ML20215E186
| ML20215E186 | |
| Person / Time | |
|---|---|
| Site: | Turkey Point  |
| Issue date: | 09/30/1976 |
| From: | FLORIDA POWER & LIGHT CO. |
| To: | |
| Shared Package | |
| ML20215E175 | List:
|
| References | |
| FOIA-87-286 NUDOCS 8706190285 | |
| Download: ML20215E186 (241) | |
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REACTOR VESSEL MATERIAL SURVEILLANCE ANil fRACTUREMECHANICSEVALUATIONSFOR,h%)
TURKEYPOINTUNITNOS,3AND14 g. W 1. D' .
NUCLEAR POWER PLANTS ,,w,- . g '.
SUBMITTED BY i; ,
THE FLORIDA POWER & LIGHT COMPANY TO THE NUCLEAR REGULATORY COMMISSION 4
BETHESDA, MARYLAND C
SEPTEMBER-1976
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ANALYSIS OF CAPSULE T FROM THE FLORIDA POWER AND LIGHT COMPANY I j TURKEY POINT UNIT NO, 3 REACTOR VESSEL RADIATION -
I SURVEILLANCE PROGRAM
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' THE CONTENTS OF THE REPORT PRESENTED HEREIN IS MAINLY FROM WESTINGHOUSR g ECTRIC CORPORATION'S WCAP-56) .
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TABLE OF C0flTENTS SECTION TITLE PAGE 1
SUMMARY
OF RESULTS 1-1 2 INTRODUCTION 2-1 3 BACt'9ROUND 3-1 4 DESCRIPTION-0F PROGRAM 4-1 5 TESTING OF SPECIMENS FROM CAPSULE T 5-1 5.1. CHARPY V-NOTCH IMPACT TEST RESULTS 5-2 5-2. TENSILE TEST RESULTS 5-9 5-3. WEDGE' OPENING LOADING (WOL) TESTS 5-9 6 DOSIMETRY ANALYSIS 6-1 6-1. FAST NEUTRON FLUX MONITORS 6-1
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6-2. ANALYTICAL METHODS 6-2 6-3. RESULTS OF ANALYSIS 6-5 6-4. DISCUSSION OF RESULTS 6-5 7 REFERENCES 7-1 APPENDIX ADDITIONAL CURVES; COMMENTS ON THE A
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LIST OF TABLES Table
/ Title Pago 41 Chemistry and Heat Treatment of Material Representing the Core Region Shell Forgins and Weld Metal from the Turkey Point Unit No. 3 Reactor Vessel 42 4-4 Chemistry and Heat Treatment of Surveillance Material Representing 6 inch thick A3028 ASTM Correlation
- Monitor Material 51 45 Charpy V Notch Impact Data for the Turkey Point Unit No. 3 Pressure Vessel Shell Forging 123P461VA 1 Irra-diated at 550*F, Fluence 5.68 x 1018 n/cm2 (E > 1 Mev) 52 53' Charpy V Notch impact Data for the Turkey Point Unit No. 3 Pressure Vessel Weld Metal Irradiated at 550*F, i Fluence 5.68 x 1018 n/cm2 (E > 1 Mev) 53 5-3 Charpy V Notch impact Data for the Turkey Point Unit No. 3 Pressure Vessel Wald Heat Affected Zone Metal Irradiated at 550*F, Fluence 5.68 x 1018 n/cm2 (E > 1 Mev) 5-4 54 Charpy V Notch Impact Data for the Turkey Point Unit No. 3 ASTM SA302 Grade B Correlation Monitor Material Irradiated at 550*F, Fluence 5.68 x 1018 n/cm2 (E > 1 Mev) 5-5 5-4 The Effect of 550*F Irradiation at 5.68 x 1018 n/cm2 (E > 1 Mov) on the Notch Toughness Properties of the Turkey Point Unit No. 3 Reactor Vessel Impact Test Specimens
(~ 56 5-9 U Irradiated Tensile Properties for the Turkey Point Unit No. 3 Pressure Vessel Shelf Forging and Weld Metal 57 Fluence 5.68 x 1018 n/cm2 (E > 1 Mev) 5 10 Irradiated Fracture Toughries: Properties for the Turkey Point Unit No. 3 Pressure Vessel Shell Forging and Weld Metal, Fluence 5.68 x 1013 n/cm2 (E > 1 Mov) 61 5 17 Results of Fast Neutron Oc:simetry Capsule T 62 64 Calculated Relative Neutrora Spectrum at Capsule T Location 67 63 Pressure Vessel Fast Neutron Exposure Based on Dosimetry Measurements and DOT Calculations 69 L
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LIST OF ILLUSTRATIONS Figure Title Page 41 Arrangement of Surveillance Capsules in the Turkey Point Unit No. 3 Reactor Vessel (Lead Factors for the Capsules are Shown in Parentheses) 4-2 Surveillance Capsule Test Specimens 42 4-3 43 Capsule T Schematic Showing Location of Specimens, Thermal Monitors and Dosimeters 5-1 4-6 Charpy V-Notch Impact Energy for the Turkey Point Unit No. 3 Reactor Pressure Vessel Shell Forging 123P461VA 1 52 55 Charpy V Notch impact Energy for the Turkey Point Unit No. 3 Reactor Pressure Vessel Wold Metal 53 5-6 Charpy V Notch impact Energy for the Turkey Point Unit No. 3 Metal Reactor Pressure Vessel Weld Heat Affected-Zone 54 57
. Charpy V-Notch Impact Energy for SA302 Grade B ASTM Correlation Monitor Material 55 5-8 Tensile Properties for the Turkey Point Unit No. 3 Reactor Pressure Vessel Shell Forging 123P461VA 1 58 5-11
, Tensile Properties for the Turkey Point Unit No. 3 Reactor Pressure Vessel Weld Metal 57 5 12 Typical Stress Strain Curve for Tension Specimens 58 5-13 Load Displacement illustrating How K Curve for a Compact Tension Specimen 5-9 Od si Calculated 5-15 Type 1 Load Displacement Curve Representative of Gross Yielding at the Specimen Crack Tip 5 10 518 Type 2 Load Disolacement Curve Representative of an Intermediate Amount of Plasticity at the Specimen Crack Tip 5 11 5 19 Type 3 Load Displacement Curve Representative of Very Little Plasticity at the Specimen Crack Tip 5 20 IV l
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ANALYSIS OF CAPSULE T FROM THE FLORIDA POWER AND LIGHT COMPANY 4<
TURKEY POINT UNIT NO. 3 3 REACTOR VESSEL RADIATION -
SURVEILLANCE PROGRAM I
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THE CONTENTS MAINLY OF THE REPORT FROM WESTINGHOUS PRESENTED HEREIN ECTRIC CORPORATION S If WCAP- 6 .
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ANALYSIS OF CAPSULE T FROM THE FLORIDA POWER AND LIGHT COMPANY !
TURKEY POINT UNIT NO. 3 REACTOR VESSEL RADIATION SURVEILLANCE PROGRAM S. E. Yanichko J. H. Phillips S. L. Anderson December 1975 i
Work performed under Shop Order No MIP 23572 APPROVED: , h* h '
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. N. Chirigos, Manahr
_{ Structural Materials Engineering WESTINGHOUSE ELECTRIC dORPORATION Nuclear Energy Systems P. O. Box 355 Pittsburgh, Pennsylvania 15230
SECTION 1
SUMMARY
OF RESULTS The analysis of. the reactor vessel material contained in the first surveillance capsule from the Florida Power and Light Company, Turkey Point Unit No. 3, reactor pressure vessel led to the following conclusions:
e The capsule received an average fast fluence of 5.68 x 1018 neutrons /cm2 (E > 1 Mev).
The predicted fast fluence for the capsule at the end of first core cycle was 5.37 x 1018 neutrons /cm2 (E > 1 Mev).
e The fast fluence of 5.68 x 1018 n/cm2 (E > 1 Mov) resulted in a 190 F increase i the 50 ft Ib reference nil-ductility transition temperature (RTNOT) of the weld metal, which is the most limiting material in the core region of the reactor vessel. The inter-mediate pressure vessel shell forging (123P461VA 1) exhibited essentially a 0 F shift s
in the 50 ft Ib nilductility transition temperature (specimens oriented in the major working direction of the forgings). The weld heat-affected zone material also exhibited a 0*F shift in transition temperature.
s Based on a ratio of 2.48 between the fast flux at the surveillance capsule location to that at the vessel wall and an 80 percent load factor, the projected fast fluence which the Turkey Point Unit No. 3 reactor pressure vessel will receive after 40-calendar year operation is 6.65 x 1018 n/cm2 (E > 1 Mov). This fluence is approximately the same as the 6.30 x' 1019 n/cm fluence calculated for 40-year operation, a The projected shift in RTNDT of the weld metal after 40 calendar year operation is
(~ 382* and 330 'F at the vessel inside surface and the 1/4 thickness location respect-ively.
e The average upper shelf impact energy of the weld metal decreased from 64.5 to 58.5 ft Ib during the first core cycle.
e The irradiated properties of forging 123P461VA-1 and the weld metal are adequate to provide for continued safe operation of the Turkey Point Unit No. 3 power plant.
e The fracture toughness as measured dynamically on WOL specimens gave results well above the limits of the KIR curve in Appendix G of ASME Code Section Ill.
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SECTION 2 INTRODUCTION This report presents the results of the examination of Capsule T, !
the first capsule of the continuing surveillance program-which monitors the effects of neutron irradiation on the Florida Power 5 Light Company, Turkey Point Unit No. 3, reactor pressure vessel materials under actual operating conditions.
The surveillance prcgrsm for the Turkey Point Unit No. 3 reactor pressure vessel materials was designed and recommended by the Westinghouse Electric Corporation. A description of the surveillance program and the preirradiation mechanical properties of the reactor vessel materials are presented in WCAP-7656. (1)
The surveillance program was planned to cover the 40-year life of the reactor pressure vessel and was based on ASTM E-185-66, "Reccmmended Practice for Surve MaterialsinNuclearReactors.".gpanceTestsonStructural Westinghouse Nuclear Energy Systems personnel were contracted for the preparation of procedures for removing the first capsule from the reactor and its shipment to the Westinghouse Research and Development Laboratory, where the post irradiation mechanical testing of the Charpy V-notch impact, tension, and Wedge Opening Leading (WOL) surveillance specimens was performed.
This report summarizes testing and the post irradiation data
, obtained from the first material surveillance capsule (Capsule T) removed from the Turkey Point Unit No. 3 reactor vessel and discusses the analysis of these data.
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i SECTION 3 BACKGROUND The ability of the large steel pressure vessel containing the reactor cor to resist fracture constitutes an important factor in ensuring safety in the nucl .
beltline region of the reactor pressure vessel is the most critical region of is subjected to significant fast neutron bombardment. The overall effects of fa -
. diation on the mechanical properties of low alloy ferritic pressure vessel ste Class 2 (base material of the Unit No. 3 reactor pressure vessel beltline) a in the literature. Generally, low alloy ferritic materials show an increase in properties and a decrease in ductility and toughness under certain conditions of A method for performing analyses to guard against fast fracture in reactor been presented in " Protection Against Non-ductile Failure," Appendix G, to ASME Boiler and Pressure Vessel Code. The method utilizes fracture me based on the reference nilductility temperature, RT N DT-
. RT i NDT s defined as the greater of the drop weight nilductility transition temperatur per ASTM E 208) or the temperature 60 F less than the 50 ft Ib (and 35 m temperature as determined from Charpy specimens oriented normal to the ma of the material. The RT NDT of a given material is used to index that material to a reference stress intensity factor curve (K K lR curve) which appears in Appendix G of the ASME Code. The IR curve is a lower bound of dynamic, crack arrest, and static fracture toughness ed from several heats of pressure vessel steel. When a given material is indexed to IR curve,
.. allowable stress intensity factors can be obtained for this material as a functio .
( Allowable operating limits can then be determined utilizing these allowable s ors.
RT NDT, and in turn the operating limits of nuclear power plants, can be adjusted to acc
.* the effects of radiation on the reactor vessel material properties. The radiation em changes in mechanical properties of a given reactor pressure vessel steel can reactor surveillance program such as the Florida Power and Light Company, Turkey Po r No. 3. Reactor Vessel Radiation Surveillance Program,N
. in which a surveillance capsule is peri-S t
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odically removed from the operating nuclear reactor and the encapsulated specimens are testert.
The increase in the Charpy V. notch 50 ft Ib temperature (ARTNDT) due to irradiatint . . a'a.1 to the original RT NDT to adjust the RT f radiation embrittlement. This adjusted RT NDT or (RT i NOT NDT nitial +4RTNOT) is used to index the material to the KIR curve and in turn to set operating limits for the nuclear power plant which take into account the effects of irradiation on the reactor vessel materials.
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i SECTION 4 DESCRIPTION OF PROGRAM J
Eight surveillance capsules for monitoring the effects of neutron exposure on the Turkey Poin Unit No. 3 reactor pressure vessel core region material were inserted in the reactor vessel prio to initial plant startup. The eight capsules were positioned in the reactor vessel between the thermal shield and the vessel wall at locations shown in figure 41. The vertical center of the capsules is opposite the vertical center of the core.
Capsule T was removed after approximately 1 1/2 years of plant operation. ThIs capsule con-tained Charpy. V notch impact, tensile, and WOL specimens (figure 4-2) from the intermediate shell ring forging (heat 123P461VA 1), weld metal from the core region of the reactor vessel, and Charpy V notch specimens from weld heat affected-zone (HAZ) material. The capsule also contained Charpy V notch specimens from the 6 inch thick ASTM correlation' monitor material-(A302 Gr. B) furnished by the U. S. Steel Corporation. The chemistry and heat treatment of the surveillance material is presented in tables'41 and 4.2.
All test specimens were machined'from the 1/4 thickness location of the forging. Test specimens represent material taken at least one forging thickness from the quenched and of the forging.
All base metal Charpy V notch and tensile specimens were oriented with the longitudinal axis of
' the specimen parallel to the principal working (hoop) direction of the forgings. The WOL test specimens were machined with the simulated crack of the specimen perpendicular to the sur-faces and hoop direction of the forgings.
Charpy V-notch specimens from the weld metal were oriented with the longitudinal axis of the specimens transverse to the weld direction. Tensile specimens were oriented with the longitudinal axis of the specimen parallel to the weld.
Capsule T contained dosimeter wires of copper, nickel, and aluminum-cobalt (cadmium shielded and unshielded). In addition, cadmium shielded dosimeters of Np m and U m were contained in the- I capsule and located as shown in figure 43.
' Thermal monitors made from two low melting outectic alloys and sealed in Pyrex tubes were in.
ciudad in the capsule and were located as shown in figure 4 3. The two eutectic alloys and their melting points are: ;
2.5% Ag, 97.5% Pb Melting Point 579"F 1.75% Ag, 0.75% Sn, 97.5% Pb Melting Point 590*F 41
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Arrangement of Surveillance Capsules in the Turkey Point Unit No. 3 Reactor Vessel (Lead Factors for the Capsules are Shown in Parentheses) 4-2 9 .
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i TABLE 41 CHEMISTRY AND HEAT TREATMENT OF MATERIAL REPRESENTING TH CORE REGION SHELL FORGINGS AND WELD METAL FROM THE TURKE POINT UNIT NO. 3 REACTOR VESSEL CHEMICAL ANALYSES (PERCENT)
Intermediate Lower Shell Shell Element As Deposited 123P461VA 1 123S266VA 1 Weld Metal C 0.20 Mn 0.19/0.21 0.076 0.64/0.64 0.61/0.62 P 0.010 1.26 S
0.010 0.011
' 0.010 0.008 Si 0.26 O.018 Ni 0.20/0.19 0.66 0.70 0.68/0.66 Cr 0.40/0.39 0.57 0.38 0.14 V 0.02 0.02 0.002 Mo 0.62 0.58/0.59 0.42 Co 0.011/0.010 Cu 0.015/0.016 0.001 0.058 (0.07) " 0.079 (0.07) "
Sn 0.010 0.31 0.008 0.004 Zn 0.001 0.001 0.003 At 0.005 oi 0.005 0.015 Ng 0.003 0.003 Ti 0.001* 0.012 0.001* 0.001 Sb 0.001* l 0.001* 0.001 As 0.005* l 0.005* 0.005 8 0.003* !
0.003* 0.003*
Zr 0.001* 0.001* 0.001 )
HEAT TREATMENT i
intermediate and 1550*F l Lower Shell Forgings 13 hours1.50463e-4 days <br />0.00361 hours <br />2.149471e-5 weeks <br />4.9465e-6 months <br /> Water-quenched '
1210*F 18 hours2.083333e-4 days <br />0.005 hours <br />2.97619e-5 weeks <br />6.849e-6 months <br /> Air cooled 1125'F 10% hours Furnace-cooled to 600*F Weldment 1125*F 10Y. hours Furnace <ooled to 600"F >
- Not Detected. The number indicates the minimum hmit of detection.
- ' Copper Content Reported By Bethlehem Steel Co.
4-4
1 TABLE 4 2 CHEMISTRY AND HEAT TREATMENT OF SURVEILLANCE MATERIAL REPRESENTING 6 INCH THICK A302B ASTM CORRELATION MONITOR MATERIAL CHEMICAL ANALYSES (Percent)
C Mn P S Mo
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Si Cu Ni Cr 0.24 1.34 0.011 0.023 0.51 0.23 ~ 0.20 0.18 0 11 HEAT TREATMENT The six inch thick plate was charged into a furnace operating at 110*F heated at a maximum rate of 63*F per hour to 1650*F, held at temperature for 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br />, and water quenched to 300*F The plate was then recharged into a fur-nace operating at 700 to 750*F and heated at a maximum of 63*F per hour to 1200*F for 6 hours6.944444e-5 days <br />0.00167 hours <br />9.920635e-6 weeks <br />2.283e-6 months <br />.
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i SECTION 5 TESTING OF SPECIMENS FROM CAPSULE T j
~ The post irradiation mechanical testing of the Charpy V notch, tensile, and WOL specimens was performed at the Westinghouse Research and Development Laboratory with consultation-by Westinghouse Nuclear Energy Systems personnel.
Upon' receipt of the capsule at the Laboratory, the specimens and spacer blocks were care-fully removed, inspected for identification nurnber, and checked against the master list in WCAP-7656.N No discrepancies were found.
Examination of the two low melting (579*F and 590*F) eutecric alloys indicated no melting i of either type of thermal monitor. Based on this examination, the maximum temperature to which the test specimens were eAposed was less trien 579'F.
