spare I

X 1, 2, 3 Spare I

Y 3, W, H Spare I

z 1, 2, 3 Fourth Region Replacement (a) Type I contains all three vessel plates. Type II contains weld metal, HAZ and one vessel plate.

(b) Material Code:

I - Plate BZ002-l; 2 - Plate B2002-2; 3 - Plate B2002-3; W - Weld Metal; H - HAZ

VI. HEATUP AND COOLDOWN LIMIT CURVES FOR NORMAL OPERATION OF INDIAN POINT UNIT NO. 2 Indian Point Unit No. 2 is a 2758 Mwt pressurized water reactor operated by Consolidated Edison Company. The unit has been provided with a reactor vessel material surveillance program as required by 10CFR50, Appendix H.

The first surveillance capsule (Capsule T) was removed during the 1976 refuelling outage. This capsule was tested by Southwest Research Institute, the results being described in the earlier sections of this report.

In summary, these results indicate that:

(1) The RTNDT of the plate materials in Capsule T increased a maximum of 140 *F as a result of exposure to a neutron fluence of 2. 02 x 1018 neutrons/cm 2 (E > 1 MeV).

(2) Based on a ratio of 2.9 between the fast neutron flux at the Capsule T location and the maximum incident on the vessel wall, the ves sel wall fluence was 6.97 x 1017 neutrons/cm 2 (E > I MeV) at the time of removal of Capsule T.

(3) The maximum shift in RTNDT after 5 effective full power years (EFPY) of operation was predicted to be 105 'F at the 1/4T and 50 'F at the 3/4T vessel wall locations, as controlled by Plate B2002-3. After 32 EFPY, the corresponding values are predicted to be 270 'F and 130 'F, respectively.

The Unit No. 2 heatup and cooldown limit curves for 5 EFPY have been computed on the basis of the RTNDT of Plate B2002-3 because it is anticipated that the RTNDT of reactor vessel beltline material will be highest for Plate B2002-3 at least through that time period. The pro cedures employed by SwRI are described in Appendix B.

The values of RTNDT for the beltline regions of Indian Point Unit No. 2 were derived from (1) the surveillance program test results, (2) ratios of fast flux at the capsule location to the maximum fast flux at the 1/4T and 3/4T locations in the vessel wall, (3) Regulatory Guide 1.99, Revision 1, trend curves of increase in RTNDT as a function of neutron fluence (E > 1 MeV), and (4) the initial RTNDT of the primary system materials.

The initial value of RTNDT for the Indian Point Unit No. 2 vessel was reported to be 60°F based on Plates B2002-1 and B2002-3. (21) As Plate B2002-3 had the highest initial Charpy V-notch transition tempera ture( 1 3 ) and the greatest radiation-induced shift in transition temperature, the unirradiated and irradiated data for that material were reduced by 35%

and replotted in Figure 10 to confirm the shift in RTNDT caused by the Capsule T irradiation. The unirradiated RTNDT (defined as 60°F less than the 50 ft-lb Charpy V-notch temperature) would then be 50 0 F. How ever, the 60°F figure given in the Indian Point Unit No. 2 Technical Speci fications was used as the initial RTNDT in this analysis.

iii, I i.

UU I I -I LL fil 4 80 0003 40

~~,60 T I L F_

' 60 r 4-41 oorl Ir - - - - - -

40 1+ 41 -. 1.

-500 5 10 10 00 50 0035040 Tempert Te1.°F FIGURE 10. ESTIMATED TRANSVERSE CHARPY V-NOTCH PROPERTIES OF PLATE B2002-3

The following pressure vessel constants were employed as input data in this analysis:

Vessel Inner Radius, r i 86. 50 in.

Vessel Outer Radius, ro 95.34 in.