A TMI Model TM 52004H impact test machine was used to perform tests on the irradiated Charpy V-notch specimens. Before initiating tests on the irradiated Charpy V specimens, the accuracy of the impact machine was checked with a set of standard specimens obtained from '
the Army Material and Mechanics Research Conter in Watertown, Massachusetts. The results of the calibration testing showed that the machine was certified for Charpy V notch impact testing.
j The tensile tests were conducted on a screw-driven instron testing machine having a 20,000-lb capacity. A crosshead speed of 0.05 in./ min was used. The deformation of the specimen was measured using a strain gage extensometer. The extensometer was calibrated before testing with a Sheffield high magnification drum type extensor'neter calibrator.
Elevated temperature tensile tests were conducted using a split tube furnace. The specimens were held at temperature a minimum of 20 minutes to stabilize the temperature prior to testing. Temperature was monitored using a chromel-alumel thermocouple in contact with the upper and lower specimen grips which were clevis pin type. Temperature was controlled within 13*F.
The load-extension data were recorded on the testing machine strip chart. The yield st'rength, j ultimate tensile strength, and uniform elongation were determined from these charts. The !
reduction in area and total elongation were determined from specimen measurements.
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Dynamic fracture toughness tests were performed at loading rates between 2.95 x 410and 3.40 x 104 ksi VTrIlsec on the WOL specimens. A hydraulic test machine desyped by t.m
" Westinghouse Research and Development Laboratory was used to perform the tests. An LVOT attached to an extensometer which was mounted on the WOL specimen was used to measure displacement. Extensometer calibration was accomplished with a Sheffield drum-type calibrator.
Load and displacement outputs were recorded on individual recording channels of a Con-
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solidated Electro Dynamics (CEC FM) instrument tape recorder. Before performing the test calibration marks were recorded on individual recording channels. During testing, load and displacement were recorded on the same channel with their respective calibration marks.
Load versus time and displacement versus time traces were then displayed on an oscillosc and photographed along with the calibration marks that had been prerecorded. The tape was then _ replayed and the load and displacement signals were then cross plotted to produce a load versus displacement curve.
Elevated temperature tests were performed on the WOL specimens using a tube-type furnace.
A chromel alumel thermocouple attached to the specimen clevis was used to monitor the i specimen temperature. Temperature was controlled within 3*F. Prior to testing, all specimens were held at temperature a minimum of 30 minutes to stabilize the temperature.
- 51. CHARPY V NOTCH IMPACT TEST REST LTS The irradiated Charpy V specimens represented reactor pressure vessel beltline forging mate rial, weld and heat affected zone (HAZ) material, and the ASTM correlation monitor material.
The results are presented in tables 51 through 5-4 and figures 51 through 5-4. The unirra-j disted data are also shown in figures 5-1 through 5-4 for comparison with the post irradiation data. A summary of the increase in the 50 ft Ib "fix" transition temperature and the decrease in the upper shelf energy is presented in table 5 5. I The test results obtained on the vessel beltline shell ring material are presented in table 51 and figure 51. The data showed that the ring forging (heat 123P461VA 1) was insensitive to irradiation at 5.68 x 1018 n/cm2 based on the RTNDT shift of 0*F and no decrease in the upper shelf impact energy.
- The test results obtained on the weld material are presented in table 5 2 and figure 5 2. I The data shovved that the fluence of 5.68 x 1018 n/cm2 received by Capsule T increased the RT NDT of the weld metal specimens by 190"F and decreased the average upper shelf '
impact energy of the wetd metal from 64.5 to 58.5 ft Ib. The test results for the HAZ material are summarized in table 5 3 and figure 5-3. The data showed that the HAZ mate-rial was insensitive to irradiation at 5.68 x 1018 n/cm 2, 52 k
i TABLE 51 CHARPY V NOTCH IMPACT DATA FOR THE TURKEY POINT UNI
-PRESSURE VESSEL SHELL FORGING 123P461VA 1 IRRADIATED 550*F, FLUENCE 5.68 x 1018 n/cm2 (E > 1 Mev) .
- +
- ! men Test Temp Energy Lateral Expansion Number (*F) Shear '
(ftIb) (mils) (%)
PS2 50 27.0 17 PS6 20 2' 33.0 20 PS3 10 10 92.0 ;
61 40 P55 10
' 78.0 45 P54 '40 25 !
112.0 89 PS8 75 50 i 130.0 71 P57 140 100 ;
158.0 67 PSI 100 208 155.0 74 100 TABLE 5 2 CHARPY V NOTCH IMPACT DATA FOR THE TURKEY POINT UNIT PRESSURE VESSEL WELD METAL IRRADIATED AT 550'F, ,
FLUENCE 5.68 x 1018 n/cm2 (E > 1 Mov) f-;+:! men Test Temp Energy Lateral Expansion Number (*F) Shear (ftIb) (mils) (%)
W18 75 10.0 4.5 W22 10 150 28.0 13.0 30 W24 208 39.0 21.0 W20 210 40 40.0 26.0 W21 50 250 52.0 39.0 60 W23 250 53.0 38.0 W17 300 50 61.0 45.0 W19 100 300 56.0 49.0 100 l
5-3
i TABLE 53 CHARPY V NOTCH IMPACT DATA FOR THE TURKEY POINT UNIT NO.
PRESSURE VESSEL WELD HEAT AFFECTED ZONE METAL IRRADIAT AT 550*F, FLUENCE 5.68 x 1018 n/cm2 (E > 1 Mov) f-;+:! men Test Temp Energy Lateral Expansion Number (*F) Shear (ft lb) (mils) (%)
H2O 95
~ 20.0 8 H22 5 75 30.0 7 10 H24 60 6.0 0 1 H18 50
' 146.0 64 H21 80 40 167.0 76 100 N19 10 i
147.0 54 HD 100 75 187.0 79 100 H23 205 150.0 72 90 TABLE 5-4 CHARPY V. NOTCH IMPACT DATA FOR THE TURKEY POINT UNIT NO.{
ASTM SA302 GRADE B CORRELATION MONITOR MATERIAL !
IRRADIATED AT 550*F, FLUENCE 5.68 x 1018 n/cm2 (E > 1 Mov)
- r+:! men Test Temp Energy l
Number Lateral Expansion Shear
(*F) (ft Ib) (mils) (%) l i
R59 40 11.0 6 1
R63 3 75 25.0 17 5 R60 140 36.0 27 R61 30 140 34.0 24 15 R57 208 56.0 47 R64 75 210 54.0 35 70 R62 300 71.0 49 R58 100 300 73.0 56 100 l
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Charpy V Notch Impact Energy for the Turkey Point Unit No. 3 Reactor Pressure Vessel Shell Forging 123P461VA 1 55
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190 W 40 O O 5' ".681038 N/CH2 E lNEV 155' 20 O e 0 I I ! I l
-200 -100 0 100 200 300 400 500 TEMPERATURE (OF)
Figure b 2.
Charpy V. Notch Impact Energy for the Turkey Point Unit No. 3 Reactor Pressure Vessel Weld Metal 5-6
8971-4 220 -
2 ~
o g O Iso -
o 160 -
e 8 O O I40 -
o S
C 120
~
~
E 100 -
00 W
w
~
80 o UNIRRADIATED 60 $ IRRADIATED 550'F 5.68x1018 M/CH2 D INEV go _ O O O-20 - --- A ,8 O
0 9@ I I I I
-200 -100 0 100 200 300 400 TEMPERATURE (O) F Figure 5 3. Charpy V-Notch impact Energy for the Turkey Point Unit No. 3 Reactor Pressure Vessel Weld Heat.
Affected Zone Metal 57 .
8971-5 100 80 -
UNIRRADI A TED O
E s
60 -
a W
9 :
i 118*
5 40 -
5 '
9 i 0 0 82 IRRADI ATED 550 F Q
g 5.68 x 1018 M/CH2 g,lMEV 20 O
O I I I O I I !
-200 -100 0 '
100 200 300 400 TEMPERATURE (O) F 1
I I
I a
Figure 5 4. l Charpy V-Notch impact Energy for SA302 Grade B ASTM Correlation Monitor Material )
i l
I, 54
' TABLE 55
' THE EFFECT OF 550'F IRRADIATION AT 5.68 x 1018 n/cm2 (E > 1 Mev)
ON THE NOTCH TOUGHNESS PROPERTIES OF THE TURKEY POINT UNIT NO. 3 REACTOR VESSEL IMPACT TEST SPECIMENS .
Average Energy Absorption 50 ft Ib Temp (*F) {
Material at Full Shear (ft Ib)
Unitradiated irradiated oT Unirradiated Irradiated Mft Ib)
Forging 25 25 0 143 -143 0 123P461VA 1 i l
, Weld Metal 52 242 190 64.5 ~ 58.5 6 HAZ Metal 40 -40 0 166 ~166 0 Correlation 72 190 118- 78 72 6 Material The test results obtained on the 6 inch thick A3028 ASTM reference correlation material'are presented in table 5 4 and figure 5 4. The 118*F increase in RT NDT si 'in agreement with the reference trend band established by the Naval Research Laboratoriesl4I
, for this material.
- 52. TENSlLE TEST RESULTS ~
The results of the tensile tests are presented in table 5 6 and figures 5 5 and 5 6. Tests were performed on specimens from forging 123P461VA 1 at room temperature and 200'F,'and the weld metal at room temperature and 550'F. Irradiation increased the yield and tensile strength of the forging by less than 5 percent and the weld metal by approximately 25 per- I cent. A typical load displacement curve obtained for the tensile tests is shown in figure 5 7. ;
- 53. WEDGE OPENING LOADING (WOL) TESTS Of the various quantitative methods that have evolved for guarding against fast fracture within the past few years, linear-elastic frecture mechanics (LEFM) is the most universally accepted. With appropriate information in the related areas of material properties. stresses, and potential defects, the concepts and expressions of LEFM can be employed in established step by-step procedures which will ensure immunity from fast fracture. The material parameter 5-9
j l
i I
l TABLE 5-6 IRRADIATED TENSILE PROPERYlES FOR THE TURKEY POINT UNIT NO. 3 l PRESSURE VESSEL SHELL FORGING AND WELD METAL FLUENCE 5,68 x 1018 n/cm2 (E > 1 Mev)
Ultimate Red.
Test 0.2% Yield Tensile Uniform Total in Specimen Temp Strength Strength Elong. Elong. Area
. Material Number *F (psi) (psi) (%) (%) (%)
123P461VA 1 P11 80 65,750 89,400 14.2 29 71 123P461VA.1 P12 200 62,250 82,450 12.2
. 28 63 '
Weld Metal WS 200 90,800 104,900 14.5 29 54 Weld Metal W6 550 86,100 102,850 10.1 21 48 K
gg, a basic material property, is dependent upon the mechanical and metallurgical condition of
~
the material. Thus, neutron bombardment is known to influence the fracture resistance of pressure vessel grade steels.
1 The fracture toughness K;c can be determined by testing a fracture mechanics specimen - WOL, Compact Tension, bend or spin disk. Mager[5] has demonstrated the fracture toughness K ie of A508 Class 2 steel to be highly temperature dependent with a rapid increase in toughness near the NDTT (nil ductility transition temperature). Valid K ic data were obtainable only up to about 10*F, even when using 8 inch thick specimens. Furthermore, it has been demonstrated by Magerl6} that for the irradiation conditions of interest to the nuclear industry, specimens with thicknesses of 4 inches or greater must be tested to attain a fracture toughness K 100,000 psi (in.)l/2 ic Of As previously mentioned, because of reactor space limitations, reactor vessel surveillance cap-sules are restricted to approximately one inch in thickness. IX-WOL fracture mechanics specimens were included in the Turkey Point Unit No. 3 reactor materials surveillance program. For the
- temperature range of interest, it is impossible to obtain valid K fracture toughness data from ic reactor vessel surveillance programs. However, WittI73 described a method to at least obtain quantitative bounding values for the fracture toughness. The basis for the method of data analysis is the e4uivalent energy concept.
5-10
1 l
9148 2 ,
1 120 LEGEND
. j Q UNIRRADIATED j
e IRRAD ATEo (sso'r, s.se x iO'8N/cw23 .
_ 100 -
4.
E 80 - --
n n v v g ULTIMATE TENSILE STRENGTH w
60 -
Nk O n V G
. 0.2% YlELD STRENGTH 40 80
@- @ n t
C U '
60 -
REDUCTION IN AREA-
^
D l 40 -
r_
g TOTAL ELONGATION i
- 20 - v o n v
UNIFORM ELONGATION v
- m. m o 7 v V 0 ! ! !
l 0 100 200 300
, 400 500 600 .700 TEMPERATURE (F)
Figure 5 5. Tensile Properties for the Turkey Point Unit No. 3 Reactor Pressure Vessel Shell Forging 123P461VA 1 5 11
i 9146-8
\
i 140 LEGEND O Q UNiRRADIATED A s IRRADIAfED (550 t, 5.68 x 10'8 N/CN )
_ l20 -
dh.
a e
- 100 -
O E
a 80 -
be' O 6 ULilMAll ifN$lLL $lRENGlH A~C ;
0.2% YifLD SIRINGlH 60
. r 80 60 -
g REDUCil0N IN ARIA
- O
" 40 -
>=
b 10f AL FLONGAi10N 20 - *
-A%
iNiiORM tLONGAIION $
0 0 100 200 300 400 500 600 700 TEMPERATURE (OF)
Figure 5 6.
Tensile Properties for the Turkey Point Unit No. 3 Reactor Pressure Vessel Weld Metal 5 12
gw-w -
9148-10 O
m C
' e o
m l c
. .9.
C H ,
8
- '*** W m D a 8
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O (ISX) SS381S 5 13
s 1
, A method is outlinedI73 for obtaining KQd (subscript d being the test specimen thicknessi value from standard fracture toughness tests. These steps are as follows:
a Measure the area under the loaddeflection curve up to maximum load of a 4 specimen of thickness d (point A in figure 54).
t a
Select any point on the linear portion of the load deflection curve (point B in figure 54). Measure the area up to this point and divide this area into i
. the area up to maximum load (point A in figure 54). Call this ratio b.
e Using the load at point B (figure 54) as Po, calculate Ko fromg the K expression for a compact tension specimen.
Ko = /2 f (51) e Multiply Ko by the square root of b. This number is K Qd-The value thus determined is unique regardless of the point selected on the linear portion of the curve.
Shabbits[8] has demonstrated that added conservatism can be produced by testing size specimens dynamically. Dynamic testing is performed at loading rates on the order of 104 to 105 ksi W/sec. Such high loading rates result in an increased dynamic yield strength based on strain rates four orders of magnitude greater than static [9,10) , and there-fore produce valid plane strain fracture toughness conditions at increased temperatures with small thickness specimens.
Shabbits has shown that the temperature dependence of dynamic toughness (K id) is similar to the static fracture toughness previously determined. Dynamic load t deflec' ion traces can also be analyzed using the equivalent energy concept.
The Turkey Point Unit No. 3 WOL specimens were tested dynamically and analyzed using the equivalent energy method. The K id (Subscript d indicates dynamic test conditions) expression used in the analysis is shown in equation (5-2).
' Pb1/2 b
K 3}
gd = B(W)D2 (5 2)
[ where K id si the dynamic equi nt to static KQd and Pb the dynamic equivalent to P in equation (51) The factor t i is applicable to both analyses.
O 5 14
p 9l i4
-A i
h l 00 N
- la
(
i I
4 l
i i
l I
i m
8 a
E r
=
-~~~----------_-- _
g b,
e-m e h.u x
w t g-e u8
-x 3
o o .
h ..b 43
_b a=
Cd SD I
l m i
'l
[
i (SON 00d)d 5 15 O 9 e
. . . . q j <
.l I
l1
. 1 The testing of each of the four irradiated WOL specimens resulted in the determ lower bound K !
id data. The WOL test results are summarized in table 5 7.
The fracture behavior of the WOL specimens in the post irradiation conditi !
eralized by three types of load displacement curves. Type 1 is shown in figure 5 i
- the loaddisplacement curve shows gross yielding at the crack tip. The specimeni j
completely break in two. This was typical for specimen P 17. Load displacement beh
. type 2 is shown in figure 510 where there is an intermediate amount of plasticity specimen crack tip. The third type of load displacement behavior is signified very little plasticity occurred at the crack tip. Specimens P 16, W 5, and W-6 exhibit behavior characteristic of a combination of types 2 and 3.
When the K t id oughness values from table 5 7 are compared to adjusted K IR curves, the toughness of both the forging and the weld metal are well above the limits of the K curves. IR I
4 e
- i 6 16
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. c . 4 e 4 4 4 0 0 0 0 1 d
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SECTION 6 DOSIMETRY ANALYSIS
- 61. -
FAST NEUTRON FLUX MONITORS To effect a correlation between fast neutron (E > 1 Mev) exposure and the radiation-induced property changes observed in the test specimens, a number of fast neutron flux monitors are included as an integral part of the Reactor Vessel Surveillance Program. In !
particular, the surveillance capsules contain detectors employing the following reactions.
Fe (n,p) Mn' 5 Nisa (n,p) Cose Cuss (n,n) Coeo Np237 (n,f) Cs'37 Uras (n,f) Cs'37 in addition, thermal neutron flux monitors, in the form of bare and cadmium shielded Co Al wire, are included within the capsules to enable an assessment of the effects of isotopic burnup on the response of the fast neutron detectors.
The relative locations of the various monitors within Capsule T are shown in figure 4 3, while the radial and azimuthal position of the capsule with respect to the nuclear core and other reactor internals is illustrated in figure 41.
The nickel, copper, and cobalt aluminum monitors were in the form of wires placed in holes drilled in spacers at several axial levels within the capsule. The iron dosimetry, on the other hand, was accomplished by drilling samples from the Charpy test specimens. The cadmium-shielded neptunium and uranium fission detectors were accommodated within the dosimeter
. block near the center of the capxle.
The use of activation detectors, such as those listed above, does not yield a direct measure of the energy dependent neutron flux level at the point of interest. Rather, the activation I process is a measure of the integrated effect that the time and energy dependent neutron flux has on the target material. An accurate estimate of the average neutron flux level 61
- . r
(-
O-l incident on the various detectors may be derived from the activation measurements only if}
the parameters of the irradiation are well-known. In particular, the following variables are of interest. 1 i
e ' The operating history of the reactor .
a The energy response of the given detector a
The neutron energy spectrum at the detector location The procedure for the derivation of the fast neutron flux from the results of the'Fe5d(n,p) I Mn'd l reaction is described below. The measurement technique for the other dosimeters, l which are sensitive to different portions of the neutron energy spectrum, is similar. 1 z
S2. : ANALYTICAL METHODS The Mn'd product of the Fe'd(n,p)Mn'd reaction has a half life of 314 days and emits .
gamma rays of 0.84 Mov energy which are easily detected using a Nal scintillator in irra-disted steel samples, chemical separation of the Mn'd may be performed to ensure freedom from interfering activities. This separation is simple and very effective, yielding sources of very pure Mn'd activity. In some samples, all of the interferences may be corrected for by the gamma spectrometric methods without any chemical separation.
The analysis of. the sample requires that two procedures be completed. First, thedis- Mn'd integration rate per unit mass of sample and the iron conterit of the sample must be measured as described above. Second, the neutron energy spectrum at the detector location must be calculated.
For this analysis, the DOT Illl wot dimensions! multigroup discrete ordinates transport code is employed to calculate the spectral data at the location of interest. Briefly, the DOT cal-culations utilize a 21 group energy scheme, an S8 order of angular quadrature and a P 3
expansion of the scattering matrix to compute neutron radiation levels within the geometry of interest. The reactor geometry employed here includes a description of the radial regio internal to the primary concrete (core barrel, thermal shield, pressure vessel, and water annuli) as well as the surveillance capsule and an appropriate reactor core fuel loading pat torn and power distribution. Thus, distortions in the fission spectrum due to the attenuation of the reactor intemals are accounted for in the analytical approach.
S2
.. ., e
Having the measured activity, sample weight, and neutron e location energy sp of interest, the calculation of the thresh <,ld flux d as follows:
The induced MnM activity in the iron monitors may be expressed as D=ff;[E o(E)$(E) J-1 Fj(14A7Jle A(T rg)
(61)
\
l where: j i'
D =
Induced Mn" activity No =
(dps/gmp,)
Avogadro's number (atoms /gm atom)
A =
Atomic weight of iron (pm/gm-atom) fj =
Weight fraction of Fe" in the detector a(E) =
Energy dependent activation cross section for the Fe"(n,p)Mn" reaction (barns)
$(E)- =
Energy-dependent neutron flux at the detector at full reactor power 2
K =
(n/cm sec)
Docsy constant of Mn" (1/sec)
Fj =
Fraction of full reactor power during the Jth time interval, rj ;
rj =
Length of the Jth irradiation period (sec)
T =
f Elapsed time between initial reactor startup and sample counting (sec) j The parameters F , ry, and T depend on the operating history of the reac between capsule removal and sample counting. r and the delay The integral term in the above equation may be replaced on by the foll p og(Eles(E)'
o (E)d(E) = u I *'
J ETH )~ 1 I ETH (6 2) 43IE)
TH '
l 63
.1 o
e where:
F
= Effective spectrum average reaction cross section for neutrons above energy, E TH I Avwage autron flux above energy, E ETH TH og(E) = Multigroup FeH(n,p)Mn" reaction cross sections compatible with the DOT energy group structure 43(E) = Multigroup energy spectra at the detector location obtained from the DOT analysis Thus, n
D= fg o IE Fj (1 e'Af J)e-A(T rj)
TH (6 3)
J=1 or, solving for the threshold flux )
. D i
ETH (6 4) b A fU i b" F; (1 e'AT 3-1 J )e X{T rj)'
The total fluence above energy ETH si then given by n-F jj dETH " ETH (65)
J=1 where:
n FJ f) = Total effective full power seconds of reactor operation up to the time of capsule removal J=1 Due to the relatively low thermal neutron flux at the capsule location no burnup correction was made to any of the measured activities. The maximum error induced by this assumption is estimated to be < 1 percent for the Niss(n p)Cose reaction, 64
- . 1
)1 l
I
- 63. !
RESULTS OF ANALYSIS The fast neutron (E > 1.0 Mev) flux and fluence levels derived from in Capsule T are presented in table 61. The relative neutron spectrum calculated to the Capsule T location is given in table 6 2 and does not apply to other capsule locations. I Although the monitors employed in the surveillance program do not all respond with a I 1.0 Mev threshold, equivalent 1.0 Mev fluence levels were derived from the measured djl by means of the following expression.
\
7
$(EldE ele > 1.0 Mev) = d(E > E '
THI] [ h (6-6)
, ETH $(E)dE ,
. where
((E > 1.0 Mev) = neutron flux above 1.0 Mev i
((E > ETH) = neutron flux above energy E TH derived from the response of a given monitor ETH = threshold energy of the monitor in question (IE)dE
= energy dependent flux calculated by the P1MG code [12)
The validity of the above expression is predicated on an accurate determination of the neutron energy spectrum over the energy range of interest. The accuracy of the energy spectrum as calculated by the DOT code is an implied assumption underlying this entire analysis. By comparing the calculated fluence levels on the basis of a common thres more meaningful conclusions may be drawn regarding the validity of the calculated spectrum
- 64. DISCUSSION OF RESULTS Using the iron data presented in table 61, the average fast neutron (E > 1.0 Mev) flux Wel at the capsule location is determined to be 1.64 x 10 11 n/cm 2-sec for Capsule T.
This value corresponds to an average fast neutron fluence level of 5.68 18 x 10 n/cm2 for Capsule T.
Because capsules are located at a position closer to the core than the pressure ves samples accumulate fast neutron exposure at a more rapid rate than the adjacent vessel.
Where possible, samples are placed at azimuthal locations of relatively high fast neutron !
65
., p
' i TABLE 61 RESULTS OF FAST NEUTRON DOSIMETRY CAPSULE T Reaction and Measured
- Monitor Average" Flux '
Activity Fluence" Location (E > 1.0 Mov)
(dps/em)
(n/cm21ec) ' (E > 1.0 Mev)
' (n/cm2)
Fe"(n,P)Mn"
' R 63 2.90 x 106 1.41 x 1011 4.&O x 1018 H 23 3.23 x 106 1.57 x 1011 H 17 5.46 x 1018 3.34 x 106 1.63 x 1011 P 58 5.64 x 1018.
3.51 x 106 1.71 x 1011 W.24 5.93 x 1018 3.71 x 106 1.81 x 1011 -
W 18. 6.27 x 1018 ;
3.51 x 106 1.71 x 1011 5.93 x 1018 Nine(n,P)Co#
Center 2.56 x 107 1.50 x 1011 5.21 x 1018 Cus3(n,alCo60 Top 9.52 x 104 1.91 x 1011 Bottom 6.63 x 1018 9.19 x 104 1.85 x 1011 6.42 x 1018 !
Npas?(n,f)Cs '37 Center 2.16 x 106 2.44 x 1011 8.48 x 1018 U23e(n,f)Cs'37 Center 2.92 x 105 2.00 x 1012 6.92 x 1018
- Measured data are accurate to 110 percent
- Subject to 210 percent measurement error
% I 64
e TABLE 6 2 CALCULATED RELATIVE NEUTRON SPECTRUM AT CAPSULE T LOCATION Neutron Group Lower Energy (Mev)
Relative Neutron Flux 7.79 1.21 6.07
. 4.11 4.72 6.75 3.68 7.61 2 87 11.7 S.23 22.0 1.74 26.4 1.35 33.6 1.05 35.4 0.821 36.4 0.388 110.0 0.111
- 141.0 4.09 x 10 2 74,7 1.50 x 10 2 60.5 5.53 x 10-3 54.9 5.83 x 104 55.6 7.89 x 10 5 182.0 1.07 x 10 5 114.0 1.86 x 104 101.0 3.00 x 10'7 130.0 0.00 x 1580.0 67
. . t
e exposure so that specimens are irradiated at a rate faster than the maximum rate un vessel follows. wall. The ratio of these rates is called the sample " lead factor" and is de The energy and spatial distribution of neutron flux within the reactor geome from the DOT two dimensional Sn transport code. The radial and azimuthal d are obtained from an R,0 computation wherein the reactor core as well as the wa steel annuli surrounding the core are modeled explicitly. The axial variations are the obtained from an R,Z DOT calculation using ihe equivalent cylindrical core co neutron flux at any point in the geometry is then given by d(E,R,0,Z) = (E,R,0) F(Z)
(67) where d(E,R,0) is obtained directly from the R,0 calculation and F(Z) is a normaliz function obtained from the R,2 analysis. The core power distributions used in ,
and R,2 computations represent the expected average over the life of the station.
Having the calculated neutron flux distributions within the reactor geometry, of the capsule as well as the lead factor between the capsule and the vessel m mined as follows:
3 The neutron flux at the surveillance capsule is given by de = d(E,Rece,0 ,Z )
(6-8) and the flux at the location of peak exposure on the pressure vessel inner diameter i (v. max = 4(E,R y,0v max,Zv-max)
(69)
The lead factor then becomes LF = #c
- v-max (6 10)
Similar expressions may be developed for points within the pressure vessel with the surveillance program dosimetry serve to correlate the radiation-indu test specimens with that of the reactor vessel.
6 7
e l
The calculated lead factors relating exposure of Capsule T to the pres (
1/4 thickness location, and 3/4 thickness location are surface.
2.48,4.17 and 17 4 respectively.
{
Using these lead factors. vessel exposure levels normalized were to caps determined and are compared with DOT calculated levels listed in table 6; .
. TABLE 64 PRESSURE VESSEL FAST NEUTRON EXPOSURE BASE DOSIMETRY MEASUREMENTS AND DOT CALCULATIONS Capsule Dosimetry Measurements Flux DOT Calculations Location Fluence Flux j (n/cm2.sec) (n/cm2) Fluence (n/cm2.sec) I Capsule (n/cm2) j 1.64 x 10I1 5.68 x 1018 1.55 x 1011 5.37 x 1018 Vessel Surface 6.59 x 1010 2.29 x 1018 6.25 x 1010 2.17 x 1018 Vessel 1/4T 3.93 x 1010 1.36 x 1018 3.72 x 1010 1.29 x 1019 Vessel 3/4T 9.43 x 109 3.26 x 1017 8.88 x 109 3.08 x 1017 Finally, a comparison of endef life peak fast neutron exposure nt of the No. 3 reactor vessel as derived from both calculations results may be made as follows. psule and
- calculated = 6.30 x 1019 n/cm2 (6 11)
(Capsule T = 6.65 x 1019 n/cm2 (6 12)
These data are based on 32 full power years of operation .
ment at 2300 M between calculations and measurements is excellent.
69 r
i n
SECTION 7 REFERENCES 1.
S. E Yanichko, " Florida Power and Light Company Turkey Point Unit No. 3
. Reactor Vessel Radiation Surveillance Program," WCAP.7656, May 1971.
2.
!, ASTM Designation E185 66, " Surveillance Tests on Structural Materials in Nuclear I Reactors" in " ASTM Standards (1967), Part 31, Physical and Mechanical Testing'of Metals - Metallography, Nondestructive Testing, Fatigue, Effect of Temperature,"
pp. 638 642, Am. Soc. for Testing and Materials, Philadephia, Pa.,1967.
3.
W.' S. Hazelton, S. L. Anderson, and S. E. Yanichko, " Basis for Heatup and C Limit Curves," WCAP 7924, July 1972. ,
. 4.
L. E. Steele and R. H. Sterne, Jr., " Steels For Commercial Nuclear Power Reactor Pres Vessels," Nucl. Eng. Des.. 10. 259 307, (1969).
5.
T. R. Mager, Westinghouse Nuclear Engery Systems, Unpublished data.
6.
T. R. Mager, " Post Irradiation Testing of 2T Compact Tension Specimens," HSST Tech-nical Report No. 9, August 1970.
7.
F. J. Witt and Toughness K T. R. Mager, "A Procedure for Determining Bounding Values on Fracture ic At Any Temperature," ORNL.TM-3894 October 1972.
8.
W. D Shabbits, " Dynamic Fracture Toughness Properties of Heavy Section A533 Grade 8 Class 1 Steel Plate, HSST Technical Report No.13, WCAP 7623, December 1970.
9.
J. M. Kraft, G. R. Irwin, " Crack. Velocity Considerations" in Fracture T'oughness Testino and its Application, pp. 114 129. ASTM STP 381,1965. _
10.
M. J. Manjoine, " Influence of Strain and Temperature on Yield Stresses of Mild Steel" Journal of Applied Mechanics Trans. Amer. Soc. Mech. Engd6, A211 A218 (1944). i l
11.
R. G. Soltesz, et al, " Nuclear Rocket Shielding Methods, Modification, Updating, and input Data Preparation, Volume 5 - Two-Dimensional, Discrete Ordinates Transport Technique," WANL-PR-(LL) 034, August 1970.
12.
H. Bohl, Jr., et af, "PIMG: A One Dimensional Multigroup Pj Code for the IBM.704,"
WAPD.TM 135, July 1959.
13.
ASME Boiler and Pressure Vessel Code, " Nuclear Power Plant Components" Section lit, Summer 1972 Addenda, Non Mandatory Appendix G " Protection Against Non ductile Failure."
71
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APPENDIX A i RT increases as the material is exposed to fast neutron rakkation. Thus, to find the most limiting RTNDT at any time period in the reactor's life a ARTNDT due tn the radiation ex-posure associated with that time period must be added to the- !
. original unirradiated RTHDT. The extent of the shift in RTNDT is enhanced by certain chemical elements (such as copper) present in reactor vessel steels, Design curves which show the effect of fluence and copper content on 6RTNDT for reactor vessel steels exposed to 550*F are shown in figure A-1. ;
Given the copper content of the most limiting material, the radiation-induced ARTNDT can be estimated from figure A-1. Fast neutron fluence (E > 1 Mev) at the vessel inner surface, the 1/4T (wall thickness) vessel locations are given as a function of full power service life in figure A-2. The data for all other
- ferritic materials in the reactor coolant pressure boundary is presented in Table A-1.
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SURVEILLAtlCE CAPSULE REMOVAL SCHEDULE TO.DATE, CAPSULE I HAS BEEN REMOVED AND THE ENCAPSULATED SPECIMENS HAVE BEEN EXAMINED. BASED ON THE POST IRRADIATION TEST RESULTS ON-CAPSULE I, THE RECOMMENDED REMOVAL-SCHEDULE FOR THE REMAINING CAPSULES IN THE REACTOR VESSEL IS AS FOLLOWS: >
CAPSULE- FACTOR BY WHICH CAPSULE IDENTIFICATION LEADS VESSEL MAXIMM EXPOSURE 1/4 THICKNESS
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- a. CAPSULES TO BE REINSERTED IN CAPSULE.S AND I LOCATIONS i AT SOME LATER TIME IN PLANT LIFE IF ADDITIONAL CAPSULE TESTS ARE NEEDED.
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CLARIFICATION ITEMS ON TURKEY POINT UNIT NO, 3 REACTOR VESSEL RADIATION SURVEILLANCE REPORT NOTE 1. Figure A-3 is included to supplement the figure on Page 5-6 and also to substantiate using the temperature shift based on 50.ft. lbs. Charpy energy.- Figure A-3 which shows a plot.of the neutron radiation induced shift in Charpy V-notch energy and lateral expansion for the irradiated weld metal. A shift in reference temperature, RTNDT, based on Charpy V-notch energy was
- used-in the analyses since the pre-irradiated RTNDT was based on the Charpy V-notch energy. . It is-unreasonable i to determine the_ shift based on lateral expansion since the initial RTNDT was based on the impact energy measure-ments. In addition, it is appropriate to use the 50 ft. lb. shift since for the irradiated material the 50 ft. lb. temperature is equal to the 35 mils LE-temperature.
NOTE 2. On Page 5-11~the irradiated testing conditions were not-extended out to 600*F as was the unirradiated samples.. Reassurance that a mechanical property change will not occur is strengthened by data in the literature (1) including W tests indicates that the yield strength-of irradiated material increases approximately the same amount at any test temperature from room temperature to 550*F. This type of behavior is also indicated-by the limited tests performed on the Turkey Point Unit
- 3 material. Westinghouse, therefore, concludes that the increase in the yield strength of the forging at 550*F will be minimal since the yield strength in-crease at 80* and 200*F was minimal.
NOTE 3. On Page'5-17 the fracture toughness tests performed on the irradiated WOL specimens were not performed at upper-shelf temperatures and, therefore, the fracture (1) Hawthdrne, J. R., and Watson, H. E., " Strength and Notch Ductility of Selected Structural Alloys after High Fluence, 550'F '(288'C) Irradiation," NRL Report 7813, December 2, 1974.
A-6
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toughness value of 177.2 ksi/In. obtained on the-irradiated weld sample does not represent an upper shelf fracture toughness. The' data is intended to show that the level of toughness for the weld material-is higher than that predicted'by the KIR curve when this
- curve is' referenced ~to the RT En-closed is a plot (Figure A-4) NDT whichofshows the material.
how the facture toughness data on the weld material compares with the KIR curve. referenced to the RTNDT of this irradiated weld material.
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Lw WESTINGHOUSE CLASS 3 I )'
l FRACTURE MECHANICS EVALUATION OF
. FLORIDA POWER AND LIGHT TURKEY POINT UNIT NO. 3 REACTOR VESSEL i
1
{
J. H. Phillips T. A. Meyer R. W. Fleming C. S. Pillar November 1975 APPROVED:
. 'N. dhirigo nah Structural Materials Engineering Work Performed Under MIP 23572
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WESTINGHOUSE ELECTRIC CORPORATION Nuclear Energy Systems P. O. Box 355 Pittsburgh, Pennsylvania 15230
I l
TABLE 0F CONTENTS SECTION TITLE PAGE 1 INTRODUCTION 1-1 1-1. BACKGROUND 1-1 1-2. ANALYSIS PERFORMED 1-1 ,
1-3.
SUMMARY
OF RESULTS 1-2 h
STEAM. [NI YSSOF bob kChh BREAK kb bbk 2-1 2-1. LOSS OF COOLANT 2-1 2-2. STEAMLINE BREAK ACCIDENT 2-4 3 THERMAL ANALYSIS 3-1 4 FRACTUREMECHANICSANALYSIS 4-1 4-1. STRESS INTENSITY FACTOR EXPRESSION 4-2 4-2. IRRADIATION EFFECTS 4-3 4-3. FRACTURE ANALYSIS 4-7 5 MATERIAL INPUT 5-1 6 RESULTS AND CONCLUSIONS 6-1 1
6-1. LOSS OF COOLANT ACCIDENT 6-1 6-2. 6-2. RESULTS 6-1 6-3. LONGITUDINAL FLxd 6-1 6-4. CIRCUMFERENTIAL FLAW 6-1 6-5. CONCLUSION 6-5 6-6. LARGE STEAMLINE BREAK 6-5 l 6-7. RESULTS 6-5 6-8. LONGITUDINAL FLAW 6-5 6-9.