Operating Pressure, Po 2235 psig Initial Temperature, To 70 OF Final Temperature, Tf 550°F Effective Coolant Flow Rate, Q = 136.3 x 106 lbm/hr Effective Flow Area, A 26. 719 ft 2 Effective Hydraulic Diameter, D = 15. 051 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 computed for cooldown rates of 0 0 F/hr (steady state), 20°F/hr, 60°F/hr, and 100°F/hr. The ZO0 F/hr curve would apply to cooldown rates up to 20°F/hr; the 60°F/hr curve would apply to rates from 20°F to 60°F/hr; the 100 F/hr curve would apply to rates from 60°F/hr to 100OF/hr.

The Unit No. 2 heatup and cooldown curves for up to 5 EFPY are given in Figures 1], 12 and 13.

2600 + +H'. f, t

-H+/-t7 T!4+

i HHHHH+ 7777

H+/-g +

+44 4FF4 HHHH1.1 M Hd+H M -7 W 2400 E4 -+ +

4-H+H4+++++4-_i .... ..

I +4i+/-

+ T 4- .....

TZ, ii H A 2200 jE -- rT--r

+H+ -- 4 -H4 2000 +Ff+

-b" HIT

+Hi HH1 1800 Lk! H 11k+

HHH HHH

• -4 +

I' l l I TI T 1,

.1600 I II i I ILI. I t it Q)

-A r 4+ A1111114 -T44-H111-IM 1400 P., 44 1. ITZT-#-#

T- H H

_0 FHHHHHHH! jo +++--f++ -111111 a, 1200 4+-Hii

1. • I * # - ----- H- +-+ + +--t::* t Till I I ITT

-+H- 44- i Ii i Hiiii TIT;

$ 1 [Till I I ITT H iiiiiii 1000 -+-4+

H+ F .........

-'oFF -H+ -- r-4+--

800 IT

+ H! i I H -0 +/-t+/-

4+-- -T

-Hiiiiii WT- HW H Till Ill 600 r-rr rr-rrr

+F+ -i++

HHH RUH: --

400

+H+ II Hl

1. t#

200

-++

-H -H+.. qT 0 IL 100 150 200 250 300 350 400 50 Indicated Temperature, deg F FIGURE 11. INDIAN POINT UNIT NO. 2 REACTOR COOLANT HEATUP LIMITATIONS APPLICABLE FOR PERIODS UP TO 5 EFFECTIVE FULL POWER YEARS

2600 2400 2200 2000 1800 #

tx

. 1600 1400 1200

-4 1000 800 600 400 200 0

50 150 200 2'50 300 350 400 Indicated Temperature, deg F FIGURE 12. INDIAN POINT UNIT NO. 2 REACTOR COOLANT COOLDOWN LIMITATIONS APPLI CABLE FOR PERIODS UP TO 5 EFFECTIVE FULL POWER YEARS

2600 2400 . . . . . . . . . . .r ... ....

'I-r-tr 2200 2000 1800 1600 C

4r+- L J~---- ~ I i l I +/- jzt: 4 -

1400

+-+--~--i-

+ sin si----- ----- s------

~ r l

~ ~

m t~ f P1 1200 03 11000 800 600 400 200 0

50 200 250 300 350 400 Indicated Temperature, deg F FIGURE 13. INDIAN POINT UNIT NO. 2 INSERVICE LEAK TEST LIMITATIONS APPLICABLE FOR PERIODS UP TO 5 EFFECTIVE FULL POWER YEARS

VII. REFERENCES

1. Title 10, Code of Federal Regulations, Part 50, "Licensing of Production 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 Determine Nil-Ductility Transition Temperature of Fer ritic Steels," 1975 Annual Book of ASTM Standards.

4. 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 Develop ment, U.S. Nuclear Regulatory Commission, April 1977.

8. Comments on Regulatory Guide 1. 99, Westinghouse Electric Cor poration, Obtained from NRC Public Document Room, Washington, D.C.

9. Position on Regulatory Guide 1. 99, Combustion Engineering Power Systems, 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 Frac ture 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. "Indian Point Unit No. 2 Reactor Vessel Radiation Surveillance Program," WCAP-7323, May 1969.

14. "Indian Point Unit No. 3 Reactor Vessel Radiation Surveillance Program," WCAP-8475, January 1975.