CIRCUMFERENTIAL FLAW 6-8 6-10. CONCLUSION 6-8 l
l -APPENDIX FPL COMMENTS ON FRACTURE MECHANICS EVALUATION A 0F TURKEY POINT UNIT NO. 3 A-1 III
l LIST OF ILLUSTRATIONS Figum Title Page 21 Safety injection System for the Turkey Point Unit No. 3 Power Plant 22 22 Flow Rate versus Pressure for the Turkey Point Unit No. 3 Safety injection System 23 23 Pressurizer Pressure and Cold Leg Temperature During a Large Steamline Break with Offsite Power Available 2-6 l
24 Pressurizer Pressure and Cold Leg Temperature During a Large Steamline Break Coincident with Loss of Offsite Power 27 j 25 Reactor Coolant Flow During a Large Summline Break Coincident with a Loss of Offsite Power 24 l j
St Temperature Distribution Through the Vessel Wall for a Loss of Coolant Accident 33 j S2 Temperature Distribution Through the Vessel Wall l for a Large Steamline Break with Offsite Power Available 34 S3 Temperature Distribution Through the Vessel Wall '
for a Large Steamline Break without Offsite Power 35 34 Thermal Hoop Stress Distribution Through the Vessel Wall for a Loss of Coolant Accident 3-6 SS Pressure and Thermal Axial Stress Distribution Through the Vessel Wall for a Large Steamline Break with Offsite Power Available 37 34 Pressure and Thermal Hoop Stress Distribution Through the Vessel Wall for a Large Steamline Break with Offsite Power Available 34 37 Pressure and Thermal Axial Stress Distribution Through the Vessel Wall for a Large Steamline Break without l
Offsite Power 39 34 Pressure and Thermal Hoop Stress Distribution Through the Vessel Wall for a Large Steamline Break without Offsite Powa 3 10 41 Longitudinal Crack in Cylinder (t/R = 0.1) 44 42 Circumferential Crack in Cylinder (t/R = 0.1) 45 !
i V j
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LIST OF ILLUSTRATIONS (Cont)
Figure Title Pego q
43 Effect of Fluence and Copper Content on RT l NDT f0f Reactor Vessel Steels Exposed to irradiation at 550*F 4-6 61 Stress Intensity Factors and Fracture Toughness for a Longitudinal Crack 500 Seconds after a Loss of Coolant -
Accident
$2 62 Longitudinal Crack initiation and Crack Arrest Depth for a i
Loss of Coolant Accident )
63
&3 Stress intensity Factors and Fracture Toughness for a J
)
Circumferential Crack 500 Seconds after a Loss of Coolant !
Accident 6-4 44 Circumferential Crack initiation and Crack Arrest Depth for a Loss of Coolant Accident 64 S5 Stress intensity Factors and Fracture Toughness for a Longitudinal Crack 600 Seconds after a Large Steamline j
Break without Offsite Power 6-9 66 )
Stress intensity Factors and Fracture Toughness for a Longitudinal Crack 600 Seconds after a Large Steamline Break with Offsite Power 6-10 j
$7 Longitudinal Crack initiation and Crack Arrest Depth for a Large Steamline Break without Offsite Power f 6-11 I 68 Longitudinal Crack initiation and Crack Arrest Depth for a large Steamline Break with Offsite Power S12 69 Stress intensity Factors and Fracture Toughness for a '
Circumferential Crack 600 Seconds after a Large Steamline .)
Break without Offsite Power 6 13
- 6 10 Stress intensity Factor and Fracture Toughness for a '
Circumferential Crack 600 Seconds after a Large Steamline Break with Offsite Power S14
$11 Circumferential Crack initiation and Crack Arrest Depth for a Large Steamline Break without Offsite Power 6-15 ~'
6 12 Circumferential Crack initiation and Crack Arrest Depth for a Large Steamline Break with Offsite Power 6-16 3
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SECTION 1 INTRODUCTION l
- 11. BACKGROUND The Turkey Point Unit No. 3 Nuclear Power Plant haw been commercially operated by the .
Florida Power and Light Company sina Ddamber, U72. The reactor vessel for this plant l was conetructed by Babcoco and Wilcox in accordance with the 1965 ASME Sciler and l
Pressure Vessel Code, Section ill, and the Westinghouse Equiptront Specification 676357. :
This report presents the results of a Westinghouse fracture mec:sanics study on the integrity l of the Turkey Point Unit No. 3 reactor-vessel belt line under two faulted conditions, i.e.,
a Loss of Coolant Accident and a Large Steamline Break. l The resul:t ci th . v!ysis were evaluated relative to current in;tistion and arrest criteria l
used by i%ewnme pertaining to the integrity of nuclear reactor vessels.
- 12. ANALYSIS PERFORMED in this integrity study of the reactor vessel for the postulated faulted conditions, the beltline region of the vessel is exposed to neutron irradiation.U.s'a result, the matenal properties
, , change, i.e. the ductile to brittle transition temperature increasesIll. The fracture mechanics I
analysis for the Loss'of Coolant Accident (LOCA) and the Large Steamline Break (LSB) therefore was applied to the core beltline region of the vessel where the potential crack instability based on en6of life conditions is more pronounced than for other regions. The fracture mechanics analysis considers, for conservatism, a postulated continuous longitudinal inside surface flaw using the most limiting material properties of the core region shell forgings, 1
as well as a postulated continuous circumferential inside flaw using the material properties of l
the circumferential core region shell weld. I The resulu of the f.w.ure mechanks analysis are presented in terms of the minirnum postulated l crack depth as a function of the time in the transient required to cause nonductile or fast fracture up to' 40 years of plant operation (end-of life).
l'
- 1. senese. L. L. Knighton. o. W., end Potapon u.. " Radiation Emtmettement of Presmer e veemis enc Peaceoutes for Umitems Wus affect en Poemer Res tors". Act. Appl., 4. 234244 (1ssal.
f l
11 l
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The complete analysis consists of the following steps:
a.
Description of the faulted condition transients for the Loss of Coolant Accident and Large Steamline Break,
- b. Thermal stress analysis based on the transient information.
- c. Fracture mechanics analysis.
- 13.
SUMMARY
OF RESULTS A fracture mechanics analysis of the Turkey Point Unit No. 3 reactor vessel for a postulated Loss of Coolant Accident and a Large Steamline Break occurring at 40 years of plant operation led to the conclusion that the integrity of the reactor vessel will not be impaired by a Loss of Coolant Accident or Large Steamline Break since:
a For the Loas of Coolant Accident, a postulated initial continuous circumferential crack between 0.18 inches and 4.3 inches in depth will arrest within 69 percent of the vessel wall thickness. For postulated continuous longitudinal cracks less than 4.0 inches in depth, unstable crack growth will not occur.
a For the Large Steamline Break without offsite power, postulated continuous cir-cumferential cracks between 0.26 inches and 1.8 inches in depth will arrest within 45 percent of the vessel wall, while for the Large Steamline Break with offsite power, postulated circumferential cracks between 0.26 inches and 1.9 inches in depth will arrest within 47 percent of the reactor vessel wall.
a For the Large Steamline Break, postulated continuous longitudinal cracks will e not be marked by unstable growth in the cases (1) offsite power not available and cracks less than 3.0 inches in depth, and (2) offsite power available and I
cracks less than 3.3 inches in depth.
I l
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12-
I #
l
. SECTION 2 i
DESCRIPTION OF THE FAULTED CONDITION TRANSIENT FOR THE LOSS OF COOLANT ACCIDENT AND LARGE STEAMLINE BREAK
' The following is a description of the safety injection system's full system performance; this is the performance or operating mode upon which the evaluation of the resistance-to-failure of the reactor vessel was based.
Figure 21 describes the pumping systems under consideration in this study. The four shared safety injection pumps are arranged to deliver the flow of refueling water down one header with branch lines to each of the three cnid legs of the unit undergoing the accident. Upon initiation, the system will deliver water from all four safety injection pumps from the two refueling water storage tanks. Figure 2 2 describes the flow rates versus reactor backpressure from the safety injection pumps to the cold legs of the reactor coolant loops.
Figure 2161so shows the two residual heat removal pumps which deliver a fbw of refueling water through a header and branch lines to each of the three cold legs.
All injection pumps take suction from the refueling water storage tanks, which are located outside the auxiliary building. Due to the outside location of the tanks, the minimum temp-ersture of the injection water can be taken as 39'F, the minimum recent historic outside air temperature for the area of the plant.
2 1. LOSS OF COOLANT ACCIDENT l During a double ended, hot leg, Main Coolant Loop Pipe Break (LOCA), the reactor vessel will blow down rapidly and the downcomer region will be refilled at least above the core centerline with 90*F water from the accumulators. Within 20 to 30 seconds from the start of the accident, the safety injection pumps will be started and will begin to deliver 155*F flow from the boron injection tank. The minimum volume of boric acid,900 gallons, will be delivered in 48 seconds (68 to 78 seconds from start of accident). The flow will then be from the refueling water storage tank, and the temperature of the delivered water will drop to 39'F. The maximum flow rate will be 780 gpm into two intact loops, assuming that the reacto'rpressure and containment pressure are near zero pressure. The injection flow will enter the reactor coolant piping through the accumulator connections.
The two residual heat removal pumps will also start on demand and will begin to deliver refueling water to loops within about 30 seconds of the start of the accident. The total flow is 2810 gpm from the low head injection system when the reactor backpressure is near zero.
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The injection phase can be considered to last for about 40 minutes from the start of the >
accident until the refueling water storage tank has emptied. At this time, long term recir-culation flow paths will be set up, the flow rates probably will be decreased by operator action. The temperature of the water will be between 100 and 250*F, depending on the decay heat at the time, since it will be recirculated from the containment sump. The analysis is based on a recirculation temperature of 100*F since this is the more conservative of the two temperatures. The flow rete will depend on the actions of the operator, but it is reasonable to assume the minimum recirculation flow would occur if the system is aligned for high head injection only, as per emergency instructions. This total flow rate would be -
about 680 gom. A minimum flow rate of 680 gpm is conservative since higher flow rates will host the vessel and reduce the stresses at a higher rate. .
It should be noted that during the injection phase the flow rate from the residual heat removal pumps is conservatively high because the refueling water storage tank was assumed to be completely full throughout the transient. The added suction boost has a significant effect on these low. head pumps.
2-2. STEAMLINE BREAK ACCIDENT i
During a Large Steamline Break accident, the reactor coolant temperature and pressure will rapidly decrease. The safety injection pumps will begin to deliver to the reactor when the reactor coolant pressure decreases below 1450 psig. The flow rates to the three cold legs
' will be dependent on the reactor coolant pressure as described by figure 2 2. Initially, the pumps will deliver boric acid from the boron injection tank (as described in LOCA, above) at a temperature of 155'F. After the stored volume of 900 gallons has been delivered, flow will begin from the refueling water storage tank, whereupon the temperature of the pumped fluid is assumed to decrease to 39'F.
The residual heat removal pumps are not expected to deliver refueling water because their developed head is not sufficient to overcome the reactor pressure expected during a steam
, break.
For the Large Steam Break, the following conditions are assumed:
a initial Conditions Noload conditions (roload temperature 547'F).
Reactor coolant system pressure 2250 psia.
Reactor coolant flow 268,500 gpm (corresponds to 199 ft3 /see loop flow).
e The largest inside containment break is assumed (4.372ft ). The check valve in the broken loop is aamamed to fail, resulting in steam flow from the unbroken 24 i
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steam generators. This steam flow is limited by the 1.4ft2 flow restrictor in the l steam lines. Isolation of the unbroken loops occurs 10 sec after the break.
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e The main feed system delivers the nominal plant flow to the broken loop for !
00 seconds following the break. The total capacity of the auxiliary feed system is supplied to the broken loop for 10 minutes.
e Thick metal heat capacity and reverse heat transfer from the intact steam generators f
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s No retum to criticality in the core is permitted and no decay heat is assumed, l I
e All four safety injection pumps operate, delivering a flow as indicated above. In addition, two charging pumps are assumed to operate, delivering a total of 150 gpm. j Safety injection and charging are assumed terminated by operator action ten minutes after initiation of the transient.
Two cases are considered in the study of the Large Steamline Break. In the first case, offsite 1 power is assumed available through the transient and reactor coolant flow is maintained, in
)
the second case, a loss of offsite power with a reactor coolant system flow coastdown is !
assumed at the time of the steamline break.
Figure 2 3 shows the reactor coolant system pressure variation and the cold leg temperature variation as a function of time when offsite power is available. Figure 2 4 shows tne reactor coolant system pressure variation and the cold leg temperature variation, while figure 2-5 shows the reactor coolant flow as a function of time coincident with a loss of offsite power.
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SECTION 3 -
THERM.!,L ANALYSIS The thermal analysis for faulted conditior> transients, such as the Loss of Coolant Accident and the Large Steamline Break, requires tcat the temperature and thermal stress profiles in the vessel be calculated as a function cf time.
The input to the thermal analysis consists of the following:
s Vesel barrel annulus volumetri: "owrote history e Flow tempeiature history
)
e Vessel-barrel annulus pressure history e Vessel and cainless steel clad geometry and material properties The temperature profiles are determined by a finite element calculation technique; the thermal stress profiles are determined by tna use of thermal stress equations for thick walled cylinders modified to include the effect of the vessei interior surface stainless steel clad. The change from two phase boiling heat transfer to convection heat transfer in the initial portion of the large LOCA transients is determine: by simultaneously calculating the heat flux from
. the vezel surface by both methods of hen transfer. The analysis assumes nucleate boiling to occur until the calculated heat flux fo' the vessel inner surface due to convection exceeds that due to nucleate boiling.
The heat transfer film coefficients for forced convection are calculated using the DITTUS-BOELTER I23 orced-convection f correlation:
0 o Pr .4 H = 0.023 f { Re .8 l
where f = safety coefficient = 1.5 ;
k = thermal conductivity of the Laid (8tu/hr.ft'F)
D = hydraulic diameter (ft) 2 Disaus. F. W. W SoWier. L. M. K.. Nest Trarufse Autornotnie Rad stors of ese Tutmelar Type. Calef Unar. Putscot.on en Eng. 2. peo.13. 44341 e te301.
51 .
l
m.
\
Re = #W = Reynolds' number ~\
. M p =
density of the fluid (Ib/ft3; V =
fluid velocity (ft/hr) p =
fluid viscosity (lb/hr/ft)
Cpu Pr = - =
g Prandtt's mmber Cp =
specific heat (Stu/lb 'F)
This correlation generally has an accuracy of 225 percent with regard to the data scatter in the correlation. An additional 25 percent margin is considered in this analysis as an added safety margin and to account for any geometry effects of the vessel barrel annulus-Therefore, an ample margin is added to the forced convection film coefficients calculated from the DITTUS BOELTER correlation before it is used in the temperature and stress i analysis.
The results of the analysis give temperature and thermal stress and combined thermal and pressure stress profiles through the vessel wall as a function of time. The temperature and combined stress as a function of time and fractional distance through the wall are needed to determine the critical crack size for initiation and arrest.
The heat transfer coefficient for forced convection for the 33*F water from the refueling water storage tank (the first 2400 seconds of the transient) is 208 Btu /hr ft2 -
- F. The heat transfer coefficient for the 100*F water during long term recirculation is 1602 Stu/hr ft .op ,
Figures 31 through 3 3 show the temperature as a function of distance through the wall at several time steps during the LOCA and LSB. In figures 3-4 through 3 8, the resulting thermal rtresses, as function of distance through the wall at the corresponding time steps are shown. ,
O e
32 -
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0 I ! I l l l l 1 0.0 1.0 2.0 3.0 11 . 0 5.0 6.0 7,o e,o DISTAliCE THROUGH THE VESSEL WALL (INCHES)
Figure 3-1. Temperature Distribution Through the Vessel Wall for a Loss of Coolant Accident !
l 33
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Figure 3 2. Temperature Distribution Through the Vessel Wall for ~
a Large Steamline Break with Offsite Power Available 3-4
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Figure 3 3. Temperature Distribution Through the Vessel Wall for a Large Steamline Break without Offsite Power SS l
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Figure 34. Thermal Hoop Stress Distribution Through the
- Vessel Wall for a Loss of Coolant Accident '
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Figure 3 5. Pressure and Thermal Axial Stress Distribution Through the j Vessel Wall for a Large Steamline Break with Offsite Power Available 3-7
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Figure 3-6. Pressure and Thermal Hoop Stress Distribution Through the
- Vessel Wall for a Large Steamline Break with Offsite Power Available
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34
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Figure 3 7. Pressure and Thermal Axial Stress Distribution Through the !
Vessel Wall for a Large Steamline Break without Offsite Power j 3-9 j l
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Figure 38. Pressure and Thermal Hoop Stress Distribution Through the Vessel Wall for a Large Steamline Break without Offsite Power {
w 3 10
)
i
~
I i
l SECTION 4 FRACTURE MECHANICS ANALYSIS l I
I l
Upon calculation of the stresses and temperatures in the reactor vessel beltline region resulting {
from the LOCA and LSB transients, fracture mechanics techniques were used to determine the limiting conditions throughout the entire course of the accidents. More specifically, !
I critical crack depths which must be exceeded before fracture initiation can occur were determined in the analysis, which was based upon the principles of the linear elastic fracture mechanics (LEFM) technology.
i The LEFM approach to the design against failure is basically a stress intensity factor considera-tion in which criteria are established for fracture instability in the presence of a crack [3) Con- .
sequently, a basic assumption employed in LEFM is that a crack or crack like defect exists in the structure. The essence of the approach is to relate the stress field developed in the vicinity of the crack tip to the applied nominal stress on the structure, the material properties and the size of defect necessary to cause failure. j 1
The elastic stress field at the crack tip in any cracked body can be described by a single para-
. meter designated as the stress intensity factor KI4I. The magnitude of K is a function of the geometry of the body containing the crack, the size and location of the crack, and the mag-nitude and distribution of the stress. The criterion for failure in the presence of a crack is that failure will occur whenever K exceeds some critical value.
For the opening mode of loading (stresses perpendicular to the major plane of the crack) the stress intensity factor is designated as Kg and the critical stress intensity factor is designated Kic. Commonly called the plane strain fracture toughness, Kge is an inherent material property which is a function of temperature and strain rate. Any combination of applied load, structural configuration, crack geometry and size which yields a stress intensity factor greater than Kie for the material will result in crack instability or fast fracture.
W hile K is associated with crack initiation and is determined by static fracture toughness ie testing, another .LEFM parameter, called the reference fracture toughness of Kgg, is associated with arrest of an unstable propagating crack.
- 3. Maeor T. R,. Suchewt. C., and Entietto. J. F., " Discuss.on et Practure Mechanses Concepts". WCAP 7a41. Maren 1972.
- 4. Grimth. A. A., "The Phenomens of Ruotw e are F ow = soids". Phil. Tram Roy. soc. Lorcon. 221. ser. A587 1& l198 (1920s.
41
, so
- )
- i i
Whenever the stress intensity factor Kg of a propagating crack becomes equal to KIR at the corresponding material temperature, crack arrest is assumed to occur. At the same temperature, K ie is invariably higher than K lR' The (static) fracture toughness versus temperature curve of K te curve essentially represents the
' lower bound K;c values of A.533 Grade B Class 1 and A 508 Class 2 reactor steels and weld.
metal. The reference fracture toughness versus temperature curve or K IR curve represents the lower bound dynamic and arrest fracture toughness values for both types of reactor steels.
The temperature scale is defined relative to the reference nil-ductility transition temperature or RTNDT. The RTNOT si a non-physical constant related to the brittle to ductile fracture trans-ition temperature; it is determined by both drop weight tests and Charpy V impact tests.
RT i NDT s defined in Article NB 2331 of the Summer 1972 Addenda to the ASME Section ill Boiler and Pressure Vessel Codel6I. : i The Kiccurve is analy'tically described [6] by:
Kje = 33.194 + 2.806 exp [0.02 (T RTNOT + 100)] (41)
The KIR curve is given[5] by:
KIR = 26.78 + 1.223 exp [ 0.0145 (T RTNOT + 160)] (4 2)
- 41. STRESS INTENSITY FACTOR EXPRESSION The fracture mechanics analysis requires the determination of the stress intensity factor solution for the continuous longitudinal and circumferential inside surface cracks which have been -
assumed for the cylinderical beltline region.
Furthermore, the steep stress gradients developed in the vessel wall during the postulated accidents. require that the actual stress profile ,e used in the stress intensity factor expressions.
The stress intensity factor for a continuous crack in a given structure subjected to an arbitrary nominal stress field a (x, o) can be written as follows:I7I Kg = k [A, Fj +f Aj F 2+ A2 F3+ A3 F)4 (43)
- 5. AsME Boder and Pressure Veeses code, section lit. Nucieer Pouer Plant cesnoonents. ASMe. New Yorm.199.
.~
- 6. AsVE Bo.ier one Preneure Vesset Code, sectmen xi. Acce%s A Eve 4uetion of Flow facicateas.
a
- 7. evene et. C.. Sam 8 erd. W. H.. strees sniens.tv Factor sos tions 4 contenuous Surface F'ews .a 8teactor r P essu e Vesseis. Preae-tes aan Natiocas serg. on Fracture Mecnaaics. Brown un v Provioence, R. l.. Act 1974 -"
i
~
r ...
t where:
a is the crack depth
. Fj, F 2i 3,F4 E are the magnification factors for the particular geometry considered .
relative to o = AO , o = Agx, a = A 2x 2, o = A x3 , respectively.
The applied stress profile a (x, o) is expressed as a third degree polynominal:
o (x, o) = Ao + A3 x + A2x +A"3 '
(*4)
C, 8. Buchalet and W. H. BamfordI7I- recently determined the magnification factors that are to be applied to the terms in equation 43 for continuous circumferential and longitudinal surface flaws in cylindrical vessels with a radius to wallthickness ration of 10. They developed a two-dimensional finite element model of a cylindrical geometry containing continuous longi-tudinal or circumferential surface cracks with depths from 2.5 to 80 percent of the wall' thickness to calculate the stress intensity factor relative to this geometry. Their results, pre-sented in terms of magnification factors versus fractional distance through the wall, are shown in figure 41 and figure &2 for the longitudinal and circumferential orientation.
- 42. IRRADI ATION EFFECTS
- Neutron ' irradiation has been shown to oroduce embrittlement which reduces the toughness properties of reactor vessels steels!Ol . The decrease in the toughness properties can be assessed by determining the shift to higher temperatures of the reference transition temperature RTNDT. Since the copper content of reactor vessel steel has now been identified as a major contributor to radiation embrittlement, copper' trend curves have been developed to relate ,
the magnitude of the shift in RTNDT to the amount of neutron fluence (see figure A3).
A surface neutron fluence of 6.3 x 10 19 n/cm2 (end of life) was used to determine the shift in RTNDT.
The reference fracture toughness curve, indexed to RTNDT. will shift along the temperature scale with a value equal to the increase in the RT NDT or f given levels of irradiation.
Based on the initial RTNDT values of the beltline weld and the intermediate shell forging and their respective copper contents, the end of life RTNDT values are determined from the copper trend curves. These final HTNDT values are subsequently used to calculate the Kic and KIR as a function of tne fractional depth through the wall.
- 7. suche'et. C Barnfoes. W. H. stress Inievoets Factoe soaviens for Cominuous survece Fiews in Reacto' Pressure venees.
Presented 8th National symo. on Fracture Mecnanics. 8 own un.v., P oveence R. l. Avpest 1974.
S. Potesovt. U., He,vthoene J. R., "The Efetet of Rescua. Esements on 55c'F leren.stion Response on seieciso P,eisure vesse. stee's one wcoments". NRL eso3. sesneneer 1se8.
43 4
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~
Ki avwa Ao F, , 2a A Fi2'a 2 Af23'1a 3 13 F.' ,
w 2 3w i Figure 41. Longitudinal Crack in Cylinder (t/R = 0.1)
)
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~
Ei " 4"
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A3 F,
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Figure 4-2. Circumferential Crack in Cylinder (t/R = 0.1) j l
)
4-5
. D
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ici l I I IIill I I I I l .1 1 I 1
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Figure 4 3. Effect of Fluence and Copper Content on RT NDT IC' Reactor Vessel Steels Exposed to irradiation at 550*F I
-l1
- l
u 4 3. FRACTURE ANALYSIS A fracture analysis method is used to determine the critical and crack arrest flaw sizes from stress intensity factor calculations for cracked bodies subjected to mechanical or thermal stresses. The analytical equations 4-1 and 4 2, for Kge and KIR respectively, are used to determine the crack size required for crack initiation and the maximum size of a running crack that will arrest.
. The minimum crack size for initiation is obtained at the first intersection of the generated stress intensity factor (Kg) curve with the fracture toughness (Kie) curve. Intersection of the Kg curve and the arrest toughness (KIR) curve yields the crack arrest size. It is noted here that the Kge and KIR curves are transformed to obtain the respective values as a function of the fractional distance through the wall rather than functions of temperature. This allows Kg, Kic, and K IR ot be plotted on the same scale from which the initiation and arrest flaw sizes can easily be determined.
Upon completion of the calculations for the applied transient, the results are presented per time step. The Kg , K ic, and KIR versus the fractional distance through the wall for the cor.
responding time steps are graphically presented.
The basic flow of logic employed in the analysis is simple. From the appropriate input, cal-culations from known analytical expressions for K g , K geand KIR are performed. Once these values have been calculated at each discrete time step, the intersections are determined and hence the critical crack sizes are determined.
The appropriate input is supplied as the temperature and the combined stresses as a function of the time and the fractional distance through the vessel wall. (figures 31 through 3 8).
47
v l
(
m SECTION 5 -
l
. MATERIAL INPUT? .
In determining the minimum critical depth for the postulated cont!nuous circumferential flaw, the following actual material properties of the circumferential weldment between the
- intermediate and lower shell forgings are appropriate:
Copper content = 0.31 weight percent initial RTNDT = 3*F For the postulated continuous longitudinal flaw, the most limitin) material properties of the shell have been applied in the presett analysis, that is, the propertie1.)f the intnmediate shell forging. The intermediate shell is made of A508 Class B mmerial.
Copper content = 0.07 weight percent
,L initial RTNDT = 40* F e
a 9
'J 51
.- , p
'l y
SECTION 6 RESULTS AND CONCLUSIONS 1
9
- 61. - LOSS OF COOLANT ACCIDENT .f The Los of Coolant Accident conditions are summarized in section 2. This analysis has been performed using methods of linear fracture mechanics to determine the postulated minimum ,
critical crack size at which unstable crack growth will occur in the Turkey Point' Unit No. 3 Reactor Vessel. )
- The stres intensity l factor Kg for a continuous longitudinal and continuous circumferential
. flaw, as a function of progressive crack depths, has been calculated from equation (4 3) at 'l time intervals of 100 seconds throughout the LOCA transient. The critical crack size (AcriticalI required for unstable growth is determined from the intersection of the Kg curve with the Kie curve, while the crack arrest value is determined from the intersection of the Kg curve with the KIR curve.
- 42. RESULTS 1 j
The results obtained for the Loss of Coolant Accident are presented for longitudinal and 1 circumferential flaws respectively (see table 6-1). <
I S3. Longitudinal Flaw - At 500 seconds after the beginning of the LOCA, a minimum ;
critical crack depth of 4.0 inches (corresponding with a/t of 0.52) is obtained from the ;
intersection of the Kg curve with the upper shelf of the Kge curve.
Figure 61 shows the Kg, Kge, and KIR curves at 500 seconds after the beginning of the )
accident.
In figure 6 2, the crack initiation and crack arrest depth are plotted as a function of time in j the transient. This figure shows that th'e critical initiation depth is approximately 4.0 inches at.500 seconds.
- 64. Circumferential Flaw - At 500 seconds after the beginning of the LOCA transient, l a minimum critical crack depth of A18 inches (corresponding to alt of 0.023) is obtained from the first intersection of the Kg curve and the Kic curve in figure 6 3. This crack, however, f arrests within 34 perant of the wall thickness, or a depth of 2.6 inches.
a 61
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' ' I I ' I I I 0.0 0.I 0.2 0.3 0.4 0.5 0.6 0. 7 0.8 0.9 1.0 FRACTIONALDISTANCETHROUGHWALL(a/t) ed Figure 61. Stress Intensity Factors and Fracture Toughness for a Longitudinal
- -i Crack S00 Seconds after a Loss of Coolant Accident
- 62
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Figure 6 2. Longitudinal Crack initiation and Crack Arrest Depth for a Loss of Coolant Accident 6-3 6 g
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FRACTIONALDISTANCETHROUGHWALL(a/t) l l
Figure 6-3. Stress Intensity Factors and Fracture Toughness for a Circumferential Crack 500 Seconds after a Loss of '
Coolant Accident G I
Figure 6 4 shows that a crack with a depth ranging from 0.18 inches (rninimum) to 4.3 inches (maximum) will initiate between 500 and 2400 seconds after the beginning of the accident.
The crack will arrest within a maximum depth of 69 percent of the wall thickness. Further-more, a crack with a depth greater than 4.3 inches will not initiate throughout this transient.
- 65. Conclusion it is concluded that the most severe crack orientation for the LOCA transient is the continuous circumferential surface flaw, postulated to exist in the weldment between the intermediate and lower shell. The severity results from the greater shift in the reference transition temperature (RTNDT) and thus, a shift in the Kje and KIR curves for the weldment at the end of life. The greater shift results from neutron irradiation of the weld material which has a high copper con-tent (0.31 wt-percent) when compared with the low copper content (0.07 wt percent) of the intermediate shell forging material.
For a postulated circumferential flaw between 0.18 inches and 4.3 inches, unstable crack growth may occur during a LOCA, but will arrest within 69 percent of the vessel wall.
For a postulated longitudinal flaw, the critical crack size is 4.0 inches. This crack depth greatly exceeds the minimum nondestructive detection limits and thus can be readily detected during a regular inspection period.
6 6. LARGE STEAMLINE BREAK The Large Steamline Break conditions are summarized in section 2.
Two cases were considered. In the fint case, offsite power is assumed available throughout the transient, in the second case, a loss of offsite power with a reactor coolant system flow coast-down is assumed at the time of the steamline break.
The methods of linear elastic fracture mechanics were used. Kg was determined for both a continuous longitudinal and a continuous circumferential flaw as was done for the Loss of ,
Coolant Accident.
- 67. RESULTS l i
The results of the Large Steam Break analysis are presented as for longitudinal and circum- )
ferential flaw, respectively (see table 6-1).
J
- 68. Longitudinal Flaw - At 600 seconds after the beginning of the LSB transient. a )
minimum critical crack depth ot 3.C inches without offsite power, and 3.3 inches with offsite )
power is obtained. l l
65 7
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Figure 6 4 Circumferential Crack initiation and Crack Arrest Depth '
for a Loss of Coolant Accident I
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r Figures 6 5 and 6 6 show the Kg and KIR curves at 600 seconds for both cases.
In Figure 6 7 and 6-8, the crack initiation and crack arrest depth are plotted as a function of time in the transient. These figures show the critical crack depths at 600 seconds,
&9. i Circumferential Flaw- At 600 seconds after the beginning of the Large Steam Break transient, a minimum critical crack depth of 0.26 inches (a/t .033) without offsite power avail-able and 0.26 inches (a/t 0.033) with offsite power available is obtained from the first inter-section of the Kg curve and K ie curve in figures 6 9 and 610. The crack initiated in the case "
without offsite power will arrest at 32 percent of the wall, while the crack initiated in the case -
with offsite power will arrest at 30 percent of the wall thickness.
.g.
Figure 611, the crack initiation and crack arrest depths plotted as a function of time for the case without offsite power, shows that a crack with a depth ranging from 0.26 inches (minimum) to 1.8 inches (maximum) will initiate between 600 and 1500 seconds after the beginning of the accident. This figure also shows that the crack will arrest within 45 percent of the vessel wall. F '
thermore, a crack with a depth greater than 1.8 inches will not initiate throughout this transient. -
Figure 612, the crack initiation and crack arrest depths, plotted as a function of time for the "
case with offsite power, shows that a crack with a depth ranging from 0.26 inches to 1.9 3 i inches will initiate between 600 and 1700 seconds after the beginning of the accident. This figure also shows that the crack will arrest within 47 percent of the vessel wall. A crack with a ,
depth greater than 1.9 inches will not initiate throughout the transient. *'
S 10. Conclusion .
it is concluded that the circumferential flaw is most severe for a Large Steamline Break (with ,
and without offsite power) for the same reasons that it is most severe in the case of the LOCA. j During a large Steamline Break without offsite power, a postulated circumferential flaw between 0.26 inches and 1.8 inches may become unstable. However, arrest of the flaw will occur within
- 45 percent of the vessel wall. For a Large Steamline Break with offsite power a circumferentia flaw between 0.26 inches and 1.9 inches may become unstable; however, this flaw will arrest )
within the 47 percent of the vessel wall. j For the postulated continuous longitudinal flaw for the Large Steamline Break, with and without offsite power, the critical crack depth exceeds the maximum nondestructive detection limits, and would be detected during the regular inspection period. For this reason, unstable 4 crack growth of a' longitudinal flaw will not occur.
i E
-d
.s e
- 7
u I
8857-88 I
300 Kg
- 200 K
_ IR C 1 a
5 i E
E W .100 0
0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0 FRACTIONALDISTANCETHROUGHWALL(a/t)
Figure 6 5. Stress intensity Factors and Fracture Toughness for a Longitudinal Crack 600 Seconds after a Large Steamline Break without Offsite Power I
~
8857-19 300 "I
bp200 m
E .
b E lR 5
N g 100 0
I I I I I I
- o. I l l 0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0 FRACTIONALDISTANCETHROUGHWALL(a/t)
,,I Figure S6. Stress Intensity Factors and Fracture Toughness for a Longitudinal -
Crack 600 Seconds after a Large Steamline Break with Offsite Power 1
1
'l Sto 'I J
i .
8857-20 n
1.0 0.9 -
0.8 -
0.7 -
0.6 ,
z !
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l E
v 0.3 - A egi7, cit = 3.0 INCHES 0.2 -
0.i o
l l l l l l l l l l l \ l l l 0 400 800 1200 1600 2000 2400 2800 i l
TIME (SECONDS) !
1 Figure 6 7. Longitudinal Crack Initiation and Crack Arrest Depth for a Large Steamline Break without Offsite Power 6 11
. . ?
-7 8857-21 f .0 -
0.9 0.4 -
0.7 -
I 20.6 a
~
- 0. 5 a
O -
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, i i
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0 l 200 400 600 600 1000 I200 140 1600 1800 2000 2200 2400 2600 2000 3000 3200 3 TIME (SECon05) 1 i
i l
.a
\
Figure 48. Longitudinal Crack initiation and Crack Arrest Depth for a Large Steamline Break with Offsite Power *J
.l i
s; i
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~
M57-22 300 200 r E Ki c7 K n
C 5
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r' 0 ' !
0.0 0.1 0.2 0.3 0.4 0.5 0.6 0. 7 0.8 0.9 1.0 l
FRACTIONALDISTANCETHROUGHWALL(a/t)
Figure 6 9. Stress Intensity Factors and Fracture Toughness for a Circumferential Crack 600 Seconds after a Large Steamline Break without Offsite Power 6 13
1 I
8857-23 l
1i e 8 300 i
I i'
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- 200 l
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'1 g 100
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0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0
. FRACTIONAL DISTANCE THROUGH WALL (a/t)
)
Figure 610. Stress Intensity Factors and Fracture Toughness for a Circumferential Crack 600 Seconds after a large ',,
Steamline Break with Offsite Power I,
6-14 s J
]
0857-24 l.0 0.9 -
0.8 -
O 0.7 7
~
0.6 -
5 5
o 0.5 -
0.4 -
MAXIMUN ARREST DEPTH 3.50 INCHES O 0.3 -
f MAXIM W fRACK INITIATIOM p DEPTH I$ l.8 INCHES
- R ')
0.2 -
~e www ?*V 1..." ::@f 0.1 . _Hff l2$ ' fh!;xCRITICAL,s 0.26 l' i i , i i i i l I I I I I 0.0 0 400 800 1200 1600 2000 2400 2600 TIME (SECONDS) l l
I i
1 i
l l
Figure 611. Circumferential Crack initiation and Crack Arrest Depth for a Large Steamline Break without Offsite Power 6 15
. ., 7
= !'
l l
8857-25 1
i l
1 1
10 0.9 - '
I 0.e -
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- 0. 2
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- 8 28' ;
0.1 I ' ' ' ' ! I I I 0.0 I I I ! ! l 0
210 400 600 800 1000 I200 1400 1600 1800 2000 2200 2400 2600 2400 3000 3200 TIME (SECOsD$) i i
1 l
l j
l l
1 .
Figure S12. ' Circumferential Crack initiation and Crack Arrest Depth for a .
Large Steamline Break with Offsite Power
- S16 O g i.m-..mm....
APPENDIX A CLARIFICATION ITEMS ON TURKEY POINT UNIT NO. 3 REACTOR VESSEL - FRACTURE MECHANICS EVALUATION NOTE 1. From Page 1-1 the current initiation and arrest criteria used by W are in compliance with the criteria presented in the Tatest addendum.to Section XI'of the
~
ASME' Boiler and Pressure Vessel Code.
NOTE 2. On Page 3-1 it is stated that the thermal stress equations included the effect of the_ vessel interior stainless steel clad while in. Figures 3-4 through 3-8 this is not indicated. The following explains what is !
meant here. I The vessel interior surface stainless steel clad. acts as an insulator for the vessel base material and is considered in the determination of the vessel base material temperature profiles. Also, the thickness of the stainless steel clad is considered in the thermal stress calculations but the clad properties are as-sumed.to be identical to the base material properties.