15. Letter, Westinghouse to Con Edison, May 16, 1975.

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

17. ASTM E 322, "Standard Method for Spectrochemical Analysis of Low Alloy Steels and Cast Irons Using an X-ray Fluorescence Spectrometer, " 1974 Annual Book of ASTM Standards.

18. 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.

19. Telecon, E. B. Norris to Ken Hogue (NRC Staff) January 19, 1977.

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

21. Indian Point Unit No. 2 Technical Specifications, Amendment No.

28, dated April 26, 1976.

APPENDIX A TENSILE TEST RECORDS A-1

Southwcst RTcsearch lnstif "c partmeint of Material: Engi,..er ing TENSILE TEST DATA SHEET T,,t No. T- -___ Est. U. T.S. psi Projcct No. " "

Spec. No. / Initial G. L. _ Z-_ Machine No. '/

Temp. -F Initial Dia. in._ Date S t ra i n Rate , / _-___.... Initial Thickness in. Initial Area,_

/

Initial \Wid th In.

Top Temp. __. __. ________ Maximum Load . 0 lb.

Bottor Tenmp. 0, 2016 Offset Load S% lb.

1 Final G.L. /, _._,in. 0.21 Offset Load lb.

FiiL D r aia./ in. Upper Yield Point lb.

I Areca ,2 . in..

.2 MaxirnLm Load U.T.S. = i Initial Area

o. zOffsc t Load "7/

c Y.S, initial Area 0.2% Offset Load Initial Area _psi U3 ppere ield Poirt Jpper Y.S. Initiaj Area psi II Fina) G. J - Initial (1. ,.

Initial G. L. X 100 i I~nitil Ara F la A

~. __ ___________________j0 7 7

.7 A-2

+

tLUMLL llt I I , , .1 ...I...........

~Zt~K2 I PdTM W, wl IF 4

&i , II , i , , i r rrr , i i , t, , i I:5-E Ii

-,44F

! -H:4# tI

+/-41 1 - 7 r i 4 j H-14 F

7L

-- l- .4-4 Lf-4 Lf4 aTl

=7.

.. qf -H: ++ -+

iH++

T

+

4-1 L: +

H,j - r -- t

-41  : 4 H- - -; r jj14 +

+

7 rl

---H-4+

-T Tl-i

+

+

4-1W

-H

4. + T-T ,iMl' I-! d-4 Lit~40M

.14-

1 - _ 4 ~+; ~

.Oe~5ore) ofr '01-0 02 5 tJ IA)O A-3

Southwcst Rcscarch Insti' 'e partmcnt of Materials Enp..ccring TENSILE TEST DATA S]1EET T -t No. T -- ___,_ Est. U.T.S. psi JProjcct No.0o-v  ?/

Spec. No. A-i Initial G. L. ___n: Machine No.1?/Zlzo Temp. -A,c15 /7 F Initial Dia. ,i n. Date I-2Qe/-2" Strain Rate . Initial Thickness in. Initial Arcaey-Initial Width in.

Top Timp. t F Maximum Load 4 2 5 lb.

Bottom Tcrnp. m b O,2% Offset Load Z 200 lb.

Final G.L. /_-2 7 in. 0.Z% Offset Load lb.

i a. in. Upper Yield Point lb.

T[ I Area _ _in.

.U.T.S. ==F Maximum Load ,

Initial Area - eq ,si 0 .0.2t1Offset Loid D

S ysl Initial Arca 2?%Offset Load 10.250 Y.S. = 0.Initial Area psi Upper Yield Point U~pper Y.S. = Initial Area Final G. ,. - Initial G.L.

I. ....... I: ....

if,?. "WT'. . '*" I ('iA .. -!f2 7 U , .Initial G. L.

Initial Arvn.a - in-a Arva A-4

A-5 Southwest Research Instif 'c particnt of Material,; E:'ngi.,ocring TENSIL)E TEST DA'fA SHEET Est.U.T.S. psi Project No .- 53/

Initial G. L. -: Machine No a Spec. No. J_

Temp. .76,(2 OF Initial Dia. In.

i Date I;'-a Strain Rate , Initial Thickness in. Initial Ae; '

Initial Width in.