The fracture mechanics analysis considers only the j carbon steel wall. The plots in the accident analysis q consider the O point on the a/t' axis to be the inter-face between the clad and the carbon steel material. f
^
NOTE 3. On Page 4-3 the end-of-life RTNDT used in.this evalua-tion is based on weld metal test data on only one surveillance' capsule was available and represented early plant life. In the accident analysis (WCAP-8580) and the ASMI Appendix G analysis (WCAP-8581) both of which were performed for the end-of-life conditions, W elected not to extrapolate the one weld metal data Foint to end-of-life and instead used the standard W~
copper trend curves (Figure 4-3, WCAP-8580) to deteEmine the end-of-life RTNDT .- Data from future surveillance .
capsules will establish the actual end-of-life shift. l l
NOTE 4. On Page 3-1 the thermal stress analys'is for thick I walled cylinders is an explicit forward finite dif-ference calculation technique and the thermal stress equations are taken directly from " Theory of A-1
Elasticity," (2nd edition), Timoshenko and Goodier, 4 McGraw Hill, 1951. The input values to the thermal-analysis are the following:
3 C'arbon Steel S.S.
Material Clad Specific Heat-(Btu /4 *F) .128 .125 Density (#/f t 3) 490 497 Thermal Conductivity 4
. (Btu /hr-ft *F)- 25.2 9.8 NOTE 5. On Pages 6-3 through 6-16 clarification of the fracture toughness values used in;the upper shelf region is amplified by the following.
A large number.of fracture toughness tests.have been completed on: pressure vessel steels and associated welds and. weld heat-affected zones (1,2 , 3 , 4 , 5 , 6 ) '. Both static and dynamic toughness data have been obtained, and a lower bound for these data in the transition region is given by the reference toughness curve (KIR) of Section III (Appendix G) of the ASME Code. The same curve is also contained ~in Section XI (Appendix A) of the Code along with a curve (KIc) which is a lower bound-of static toughness only. The upper shelf tough-
' ness for these materials has not been characterized as yet in the ASME Code, because of the difficulties -
involved in testing in this area.
Fracture toughness testing results in two distinct types
, of fracture behavior. At low and transition temperatures fracture is by cleavage, and the onset of crack exten-sion is abrupt and unambiguous. . The maximum load corresponds to the fracture initiation point, and there-is no stable crack growth. Thus, methods for inter-
. preting data from toughness specimens such as the equivalent energy method (7) and the J integral l method (b) based on the maximum load point in the L
load-deflection record produce toughness values con-sistent with those obtained according to the require-l 1
ments of ASTM procedure E399, (9) Instrument precracked Charpy test results (4,5,6) obtained in the transition region.are also consistent with the other results.
At temperatures higher than the transition temperature fracture occurs by ductile tearing, and the onset of crack growth cannot be ascertained from the appearance L
l A-2 O
4
=
F of the load-deflection record. Thus, the methods of toughness determination based on the maximum load point (including precracked Charpy results) lose their validity. The test methods presented in ASTM procedure E399 are impractical for obtaining these " upper shelf" toughness values, because of the size requirements.
Results of the Heavy Section Steel Technology Program (1,2,3) have shown that valid toughness values cannot be obtained in this region with compact specimens as large as 12 inches thick. Testing of larger specimens to obtain a valid toughness measurement is both impractical and prohibitively expensive. Therefore, great care must be exercised in interpreting fracture toughness results in the upper shelf region. Two methods are available for reliable determination of upper shelf fracture toughness for reactor pressure vessel steels.
Data from large compact specimens (2T and 4T) can be used to measure toughness in the transition region based en the maximum load point, and will give accurate results as long as the fracture is cleavage contre 11ed.
Both static and dynamic toughness results are avail- ;
able (4,5,6) on unirradiated plate and forging material ;
and welds and heat affected enes which show that the onset of the upper shelf occurs at toughness values which range from 210 to 250 ksi/En. A careful study of the onset of upper shelf toughness using 1.9 and 4T compact specimens was recently made on irradiated plate and weld material (10) and results showed that the onset of the uo cess of 215 ksi/IE'.per for shelf occurred the weld at values material and 220in ex-ksi/In. for the base metal. l 4
Another reliable method of determining upper shelf 1 toughness is the J-integral method proposed recently by Landes and Begley (11). This method involves measure-ment and reporting of the actual amount of suberitical crack growth associated with the fracture process. A rather extensive test program (12) has been completed to obtain Jre fracture toughness results for pressure '
vessel steels. These results can be used to calculate Kre fracture toughness, through the relationship JIe = (I -E v2) K 7C 2 A-3
L I In choosing'J-integral values which correspond to some' minimal crack extension, we should obtain toughness values which,_when employed in fracture analyses for vessels at upper shelf temperature, predict the conditions for the onset of-slow crack growth rather'than unstable fracture. In this sense these values will be conservative. Choosing 120 mils as'a conservatively small. amount of crack extension, we obtain a minimum upper shelf toughness ofs200 kai/In. from reference (12).
In' conclusion, therefore, the fracture toughness of irradiated and unirradiated pressure vessel steels can be conservatively' determined from the following .,
procedure:- l 1
--In the lower temperature and transitionfregions use l the fracture toughness values given by the ASME-Code. i
--In the upper shelf region, let the fracture toughness' be constant, and-equal to 200 ksi/In.
This interpretation is consistent with the most recent fracture toughness testing results, and will provide i conservative toughness values throughout the range of temperatures at which the reactor vessel operates.
The upper shelf fracture toughness values of the reference fracture toughness curves presented in Figure 6 of enclosed ASME Publication referenced under question C.ll, were based on older and very little data. Therefore, these data are considered inap-propriate and are superseded by the recently obtained upper shelf fracture toughness values presented in reference (12).
i NOTE 6. On Figures 6-5 and 6-6 crack arrest is indicated by ;
the intersection of K r and the KIR curve as long as Kr for increasing crack depths becomes less than Krg.
The curves in question do not indicate crack arrest since after.the intersection Kg for progressive crack depths becomes greater than Kvg. The upper shelf of the KIe curve is 200 ksi/En. as is' the upper shelf of the KIR curve. Therefore, the crack initiation depth is obtained from the intersection of KI with the . constant 200 ksi/In. KIe, KIR upper shelf.
A-4
v APPENDIX A REFERENCES I
i
- 1) Shabbits, W. O., Pryle, W. H., Wessel, E. T.,
" Heavy Section Fracture Toughness Properties of A533B Class 1 Steel Plate and Submerged Arc Weldment," Heavy Section Steel Technology Report 6, December 1969. ,
- 2) Mager, T. R., " Fracture Toughness Characterization Study of A533B Class 1 Steel," HSST Technical Report 10, October 1970.
- 3) Shabbits, W. O., " Dynamic Fracture Toughness Properties of Heavy Section A535B Class 1 Steel Plate," HSST Technical Report 13, December 1970.
- 4) Marston, T. U., et al, " Fracture Toughness of Ferritic <
Materials in Light Water Nuclear Reactor Vessels,"
EPRI-232-2, December 1975.
- 5) Wu11aert, W. A., et al, " Fracture Toughness of Ferritic Materials in Light Water Nuclear Reactor Vessels,"
EPRI-232-1, Task B, February 1976.
- 6) van der Sluys, et al, " Fracture Toughness of Ferritic Materials in Light Water Nuclear Reactor Vessels,"
EPRI-232-3 (To be published).
- 7) Witt, F. J. and Mager, T. R.,
"A Procedure for Determining Bounding Values on Fracture Toughness K rc at any Tempera-ture," ORNC-TM-3894, October 1972.
- 8) Begley, J. A., and Landes, J. D., "The J Integral as a Fracture Criterion," ASTM STP 514, American Society for Testing and Materials, Philadelphia, 1972.
- 9) ASTM Standard E399-72, " Standard Method of Test for Plane Strain Fracture Toughness of Metallic Materials, in ASTM Standards, Part 31, pp 960-979, American Society for Testing and Materials, 1973.
- 10) Davidson, J.A., "The Irradiated Dynamic Fracture Toughness of ASTM A533, Grade B, Class 1 Steel Plate and Submerged Arc Weldment," HSST Program Technical Report No., 1976, in progress.
A-5 l
1
- \
.j i
- 11) Landes, J. D. and Begley, J. A., " Test Results from J-integral Studies, an Attempt'to Establish a J7e Testing Procedure," ASTM STP 560, 1974, pp 170-186.
- 12) Landes, J. D., Logsdon', W. and Begley, J. A., " Upper Shelf J 7e Behavior of A533B and A508 Class 2 Steels" (to be published). I f
)1
{
l~
,)
A-6 e
WESTINGHOUSE CLASS 3 t
W
. FRACTURE ANALYSIS FOR NORMAL, UPSET, AND TEST CONDITIONS FOR TURKEY POINT lll NUCLEAR STEAM SUPPLY SYSTEM BASED ON APPENDlX G, ASME CODE SECTION til i
J. Phillips S. Palusamy W.Ma ,
i October 1975 i
APPROVED: \, g K N. Chirigos, Managed Structural Materials Engineering Work Performed Under Shop Order MIP.23572 WESTINGHOUSE ELECTRIC CORPORATION Nuclear Energy Systems ,
P. O. Box 355 i Pittsburgh, Pennsylvania 15230
'l
. . r
)
4 TABLE OF CONTENTS SECTION TITLE PAGE 1 INTRODUCTION 1-1 l.-l.
.- 4 .
BACKGROUND l.-l
SUMMARY
OF CONCLUSIONS . .-4 2 ANALYSIS 2-1 2-1. STRESE JNTENSITY FACTOR DETERMINATION PER AdMt APPENDIX G 2-1 2-2. DESCRIPTION AND APPLICATION OF THE' REFERENCE STRESS INTENSITY CURVE 2-3 2-}. DERIVATION OF THE K4 p CURVE 2-6 4-4. ANALYTICAL DESCRIPTLON OF THE Kip CURVE -
- 5. IRRADIATION EFFECTS -
2b.
4- APPLICATION OF IR CURVE 2-7. CONSERVATISMOFTHEAb?kK APPENDIX G ANALYSIS 2-8 2-8. ELABORATION OF THE APPENDIX G ANALYSIS i FOR TURKEY POINT III NUCLEAR STEAM SUPPLY SYSTEM 2-14 2-9. DESIGN TRANSIENTS 2-14 2-10. GEOMETRY DIMENSIONS 2-16 2-11. MATERIAL DATA 2-16 s-2-12. CALCULATION OF MEMBRANE AND 2-16 2-13. BENDING STRESSES DETERMINATION OF THE MODE I STRESS INTENSITY FACTORS FOR PRIMARY AND SECONDARY MEMBRANE AND BENDING STRESSES 2-16 3 RESULTS AND CONCLUSIONS 3-1 APPENDIX COMMENTS ON FRACTURE ANALYSIS A 'ASME - SECTION III - APPENDIX G A-1 l
.\
LIST OF ILLUSTRATIONS Figure Title Page 11 The Four Analyzed Critical Locations in the Pressurized-Water Reactor Vessel 13 12 Analyzed Cross Section of Closure-Head to-Upper Flange Area 1-4 13 Analyzed Cross Section of Outlet Nozzle 15 14 Analyzed Bottom-Head to Shell-Course Region 16 21 Semielliptic Surface Flaw with Aspect Ratio'a/2b of 1/6. 22 l 22 Membrane and Bending Stress Multiplication Factors vs QThickness for Various Stress / Yield-Stress Ratios 24 1 23 Linearization of Stresses Through the Vessel Wall 25 2-4 KIR Reference Stress Intensity Factor Curve 27 25 Effect of Fluence and Copper Content on ARTNDT for Reactor Vessel Steels Exposed to irradiation at 550*F 2-9
{
26 Fracture Mechanics Analysis Upper Head Region 2 10 i 27 Fracture Mechanics Analysis Outlet Nozzle Region 2 11
{
28 Fracture Mechanics Analysis Core Beltline Region 2 12 l
, 29 Fracture Mechanics Analysis Bottom Head Region 2 13 1
)
l l
1 i
l I
1 I
I i
J l
1 LIST OF TABLES
)
, Table This Pop
~
.j 21 Transient Temperatures i 22 2 15 Material Data
, 23 2 18 Stress Intensity Factor Determination for Upper Head 2 19 -
24 Stress Intensity Factor Determination for Outlet Nozzle Region 2-5 2 20 Stress intensity Factor Determination for Core Beltline Region )
2-6 2 21 Stress intensity Factor Determination for Bottom Head Region 2 22 i e
l i
)
s l l
Vii ~
i]
4
. . i i
SECTION 1 INTRODUCTION
- 11. BACKGROUND Appendix G, "P otection Against Nonductile Failure," Section til to the ASME boiler and Pressure Vessel Codelll presents a method for obtaining the allowable loadings for protec-tion against nonductile failure for ferritic pressure-retaining materials in Class 1 components.
The ASME Appendix G method is based on the principles of linear elastic fracture mechanics.
For a postulated defect size, for each location being investigated, and for assumed specified loadings as calculated, the tension mode stress intensity factor Kg is calculated. The sum-motion of Kg due to primary and secondary stresses resulting from mechanical and thermal loading during normal, upset, and test conditions is then compared to a reference value KIR-The reference stres intensity factor K IR si the highest critical value of Kg for the material and the temperature involved.
. One of the recently published amendments to the 10CFR Part 50 regulations, marked as Appendix G " Fracture Toughness Requirements," specifies minimum fracture toughnem re-quirements for Nuclear Steam Supply Systems. Compliance with these requirements provides adequate margins of safety during normal and upset transients and test conditions to which
, a NSS System may be subjected over its service lifetime. The rules of this amendment set minimum material acceptance criteria and define minimum fracture toughness requirements during the service lifetime of the NSS System. In addition, the rules require that "the cal.
culated stress intensity factors shall be lower than the reference stress intensity factors by the margins specified in the ASME Code Appendix G".
This report presents an ASME Section til Appendix G Analysis for normal, upset, and test conditions for the Florida Power and Light Company, Turkey Point til Reactor Vessel, which is made of SA 302 Grade B and A508 Clas B material with a minimum specified yield strength at room temperature of 50 ksi or less.
Performing an Appendix G fracture mechanics analysis requires data on the magnitude and distribution of stress components through the reactor vessel wall for all cases of loadings.
I i
III AsME aooer one Pressure Veemi Cooe. Section m. Appendes G Winer 1s74 Addends )
11
Examination of Turkey Point 111 reactor vessel stress reports ovealed that complete infor- -
mation on stress distribution could be obtained for most but not all cases of loadings.
Therefore, the results from a more complete stress report, namely Ringhals il stress reports, were adapted as described in the following. Comparison of reactor vezel geometries showed that all dimensions except the outlet nozzle thicknem are identical within a deviation of less j than 10 percent. The Turkey Point 111 outlet nozzle wall is about 34 percent thicker than j that of Ringhals 11. The Ringhals 11 outlet nozzle stresses due to mechanical loads will pro-l wide a conservative estimate of the corresponding stresses for Turkey Point 111. However, j since greater wall thicknem leads to greater thermal stress, the thermal stremes must be revised. j Comparison of Turkey Point 111 and Ringhals 11 outlet nozzle heat-up/cooldown stresses dic-tated multiplication of the Ringhals ll thermal stres by a factw of 1.5. Further, comparison i
of specified mechanical and thermal loadings showed that in general Ringhals 11 is more severely loaded than Turkey Point 111.
The reactor vessel geometry and the four critical regions that are analyzed for protection against nonductile failure are shown in figure 11. Figures 1-2,13, and 1-4 are detail views of these critical regions.
For the core region and outlet nozzle, neutron irradisticir effects have been taken into account in the analysis. The fracture properties of the other two areas, remote from the core region, are essentially unaffected by irradiation.
- 12.
SUMMARY
OF CONCLUSIONS It is concluded that the Turkey Point lli reactor vesel complies with ASME Appendix G.
Plots of stress intensity factors at the four specified locations lead to maximum stress inten-sities, for all time steps of the transients, which meet the Appendix G fracture toughness re-guirements with a good margin of safety.
121 op;,,, gy,,33,,,,, , ,,,,,y,,,,, ,,,,,,,,,. ,,,,,,,, ,n, g y ,, e ,
12
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l NECHANISMS l
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CLOSURE HEAD l ,
!- i*@ CLOSURE HEAD REGION
'O,; THICKNESS CRITicb
[ TRANSITION r LOCATION
- h L. _ _ _ _ _ _ _ _ _ !
i
[
i [ CLOSURE FLANGE ICRITICAL
' LOCAT1011
.l l', N0ZZLE h SHELL-COURSE REGION - ~~
l jN f OUTLET N0ZZLE f/
U e
(
g 7_.
VESSEL WALL
- THICKNESS g,,7,e,t TRANSITION Location REACTOR s SELTLINE(
REGION % .
' s CORE . ,
l
+ +- THICKNESS = 7.75" RADIUS * '
'77.906" k VESSEL WALL T0 i LOWER HEAD j caTicALj r THICKNESS LOWER HEAD I LOCATION W,f, TRANSITION REGION '
IN-CORE L g,)- l INSTRUMENTATION L ,, , , , , , , , , , , , , ,
PENETRATIONS 1
i- l c
l
'- Figure 11. The Four Analyzed Critical Locations in the Pressurized Water Reactor Vessel 1
l 13
.m .
5 4
M07 13 i
s
'N 6.19" I
CRITICAL CROSS SECTION TOP CLOSURE JUNCTION v
4 Le ~
, .a Figure 12. Analyzed Cross Section of Closure- 'I t
,y Head to-Upper Flange Area 1
_a 14 5 J
, a #
8407-14 t
i CRITICAL CROSS SECTION 9.0" NEAN VESSEL RADIUS 78.75" Figure 13. Analyzed Cross Section of Outlet Nozzle 15
. t ,
. ~
8407-15 "M
/
\
, y,:##
t a : 7.75"
, , , . . r CRITICAL CROSS SECTION
,e 4.75" .
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Figure 1-4. Analyzed Bottom Head-to Shell<ourse Region i 1, "
14 ,;
. h 4
i
SECTION 2 ANALYSIS i
- 21. STRESS INTENSITY FACTOR DETERMINATION PER ASME APPENDIX G The stress intensity factor for the tension mode, indicated by K ,i depends on the geometry of the body in which the crack is postulated, on the size and the shape of the crack, and 1
, on the magnitude and shape of the stress distribution. '
The postulated defect is a sharp surface defect normal to the direction of maximum stress.
The defect shape is semielliptical with a depth one-quarter of the section thickness, and a major axis of length six times the depth, figure 21.
For cylindrical geometries (t < 0.1 R) with no discontinuities, the Appendix G expression is based on the following equation.
Ki [oM AM+88 AB) with: (21) a = flew depth O = the flew shape factor modified for the plastic zone size A M, A B: = the membrane and bending correction factors, respectively I eM* 88: = the membrane and bending stresses, respectively Q depends on the aspect ratio of the semielliptical flaw and on the ratio of acting stresses to j the nominal yield strength of the material, j
(
The parameters AM and AB are dependent on the aspect ratio of the semielliptical flaw and j on the flaw depth to thickness ratio, alt. '
For a given alt ratio of 0.25 and the given aspect ratio, the two correction factors can be combined with the term (ra/Q)l/2, yielding the ASME Appendix G expression:
I Ki = My oM + MB08 (2 2) l l
1 l
21
w J
8%7*9 >l i
i ii i
a !
~
i
+ a ->
i 2b 1 i
i y
, l l
Figure 21. Semiel11ptic Surface Flaw with 4
Aspect Ratio a/2b of 1/6. *j j
i Il l
<\
22 l'
, i J
For the primary stresses, a safety factor of 2 is applied for normal and upset conditions, and l
of 1.5 for test conditions, while for the secondary stresses a factor of 1 is used for all three J
conditions. The factors MM and MB are presented in graphical form in ASME Appendix G for various ratios of stres to yield strength (figure 2 2).
The membrane and bending stremes are determined from the hoop stres ' values at the inside and outside surfaces of the wall in each region of interest, implying the asumption of a linear .l stress distribution through the wall thickness as indicated in figure 2 3.
l However, stress intensity factors for the design transients in the nozzle region are not calcul- {
'sted in this manner, in the geometrical discontinuity of the nozzle region, the pressure can induce stresses as great as two or three times the membrane stress in the shell region. ASME Appendix G, recognizing that the nozzle region can not be expected to meet its requirements l for a onequarter-thickness defect, states that " smaller defect sizes may be used on an individual I case basis if a smaller size of maximum postulated defect can be assured".
WRC Bulletin 175, "PVRC Recommendations on Toughness Requirements for Ferritic l
Materials"[3] provides procedures for considering postulated defect sizes smaller than one- '
quarter of the wall thickness. Following these procedures, the reference flaw for the nozzle region was postulated to be a semielliptic surface flaw with a maximum depth of one fifth of -
the cross section thickness and an aspect ratio of 1/6.
The justification for a 1/ST defect for the nozzle is based on the present highly reliable non-destructive inspection techniques that ensure capsbility of detecting such a flaw. Because of the greater cross section thickness at the nozzle-shell junction, this flaw size is negligibly smaller than a 1/4T defect in the other areas of interest.
The' stress intensity factors for the design transients in the nozzle region were calculated by equation (2 2) for the modified membrane and bending correction factors which allow for a smaller defect sizel3) ,
2 2. DESCRIPTION AND APPLICATION OF THE REFERENCE STRESS INTENSITY CURVE Sections 2 3 through 2-6 develop the means for establishing a reference stress intensity curve which will enable evaluating the reactor vessel material for conformance to Appendix G stress requirements with respect to temperatures and irradiation levels.
WRC sulistin No.1751Auws 1972, PVRe Recommendations on Toughnen Moouirement for Ferritic Meterials-23
1 l
8407-10 I
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utunnut g,, = u, , ,
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YTHICKNESS(INCHES) 1 f
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Figure 2 2. Membrane and Bending Stred Multiplication Factors 1 j
vs Vthickness for Various Applied Stress / Yield Stress Ratios 1 4
a }
4 2-4
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. 4W7-25 e ,
lilSIDE OUTSIDE SURFACE SURFACE l
LillEAR APPROXIMATION JL
\
\ ACTUAL STRESS PROFILE l
J L N
$ \
E m JL \ jk u .
i JL \
N -,,
JL \
Jk N V
- e JL JL v
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l Figure 2 3. Linearization of Stresses Through The Vessel Wall 25
, f
T 3 O' 6
- 23. Derivation of the K IR Curve '
The principles of linear elastic fracture mechanics (LEFM) serve as a basis for the analysis methods of Appendix G of Section ill to the ASME Code. The central parameter of LEFM is the crack opening mode stress intensity factor K g. This single parameter defines the elastic stres field in the vicinity of a crack tip. K g is dependent on the geometry of the body con-taining the crack, the crack size and shape, and the magnitude and distribution of the stres.
A flaw will grow unstably whenever the Kg exceeds a critical value, Kic, the fracture toughness.
The fracture toughness is a material property, dependent on strain rate and temperature. K ie is also dependent on the metallurgical condition, i.e., it changes with the microstructure, neutron irradiation,' etc.
For stress intensity factor rates below 2.5 ksi vGi/sec, the fracture toughness is indicated by Kge, whereas for higher strain rates (the dynamic range), the critical stress intensity factor is ,
indicated by Kid-A third LEFM parameter, the arrest fracture toughness Kg ,, is the value at which a fast running crack (unstable propagation)_ will evantually stop. K ie values are invariably higher than Kid or Kg, values.
The reference stress intensity curve or KIR curve essentially represents the lower bound K ,
ic K id, and K g, values of A533, Grade B, Class 1 and A508 steels as a function of the -
temperature. The temperature scale is defined relative to the reference transition temperature or RTNDT. The RTNOT, a nonphysical constant which is related to the brittle-to ductile fracture transition temperature, is determined by both drop weight tests and Charpy V im-pact tests. The reference transition temperature is defined inIll, Article NB 2331 of the Winter 1974 Addenda. ASME Appendix G specifies that this KIR curve may be used for "
ferritic steels with a specified minimum yield strength at room temperature of 50 ksi or less. .
- 24. Analytical Description of the K IR Curve Figure 2-4 shows the KIR vs RTNDT adjusted temperature curve. To facilitate analytical cal-culations, the equation representing this curve can be expressed as:
KIR = 26.777 + 1.223 exp 0.014493 (T-(RTNOT-160)] (2 3) ,)
I l ,
I AsME Bode ana Preiwe vessee Coas. section lH. Accendia G. J
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.. j 24 1 J
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(KiR-as.m)=i.22se 0'*8IT-( "TuoT-'"i i WNERE 130 r ,=
i REFERENCE STRESS INTEN$lTY FACTOR T = TEMPERATURE AT WHICH r ig IS PERMITTED, *F 110 IInoT = REFERENCE NIL DUCTILITY TEMPERATURE 7
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- 70 -
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-240 -160 80 -W 0 W 80 160 240 TEMPERATURE RELATIVE TO ATuoi, (T-ATuot),FAHRENHEITDEGREES Figure 24. KIR Reference Stress Intensity Factor Curve t
27
- s. . l
,a Saeed on Westinghouse data, a KIR upper shelf value of 200 ksi v7 has been adopted for ~'
unirradiated material. For highly irradiated material, the upper shelf has been fixed at.
170 ksi vTri" 2 5.- Irradiation Effects -I Neutron irradiation adversely affects the ductility of reactor vessel steels. The neutron embrit-tlement of reactor vessel steels has been shown to be a function of the copper content for. -
]
given fluences.
j i
One of.the consequences of decrease in ductility is a drop in the fracture toughness. Quanti-tatively, this drop can be assessed by determining the shift to higher temperatures of the ref- I erence transition temperature. Copper trend curves have been developed to relate the mag- -
nitude of the shift in RTNDT ot the amount of neutron fluence (see figure 2 5).
l The reference fracture toughness curve, indexed to RTNDT, therefore will shift along the '
horizontal scale with a value equal to the increase in the RTNDT or f given levels of irradiation.
.Gince the RTNDT value at the end of life differs among the four analyzed regions of the reactor vessel, similarly differing reference. fracture toughness curves are required. The adjusted ,
Ki g curves are shown in figures 2 6 through 2 9.
2 6. Application of the KIR Curve -
The Kg values calculated for the closure head, outlet nozzle,. beltline region, and bottom head at each design transient, are plotted on the respective KIR curves (indexed to the RTNOT Of each region) at the temperature of interest (see figures 2 6 through 2 9). Protection against ,
d nonductile failure is then conservatively assured if the Kg values are below the corresponding K i g curve.
2 7. CONSERVATISM OF THE ASME APPENDIX G ANALYSIS .
A fracture mechanics evaluation in compliance with the rules of Appendix G to the ASME Code inherently includes a number of conservatisms.
1{
The most evident ones are obviously the safety factor of two applied to the primary bending '
and membrane stresses and the large postulated defect size. -
Further conservatism is introduced into the analysis by the expression used to calculate the 14 stress intensity factor (equation 21). This expression was derived for semielliptical flaws in )
plates, whereas the cylindrical geometry of a vessel or a nozzle is more constrained. This constraint results in lower actual Kg values than result from calculation by the methods of i appendix G. *J l T'
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1 The separate calculation of the primary and secondary stresses and their respective Kg v '
involves another conservatism. It is assumed that negative stresses do not contribute at all to the total stress intensity factor, thus implying that the total Kg will be equal to that result-ing from the positive stress rather than the total stresses.
Additional conservatisms are also introduced into the analysis through the use of the reference 3i fracture toughness curve. This curve is a lower bound curve determined from crack arrest
}
toughness values, dynamic fracture toughness data, and static toughness data, g including K values ensures that a flaw as large as the reference flaw will not propagate'and that a pro- '
pagating crack will arrest before it exceeds the size of the reference flaw. J A last point will be made with respect to the adopted value of 200 ksi A and 170 ksi </Iii for the fracture toughness upper shelf of unirradiated and irradiated material, respectively. _ )<
d i
Westinghouse data Indicate that the unirradiated upper shelf dynamic fracture. toughness for '
reactor vessel materials is substantially above 200 ksi </In7 Values between 250 and 450 l ksi A have been determined in actual practice.
- Therefore, it is concluded that the upper shelf value of 200 ksi </iri for the K ]
IR curve is a J highly conservative restriction as to the allowable Kg values determined for closure head and ,
bottom head since these regions are essentially not exposed to neutron irradiation, s
For highly irradiated material, the upper shelf value drops to an estimated lower-bound value of about 170 ksi </Tn7
- 28. f ELABORATION OF THE APPENDlX G ANALYSIS FOR TURKEY POINT lil '
i NUCLEAR STEAM SUPPLY SYSTEM a
The Turkey Point 111 reactor vessel was built by Babcock and Wilcox Company in confor- ,
I mity with the ASME Boiler and Vessel Code, Section lil, Nuclear Vessels,1965 edition and the Westinghouse Equipment Specification 676357I41 ', ,
2 9. Design Transients 1
.2 ,
The normal, upset, and test transients used in the analysis are listed in table 21, together with the governing transient temperatures for each of the analyzed areas.
.s 1 9
i I
Weeunghouse NES Eauoment specification 6763s7. '
. t 4
2 14 a ,l 1
TABLE 21 TRANSIENT TEMPERATURES j i
Temperatures, 'F i Cold Leg (For Top Head, Hot Leg (For 4 Beldine, and Bottom Head) Oudet Nonle) !
. Transient -
Heatup Cooldown 70 70 Plant Loading and Unloading 547 547 Small Step Load increase Decrease 533 599 ;
Large Step Load Decrease 535 528 Loss of Load 549 544 Loss of Flow 535 492 Reactor Trip 537 529
{
Turbine Roll 475 475 i
Steady State '
Fluctuations 541 607 i
+
l NOTE: Transient temperatures are not shown for the hydrostatic test conditions since these transient .-
temperatures are determined to insure compliance with ASME appendix G requirements.
l l
l 2 15 j
t n
2 10. Geometry Dimensions 1 Figures 11 thru 14 present the geometry dimensions of the investigated areas as well as the indications of the critical cross sections in the closure head to upper flange region, the out. I let nozzle-shell junction, the core beltline region, and the bottom-closure head to-shell course -
region.
2 11. Material Data I I
i Table 2 2 lists, for the regions of interest, data on Cu content, the initial RT !
NOT as deter- j mined by drop weight testing and Charpy V impact tests, and the predicted shift in RT '
NOT at the end of life for the corresponding total integrated fast neutron flux. The predicted shift in RT ,
NOT has been obtained from the copper trend curve as shown in figure 2 5. Since the core beltline region sustains substantial neutron irradiation, the predicted shifts for base metal and weld metal from this region are listed in table 2 2. The respective K IR curve has correspondingly been shifted for the end of-life RTNDT of the weld metal, since the shift is greater than the shift for the base metal.
The nozzle to-shell region is also subject to neutron irradiation, but due to its attenuation bya j water and reactor materials, the end-oflife fluence is much less here than in the beltline ~
region. The end of-life fluence for the nozzle region and predicted shift for this material are also listed in table 2 2.
The K ,
IR curves for the remaining regions have been determined according to their respective initial reference transition temperature. The KIR curves for the four regions are shown in '
figures 2-6 thru 2-9.
2 12. Calculation of Membrane and Bending Stresses '
For the four regions of interest, the primary membrane and bending stresses have been cal- '
culated from the inside and outside stresses presented in the stress reports for each design '
transient. Since for most transients the stresses are given for several time steps during each ,
transient, only those stress combinations resulting in the highest Kg value have been listed in the summary tables,2 3 thru 2 6. ,
2 13. 1 Determination of the Mode i Stress intensity Factors for Primary and Secondary Membrane and Bending Stressas ,
Equation (2 2): K; = MM8M+MB '8 has been used for the calculation of the primary and secondary stress intensity factors. The summation of both Kg values, witn a safety -
factor of two applied to the primary stress intensity factor, yields the ASME Appendix G <
stress intensity factors. These are listed in tables 2 3 thru 2-6 for the most severe time
.m steps during any of the design transients for the four regions of interest.
)
2 16 a
r
l The multiplication factors My and M 8 or f the primary and secondary stresses, derived for any region by means of figure 2 2Ill, are based on the combination of the highest primary mem-brane plus bending stresses and the highest secondary membrane plus bending stresses, respec-tively, resulting from any of the derign transients.
l l
i l
l I
l Ill AsME soae, ana Pr w. v ' coo.. s.ct.on m Appengen G Winer 1974 Adoence.
2 17 l
r . 'T TABLE 2 2 MATERIAL DATA initial Predicted End of Life Cross Section Cu RTNDT End of Life Fluence at Thickness (Wt %) (*F) RTNOT (*F) '1/4TI4) (Inches)
Closure Head To Flange N.A.Ill 60(2) 60 Negligible 6.19 Outlet Nozzle (3) 0.30(2) 50(2) 154 1.61 x 1018 g,o Beltline Region 0.07 40 149 3.76 x 1019 7.75 8eltline Weld 0.31 3 283 3.76 x 1019 7.75 Bottom Head N.A. 60 60 Negligible 4.75 j '
s (1) Not Available 1
(2) Estimated ,
(3) End of Life Fluence (col. 4) is for 1/5T -
(4) End of Life Fluence (n/cm )2 based on load factor of 0.8 for 40 years plant operation I
.)
1 4
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TABLE 2 3 '
STRESS INTENSITY FACTOR DETERMINATION FOR UPPER HEAD l
STRESS (KSI)
Primary Stress Intensity Secondary Factor, K Membrane Bending Membrane Bending (ksiVTn) g Transient Hestup-Cooldown (End of He,atup) 19.3 - 3.7 -12.6 -9.0 92.6 Plant Loading -
and Unloading 19.3 3.7 0.1 0.4 93.5 Step Load increase and Decrease 21.2 4.1 2.7 3.7 114.1 Large Step Load Decrease 21.8 3.6 -0.9 1.2 104.6 Loss of Load 25.5 -4.9 2.4 3.0 122.4 Loss of Flow 19.3 3.7 0.0 0.0 92.6 Reactor Trip 19.3 3.7 0.0 0.0 92.0 Steady State Fluctuations 21.8 3.6 0.4 0.6 106.5
. Hot Hydro Test
- 21.4 4.1 -5.2 7.2 88.5 Turbine Roll Test 19.3 3.7 0.0 0.0 92.6 Cold Hydro Test' 26.0 4.7 0.0 0.0 93.6
' Safety factor of 1.5 apphed to these transients as reowered tv Appenden G.
2 19
'{ l 1
\
TABLE 2 4 i' STRESS INTENSITY FACTOR DETERMINATION FOR OUTLET NOZZLE REGION l
URESS (KSI) 1 Stress intensity l Primary Secondary Factor, Kg '
Transient Membrane Bending Membrane Bending (ksiW Heatup Cooldown (End of Heatup) 28.2 4.7 2.3 13.8 157.3 Plant Loading and Unloading 28.2 4.7 3.8 -6.0 157.3 Step Load increase and Decrease 29.1 4.9 2.0 2.4 162.5 Large Step ,
Load Decrease 29.5 4.9 0.2 3.3 171.0 Loss of Load 31.4 5.2 -0.9 1.2 177.3 Loss of Flow 28.2 4.7 1.1 5.1 109.2 Reactor Trip 28.2 4.7 0.2 3.2 163.1 Steady State Fluctuations 29.5 4.9 -0.9 1.1 164.5 Hot Hydro Test' 31.3 5.2 2.0 12.3 130.9 Turbine Roll Test 28.2 4.7 2.0 9.5 179.4 Cold Hydro Test
- 39.4 6.6 0.0 0.0 164.9 l
se.w rum, et ts owi.e e e=e u.a.ma a reoui=e av Ase.aois c.
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r -
TABLE 2 5 STRESS INTENSITY FACTOR DETERMINATION FOR CORE BELTLINE REGION STRESS (KSI)
Stress Intensity Prirnary M ry Factor Kg Membrane Bending Membrane Bending (ksi m)
Heatup Cooldown '
(End of Heatup) 22.9 1.0 0.7 1.6 130.5 Plant Loading and Unloading 22.9 1.0 0.0 0.0 125.8 Step Load increase and Decrease 23.7 1.0 2.0 1.3 137.7 Large Step Load Decrease 24.0 1.0 -1.5 -1.5 131.7 Loss of Load 25.5 1.1 3.3 3.8 140.0 Loss of Flow 22.9 1.0 0.0 0.0 125.8 .
Reactor Trip 22.9 1.0 0.0 0.0 125.8 Steady State Fluctuations 24.0 1.0 0.7 1.0 135.3 Hot Hydro Test' 25.5 1.1 0.0 0.0 105.0 Turbine Roll Test 22.9 1.0 0.0 0.0 125.8 j Cold Hydro Test' 31.9 1.4 0.0 0.0 131.4
- s '.tv r cio, of 1.s soon.e to in. ir n.i.nts ei r.auir.o tv Apo.ndi= c.
2 21
=
TABLE 2 6 -
STRESS INTENSITY FACTOR DETERMINATION FOR BOTTOM HEAD REGION
- b STRESS (KSI)
Mmary Stress intensity Secondary Factor, K Membrane Bending Membrane Bending (ksi vin) g Transient -
Heatup.Cooldown (End of Heatup) 19.1 1.4 0.0 0.0 84.1 Plant Loading and Untoading 19.1 1.4 0.0 0.0 84,1 e Step Load increase .
and Decrease 19.7 1.4 1.2 1.4 86.6 Large Step .'
Load Decrease 20.0 1.4 1.1 1.3 87.9 Loss of Load 21.2 1.5 2.9 3.2 93.2 Loss of Flow 19.1 1.4 0.0 0.0 84.1 Reactor Trip 19.1 1.4 0.0 0.0 84.1 -
Steady State
- Fluctuations 20.0 1.4 0.5 -0.6 87.9 .