'9 Top Temp. 5S 0 .F Maximum Load 6I/39o lb.

Bottom Temp. OF 0,Z% Offset Load ,cc'O lb.

Final G. L. /, /c in. 0.-2% Offset Load lb.

Zinlk Dia. ,/. in. Upper Yield Point lb.

F1 I Area c/, / in.Z UTS U.T.S.= Maximulm Load~

psi Initial Area

0. vy Y.s. -=0. Z%' Offset Load Initial Area ,7? ps!

0. Z%Y.. S. = 0. 2 i" Offset Load psi Initial Area UppCr . Upper Yield Point _Psi pnitial Area Final G.L,. - Initial G. L.

Initial G. L.

( Initial Ar.a - Fihal Area "

hiii Area i

-_____________--_ -__0 A-6

-i-H, '-~ L A-

- I 11521 JI 7i -H-

++

i.i+-HH r---

-]Tq

-1 P71T 11 i 1!-H- 1tt i -r-, I -!!!- -L -

- 4 II'!I F F' I 14X Hh+. L~4~I~5~IL -I-4-~'

-T ll4

.1-mmil 1,

~4 IT-,

I 11#41L 1111!111,1Ill 11111111 111 #IL#J#1-11:44-1 I. U .L.,

' OIS7-.4010 ,0 2-!5-A-7

APPENDIX B PROCEDURE FOR THE GENERATION OF ALLOWABLE PRESSURE-TEMPERATURE LIMIT CURVES FOR NUCLEAR POWER PLANT REACTOR VESSELS B-1

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. Background The basic parameter used to determine safe vessel operational conditions is the stress intensity factor, K I , which is a function of the stress state and flaw configuration. The K I corresponding to membrane tens ion is given by KIm = Mm om where Mr n is the membrane stress correction factor for the postulated flaw and a-m the membrane stress. Likewise, K I corresponding to bend ing is given by KIb = Mb (b where Mb is the bending stress correction factor and 0-b is the bending stress. For vessel section thickness of 4 to lZ inches, the maximum B-Z

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. 1).

These curves are a function of the stress ratio parameter O/c-y, where 0 y

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 XI, "Rules for Inservice Inspection of Nu clear Power Plant Components, " Article A-3000.

The Code specifies the minimum K I 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 10 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 Relationships

1. Inservice Leak and Hydrostatic Test During performance of inservice leak and hydrostatic tests, the reference stress intensity factor, KIR, must always be greater than B-3

1.0 1.2 1.4 1.6 1.8 2.0 2.2 2.4 2.6 2.8 3.0 3.2 3.4 3.6 3.8 4.0 FU TR CKNSESS (IN.)

FIGURE 1. STRESS CORRECTION FACTOR B-4

130 ]-- ' REFERENCE STRESS INTLNS!IY FACTO,------/-

120 T. TEMPERATUR AT W1HICH ,<

10 .... . IS PERlITTE),0F 10 RTaDT REFERENCE NIL-DUCEILlTY 00 . TEMPERATURE 70 30 Go l _ _ _

40 30 "

20 10 . . . .. 1-~-~ - --

o L 1 __ .

I .,_ ___

240

-240 -200 -I0O -120 -80 -40 0 40 80 120 IGO 200 TEMPERATURE RELATIVE TO RTHOT, (T-RT.DT), FAHRE ,9EIT DEGREES FI'?IURE 2. R.FFIERENCE STRESS INTENSITY FACTOR B-5

1. 5 times the K, caused by pressure, thus

1. 5 Kip < KIR (4) or

< (5) 1.5 Mm Tm KIR" For a cylinder with inner radius r i and outer radius r o ,

the stress distribution due to internal pressure is given by S ri 2 " o r 6 cr(r) ( - riZ)( r With 1/4T flaws possible at both inner and outer radial locations, i.e.,

at rl/ 4 =r i + 1/4(r O - ri) and r 3 / 4 = r i + 3/4(r o - ri), the maximum stress will occur at the inner flaw location, thus 2 (7)4i)

T ( ri [ro2 +( (/4ro amax= Po(roZ--riZ) ( 1 /4ro+3/ri)Z ] (7)

With the operation pressure known, i.e., Po, we deter mine the minimum coolant temperature that will satisfy Equation (4) by evaluating KIR = .5 Mm omax (8) 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 = RTNDT(1/4T)- 160. +68. 9988 in K - 6 7 L 223-77 " ] (

9 B-6

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

2. Heatup and Cooldown Operations At all times during heatup and cooldown operations, the ref erence stress intensity factor, KIR , must always be greater than the sum of 2 times the Kip caused by pressure and the Kit caused by thermal gra dients, thus 2.0 Kip + 1.0 Kit < KIR (10) or 2.0 Mm O-max = KIR - Kit (11) where 0rmax is the maximum allowable stress due to internal pressure, 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 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 B-7

OUTER RADIUS 3/4T 1/4T INNER RADIUS Pressure stress distribution Thermal stress distribution (a) Heatup OUTER RADIUS 314T

+

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 mnum 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 Pmax(l /4T) 12)

(2KIR 2

Mm -2r (i/4r° + 3/4ri)2 )

where Mm is determined from the curves in Figure l and KIR 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-/oy).

At the 3/4T location, the maximum pressure is determined from Equation (11) as P (3/4T) KIR - Kit (13) max (M,ri Vro2 + (l/4ri+ 3/4ro)\

,Znkr2rZA (l/4ri+ 3/4r0 )Z )

where KIR is obtained from Equation (2) using the material temperature and RTNDT at the 3/4T location and Kit 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 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 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 cooldown rate, the maximum pressure is evalu ated at the 1/4T location from KIR - Kit Pmax(1 /4T) = M (ri 2 Z)(oi;2/4 1/4o) (14) 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 B-10

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 Operation 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 Analysis 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 pcTt - K [(I/r)Tr + Trrl 0 (15) subject to initial condition T(r, 0) = T O s (16) and boundary conditions

-KTr(ri,t) = hFTc(t) - T(rit)] (17)

B-ll

and Tr(ro,t) = 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, T o the initial coolant temperature, T(rt) 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 Ar apart, across the vessel are:

for 1 <n < N Tt+At ArK Ar ] t n =1- pc(Ar) Z ( +rn n for n = 1 t+At F1 pc(Ar)

AtK 2 (i+- Ar cAth-(A pc(Ar) r I

+ (p+-) T +- t (21) pc(Ar) rl 2 K c B-12

and for n = N t+At AtK KAt TN =[ p(Ar)2 N TNN + K 2 t TN-N pc(Ar) (22)

For stability in the finite difference operation, we must choose At for a given Ar such that both AtK +Ar)

A2 (2+ < 1 (23) pc(Ar)Z r1 and AtK +Ar) Ath p(Ar) 2 (1 + < 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 (K/pc) called the thermal diffusivity and K, 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-OF) (25)

B-13

and 2 /HR) (26) k(T) (K/pc) = 0.6942 - 0.000432

• T (FT 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*

) v0 8 /Do'2 (27) h(T)= 170(1+1072*T- 10-5*T-where v is in ft/sec, D in inches, the temperature is in °F, and h is in Btu/hr-ftz 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 temperatures up to 600°F. The mean velocity of the coolant, v, is generally given in terms of the effective coolant flow rate Q 2 Given the relationship (Lbm/hr) and effective flow area A (ft ).

p(T) 62.93 - 0.48x 10 - 2 *T - 0.46 x 10 -4

• T 2 (28),

for the density of water as a function of temperature, the mean velocity of the coolant is obtained from v Q/(3600 *pT) *A) . (29)

Glasstone, S., Principles of Nuclear Reactor Engineering, D. Van Nostrand Co. , Inc. , New Jersey, pp. 667-668, 1960.

B-14

The thermal stress distribution is calculated from rT(r, t) = -a 1 r +1 r 2 +ri 2 rro

- r?-21 T(r t) rdr - T(r, t)+ r2rZ - ri2 j T(r, t) rdr (30) ri 0r i 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 Theory of Elasticity 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 x and E are evaluated as a function of the average temperature across the vessel 2 rro Tavg - (to2 ri2)J T(r)rdr (31) ri 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 T (32) and the dependence of Young's modulus on temperature is taken to be E(T) = 27. 9142 + 2. 5782 x 10 - 4

• T - 6.5723 x 10-6

• T 2 (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

stress distribution is given by r o M(t). = b O-T (r, t) rdr ri 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 stress distribution we have that T'max Mc (35) where a is the maximum outer fiber stress, c the distance from the neutral axis, taken to be (r o - ri)/Z, and I the section area moment of inertia which is given by bh 3 b(ro - ri)3 12 12 (36)

Combining these expressions results in the equivalent linear stress due to thermal gradients C'max = rbt =(r o 6 ri) o (rT (r, 0) rdr (37) r The thermal stress intensity factor Kit is then defined as Kit = Mb a-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 B-16

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, Tavg, by T= T"Tavg (39)

(T - Tavg)max 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. Example Calculations The following example is based on a reactor vessel with the follow ing characteristics:

Inner Radius 82. 00 in. (ri)

Outer Radius 90. 00 in. (r 0 )

Operating Pressure 2250 psig (Po)

B-17

,OUTER WALL 1.0 0.8 0.6 0.4 0.2

-1.0 1.0 -1.0 1.0 INNER WALL Normalized temperature Normalized stress

.distribution ( AT / ATmax) distribution ( 0/max Figure 4. Typical Normalized Temperature and Stress Distribution During Heatup B-18

Initial Temperature = 70°F (T 0 )

Final Temperature = 550°F (Tf)

Effective Coolant Flow Rate = 100 x 10 6 Lbm/hr (Q)

Effective Flow Area = 20.00 ft 2 (A)

Effective Hydraulic Diameter = 10. 00 in. (D)

RTNDT (1/4T) = 200°F RTNDT (3/4T) = 140°F In the thermal stress analysis Z 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 Kit are required to compute allowable pressure via Equations (13) and (14). The material temperatures versus coolant temperature during the 100oF/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 temperature.

B-19

200 160 -RTNT 1/4T) 200OF 120 108 80 40 01 250 300 350 400 50 100 150 200 TEMPERATURE (F)

Indexed to RTNDT( 114T)

Figure 5. Reference Stress Intensity Factor as a Function of Temperature

400 -

- 100°F/HR HEATU P ( 3/4T Location ) fle 0-0

--- IO0F/HR COOLDOWN ( 1/4T Location )

300 -

2001-100 .00*

0000 100 100 150 200 250 300 350 COOLANT TEMPERATURE (OF)

Figure 6. Vessel Temperature-at 1/4T and 3/4T Locations as a Function of Coolant Temperature

10 8 - 0000 6

- 100OF IHR HEATUP ( 3/4T Location)

-- 100 OF/HR COOLDOWN (1/4 Location) 4 2

01 200 250 300 350 50 100 150 COOLANT TEMPERATURE (OF)

Thermal Stress Intensity Factor at 3/4T and 1/4T Locations as a Function of Coolant Temperature Figure 7.

Figures 8 and 9 demonstrate the construction of the allowable com posite pressure and temperature curves for the 100OF/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 LIMIT 2400 2000 COMPOSITE CURVE'-100°F/HR HEATUP

( Margins of +10°F and -60 psig for instrument error) 1600 STEADY STATE CRITI CALITY 1200 LIMIT HEATUP 400 100 350 400 INDICATED TEMPERATURE ('F)

Figure 8. Pressure-Temperature Curves for I00°F/Hr Heatup

2400 2000 COMPOSITE CURVE -100 0 F/HR COOLDOWN (Margins of +10°F and

-60 psig for instrument error 1600 1200 COO LDOW N 800 STEADY STATE

- -W-ol 400 150 200 250 300 INDICATED TEMPERATURE (.F)

Figure 9. Pressure-Temperature Curves for 100°F/Hr Cooldown