Hot Hydro Test
- 21.2 1.5 -4.5 9.0 69.9 <
Turbine Roll Test 16.4 1.2 3.0 8.2 90.4
)
Cold Hydro Test
- 26.5 1.9 0.0 0.0 87.4 I
)
M s v.w en- of u eaw to m ven..no on.o e,r 4 o.nois o.
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SECTION 3 RESULTS AND CONCLUSIONS I
. This analysis has considered the design transients that the Turkey Point til Nuclear Steam Supply System could experience during its design lifetime; the degree of compliance with the requirements of Section lil, Appendix G of the ASME Code has been determined. l The results of the analysis of the closure head, outlet nozzle region, core beltline, and bottom head are plotted in figures 2-6 thru 2 9. Each transient is represented by a point corresponding to the maximum stress intensity resulting during that transient. (The calculated stress intensity factors for the heatup and cooldown transients and the hydrostatic test conditions are not shown since these transients are structured to ensure compliance with ASME, Appendix G.)
For the four regions considered in this analysis, the results show that the maximum stress in-tensity for any time step during the design transients meets the fracture toughness requirements set by ASME, Appendix G with a good margin of safety. Therefore, it is concluded that the Turkey Point lli Reactor Vessel does comply with the ASME Appendix G requirement.
i t
3-1
r a e-APPENDIX A CLARIFICATION ITEMS ON TURKEY POINT UNIT NO. 3 REACTOR VESSEL - FRACTURE ANALYSIS BASED ON APPENDIX G. ASME CODE SECTION III NOTE 1. On Pages 2-14 and 2-15 the transient loadings used represent the official design specification values.
from the applicable Equipment Specification for all loading conditions except where noted.
e 4
A-1 !
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E l
SOUTHWEST RESEARCH IN S TIT U T E Post Office Drower 28510, 8500 Culebro Road Son Antonio, Texas 78284 REACTOR VESSEL MATERIAL SURVEILLANCE PROGRAM FOR TURKEY POINT UNIT NO.4 ANALYSIS OF CAPSULE T by E. B. Norris
{
FINAL REPORT.
SwRI Project No. 02-4221 l to Florida Power & Light Company
+
P. O. Box 3100 Miami, Florida 33101 June 14,1976 Approved:
U. S. Lindholm, Director Department of Materials Sciences
E 4
TABLE OF CONTENTS M
LIST OF TABLES ', 'iii LIST OF FIGURES iv I.
SUMMARY
OF RESULTS 1 IL BACKGROUND 4 III.
DESCRIPTION OF MATERIAL SURVEILLANCE 8 PROGRAM' IV. TESTING OF SPECIMENS FROM CAPSULE T 15 V. ANALYSIS OF RESULTS 33 VI. REFERENCES 42
+ '
l 11 e
r
/,
f 4 t I
, r1 n LIST IOF TABLES <-
3 Table >
Page
, I FP&L Turkey Point Unit No. 4 Reactor 11 Vessel Surveillance Materiala(13)
H Summary of Plant Operations - Turkey ' 19 Point Unit 4
, ( '
IA ).L, Summary of Neutron . Dosimetry Results 7.0 IV ' Charpy V-notch Impact Properties - Shell
~
23 Forging 1225180VA-1
)
V Charpy V-notch impact Properties - Heat 24 Affected Zone Material
- VI Charpy V-notch Impact Properties W61d 25-Metal VII , Charpy V-notch Impact Properties - A 533 26 Correlation Monitor n -
3<VIH Notch Toughness Properties of Capsule T 31 Specimens - Turkey Point Unit No. 4 .
IX Tensile Properties of Surveillance Materials 32
+
X Fluence and Weld Metal Cy Upper Shelf Energy 36 Projections - Turkey Point Unit No. 4 XI Projected Shifts in RT NDT or f Turkey Point 39 Unit No. 4 'i '
XH Proposed Reactor Vessel Surveillance Capsule 41 Schedule - Turkey Point Unit No. 4 i
1 l
iii l
lq.
c, LIST OF FIGURES Figure Pm 1 Arrangement of Surveillance Capsules in 9 the Pressure Vessel 2 Vessel Material Surveillance Specimens 12 3 Arrangement of Specimens and Dosimeters 13 in Capsule T 4 Effect of Irradiation on Cy Irnpact Properties 27 of Turkey Point Unit No. 4 Lower Shell Forging 5 Effect of Irradiation on Cy Impact Properties 28 of Turkey Point Unit No. 4 Weld Metal v.; 6 Effect of Irradiation on Cy Irnpact Properties 29 of Turkey Point Unit No. 4 HAZ Material 7 Effect of Irradiation on Cy Impact Properties 30 of Turkey Point Unit No. 4 Correlation Material 8 Dependence of Cy Shelf Energy on Neutron 35 Fluence
+ 9 Effect of Neutron Fluence on RT NDT Shift of - 37 Weld Metal iv 9
I.
SUMMARY
OF RESULTS AND CONCLUSIONS The analysis' of Capsule T, the first reactor vessel material surveil-lance capsule removed from Florida Power & Light Company Turkey Point Unit No. 4 nuclear power plant, led to the following conclusions:
, (1) The vessel beltline weld metal (0.30% copper) was more sensi-tive to radiation-induced embrittlement than the lower shell forging and heat affected zone materials (0.056% copper). As a result, the weld metal will control the future operating limitations on the pressure vessel.
(2) Based on the results from Capsule T and trend curves for like .
materials, the irradiated properties of the weld metal are adequate to meet the requirements of 10CFR50, Appendix G, through the end of Core Cycle IV.
(3) Capsule T received a fast fluence of 6. 05 x 1018 neutrons per em 2 (E > 1 MeV) during fuel Cycle I, equivalent to 1.18 Effective Full Power Years (EFPY) of operation at 2200 Mw g.
(4) The Charpy V-notch (Cy) upper shelf energy of the weld metal was reduced from an unirradiated value of 66.3 ft-lbs to 43.5 ft-lbs as a re-sult of exposure in Capsule T. The forging and heat affected zone materials retained Cy shelf energies of approximately 130 and 145 ft-Ibs respectively.
(5) The shift in RT NDT f r Turkey Point Unit No. 4 of 275 deg F l l
was based on the following data obtained from Capsule T:
4-2 Increase in RT NDT Material Crite ria of Weld Metal
. Weld 42 ft-lbs Cy energy .
2 75 deg F .
Weld . 35 mil Cy lateral expansion 240 deg F Forging ,50 ft-lbs C y energy 10 deg F Forging . 35 mil Cy lateral expansion 25 deg F HAZ .50 ft-lbs Cy energy 80 deg F HAZ 35 mil Cy lateral expansion 85 deg F i4 (6) Based on a lead factor (Capsule T flux + Predicted Maximum Vessell.D. flux) of 2.48, the fluence received by the vessel wall each -
EFPY is 2. 07 x 1018 neutrons per cm2 (E > 1 MeV). Assuming an 80%
load factor over the design lifetime of 40 calendar years, the end-of-life l 4
fluence on the vessel l.D. is predicted to be 6. 6 x 1019' neutrons per em2 (E > 1 MeV). This is approximately 30% higher than the design value of 5.1 x 1019 neutrons per em2 (E ) 1 MeV) given in the FSAR. Howe ve r, more recent calculations by Westinghouse revised the predicted fast flu-ence to a value of 6.3 x 1019 neutrons per em2 (E > 1 MeV) which is within
+
, 5% of the value calculated from Capsule T results.
-(7) Based on a 1/4T lead factor (Capsule T flux + Predicted Maxi-mum 1/4T flux) of 4.17, and an 80% load factor over the 40 calendar year
- design life (= 32 EFPY), the predicted end-of-life fluence at the 1/4T posi-tion in the pressure vessel wall is 3. 9 x 10 19 neutrons per em2 (E > 1 MeV).
(8) , Based on currently-available information which correlates change in RTNDT with copper content, the predicted shift in RTNDT of the weld metal at the vessel 1/4T and 3/4T locations for various periods of !
- vessel operation are as follows: i
. . _ . . . . o ;
L 3
f I.oc ation . Lead Shift in RTNDT (deg F) in Wall' Factor 3 EFPY 5 EFPY 10 EFPY 32 EFPY 1/4T 4.17 242 281 342 467
-3/4T 17.4 162 188 230 312 i
These values were used as the bases for computing heatup and cooldown limit curves for Turkey Point Unit No. 4. _ (Three EFPY will not be
.reac eh d until near the end of Core Cycle IV as estimated from both computer predictions and past operating experience. )
(9)' Assuming that the percent change in Charpy V-notch upper shelf energy is proportional to the ' square root of the neutron fluence, the weld metal upper shelf energy at the 1/4T position is predicted to reach '
the 50 ft-lb level at approximately 2. 7 EFPY of operation.
(10) Although the surveillance pregram is in compliance with Ap-pendix H of 10CFR50, it is recommended that a replacement capsule with additional weld metal specimens be placed in the Capsule T slot if archival material is available. An alternative is to move Capsule V into the Cap-sule T slot at the end of Core Cycle III (April 1977) and remove it for test-ing at the end of Core Cycle IV (April 1978), at which time the estimated
, fluence on Capsule v would be 8.25 x 10 18 neutrons per em2 (E > 1 MeV).
I (11) On the basis of NRC recommendations, the WOL fracture mechanics specimens have been stored untested pending development of recommendations concerning test procedures.
e
4 II. BACKGROUND The allowable loadings on nuclear pressure vessels are determined by applyirig the rules in Appendix G, " Fracture Toughness Requirements,"
of 10CFR50.(II* 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 ASTM E208;(3)
(2) 60 deg F below the 50 ft-lb Charpy V-notch (Cy) temperature:
(3) 60 deg F below the 35 mil Cy tempe rature.
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.
5 to neutron fluences in excess of 10I7 neutrons per em2 (E > 1 MeV). (4) ,
1 Also, it has been established that tramp elements, particularly copper and phosphorous, affect the radiation embrittlement response of ferritic mate rials. (5-7) Unfortunately, there are disagreements concerning the i
relationship between increase in RTNDT and copper content. For example,
. {
1 Regulatory Guide 1. 99I7) proposes an adjustment to RTNDT proportional to l the square root of the neutron fluence. Westinghouse Electric Corporation, in their comments 'on Regulatory Guide 1. 99(8), feels that the proposed re-lationship overestimates the shift at high fluences (above 103 9) and under-estimates the shift at low fluences (below 1019 ). On the other hand, Com-bustion Engineering, in their comments on Regulatory Guide 1. 99(9), sug- i gest that the proposed relationship is overly conservative at fluences below 1029 neutrons per em2 (E > 1 MeV). There is also disagreement concerning the prediction of Cy upper shelf response to exposure to neutron irradiation. (7-9I _,
. It is important to resolve these questions because the analysis of reactor vessel material surveillance program data requires that estimations be made of shifts in RTNDT and Cy upper shelf energy at fluences other than that received by the surveillance capsule.
In general, the only ferritic pressure boundary materials in a nuclear plant which are expected to receive a fluence sufficient to affect RT NDT are those materials which are located in the core beltline region of the reactor ,
pres sure vessel. Therefore, material surveillance programs include spec- i l
imens machined from the plate or forging material snd weldments which are l l
l 1
- \
l
I s
6 located in such a region of high neutron flux density. ASTM E185(10) describer the current recommended practice for monitoring and evalu-ating the radiation-induced changes occurring in the mechanical properties of pressure vessel beltline materials.
Westinghouse has provided such a surveillance prograrn for the l Turkey Point Unit No. 4 nuclear power plant. The encapsulated Cy spec-imens are located on the O. D. surface of the thermal shield where the fast f neutron flux density is approximately twice that at the adjacent vessel wall i
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, some capsule fluer>ces may be less than the maximum vessel fluence in a corresponding exposure period.
For example, Capsule T (removed during the 1975 refuelling outage) is re-ported to lead the maximum exposure point on the vesselI.D. by a factor of 2.48 while Capsule V (scheduled to be removed at a later refuelling) will receive about 79% of the fluence accumulated at the point of maxirnum vessel exposure. The capsules also contain several dosimeter materials for exper-imentally determining the average neutron flux density at each capsule loca-tion during the exposure period.
The Turkey Point Unit No. 4 material surveillance capsules also in-clude tensile specimens as recommended by ASTM E185. At the present time, irradiated tensile properties are used only to indicate that the materials l
l i .. . .. .
7.
tested continue to meet the requirements of the appropriate material 1
l specification. In addition, the material surveillance capsules contain {
r wedge opening loading (WOL) fracture mechanics specimens. Current ,
technology limits the testing of these specimens at temperatures well <
\
below the minimurn service temperature to obtain valid fracture mechan- I 1
ics data per ASTM E399Ill) " Standard Method of Test for Plane-Strain Fracture Toughness of Metsllie Materials. " However, recent work re- t 1
ported by Mager and Witt(12) may lead to methods for evaluating high- 4 toughness materials with small fracture mechanics specimens. Currently, the NRC suggests storing these specimens until an acceptable testing pro-cedure has been defined.
4 This report describes the results obtained from testing the contents I
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 1975 refuelling outage as well as predicting the changes expected to oc-cur at selected times in the future operation of the Turkey Point Unit No. 4 power plant. I j
i i
i t
8
8 i
HL DESCRIPTION OF MATERIAL SURVEILLANCE PROGRAM The Turkey Point Unit No. 4 material surveillance program is de-scribed in detail in WCAP 7660II3I, dated May 1971. Eight materials sur-veillance capsules (five Type I and three Type H) 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 cen-ter of the core. The neutron flux density at the capsule location leads the maximum flux density on the vesselI. D. by the factors given in parentheses following the capsule identification letter in Figure 1. The Type I capsules each contain Charpy V-notch, tensile and WOL specimens machined from the two vessel forgings located at the core beltline plus Charpy V-notch opecimens machined from a reference heat of A533 Grade B, Class I steel utilized in a number of Westinghouse surveillanc,e programs. The Type H 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 forging material. Capsule T, one of the Type H capsules, was removed during the 1975 refuelling outage.
Babcock and Wilcox Co. supplied prolongations from two 7-7/8 in, thick forged rings (Heat 123P481VA-1 and 122S180VA'-l produced by Beth-lehem Steel Co. ) of SA 508, Class 2 steel used for the FPL Unit No. 4 reac- l tor pressure vessel intermediate and lower shell course, respectively, and i a weldment which joined sections of the two forgings. Correlation monitor 1
i
)
0 i
T (2. 48) 9 i S (1.61) 1
- 7. (O. 34 ) #
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1 10* 10*
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180'
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/
4
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-. U (0. 4 9 )
/ - W ( 0. 3 4 )
Y (0. 49) // - X (0. 34) 90*
Reactor Vessel Thermal Shield Core Barrel Note: Numbers in parentheses indicate lead factors for the vessel 1. D. I FIGURE 1. ARRANGEM ENT OF SURVEILLANCE CAPSU LES IN THE PRESSURE VESSEL
6'
)
I 10 i l
material was supplied by the Oak Ridge National Laboratory from plate material used in the AEC-sponsored Heavy Section Steel Technology (HSST) l Prog ram. This material was obtained from a 12 in. thick A533 Grade B, .
1 Class I plate (HSST Plate 02) which has been provided to Subcommittee II j
(
of ASTM Committee E10 on Radioisotopes and Radiation Effects to serve I as correlation monitor material in reactor vessel surveillance programs.
1 The plate was produced by'the Lukens Steel Co. , and heat treated by Com- '
busion Engineering, Inc. The chemistries and heat treatments of the vessel t i
surveillance materials contained in Capsule T are summarized in Table I.
All test specimens were machined from each of the materials at the quarter-thickness (1/4 T) location. The base metal tensile and Cyspecimens were ,
i
)
oriented with their long axis parallel to the principal working direction; the ~
Cy notches were perpendicular to the major forging surfaces. The WOL i
specimens were machined with the simulated crack perpendicular to the '
principal working direction and to the forging surfaces. All mechanical
]
test specimens, see Figure 2, were taken at least one plate thickness from i q
the quenched edges of the forging material. l j
Capsule T contained 32 Charpy V-notch specirnens (8 from each of the vessel beltline materials plus 8 frorn the reference steel plate); 4 tensile spec-imens (2 forging and 2 weld metal); and 4 WOL speciment (2 forging and 2 weld }
J metal). The specimen numbering system and location within Capsule T is shown in Figure 3.
I l
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]
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, .t .
E , .'q
' TAllI.M 1 FPkL Turkey Point Unit No. 4 Reactor Venmut Surveillance MhterininII3I i
Heat Treatment History Lowe r,Shell 1550 F 1/4 hours - water-quenched
~
""~"#~ *
(Heat 122S180VA'-1) 1125 F 1/2 hours - furnace-cooled to 600 F Weldment 1125 F .10-1/4 hours - furnace-cooled to 600 F. 1675
- 25 F. - 4 hours - air-cooled Correlation Monitor .1600
- 25 F - 4 hours - water-quenched 1225
- 25 F - 4 hours - furnace-cooled- 4 1150
- 25 F - 40 hours - furnace-cooled to 600F i
Chemical Compos'ition (wt-%) Lower ; Shell Correlation Element 122S180VA- 1 - Weld Metal Monitor C 0.21 0.098 0.22 Mn 0.67 1.44 1.45L P 0. 011 0.014 0. 011 S 0.009 0.011 0. 019 31 0.23 0.50 0.22 Ni 0.70 0.60 0.62 Cr 0.31. 0.14 - V 0. 001 0.002 - Mo 0.56 0.36 0. 53 Co 0.015 0.008 - Cu 0.056 0.30 - Sn O.008 0.003 -
]
Zn '0.001* -
- 1 A1 0.008 0. 014 -
N2 0.002- O. 012 -
.Ti 0.001* 0.001* -
)
Pb 0.001* 0.002* - I
,As 0.005 0.005- -
B 0.003* 0.003* - Zr 0.004 0.004 - W 0.002* 0.001* - Nb 0.001 0.001* - Ta 0.002 0.002 -
* - Not detected. The number indicates the minimum limit of detection.
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(b) Tensile Specimen l.45 __"51 _M 375 0 - usO ,,,, my= IE
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.0667 (c) Wedge Opening Loading Specimen FIGURE 2. VESSEL MATERIAL SURVEILLANCE SPECIMENS
p d T C I U oo( l8 lJ 4 7 4 6 CC i1I L S R. u i l ga
' : r - 3 I
i C j g ,t o s 3 7 6 i t C S R I n 2 2 o 7 6 y 3 R i M C M 1 1 F 7 6 l l S R
+ a 9 e 6 7 l 1 i 5 i S r s e n
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+ f n
e r e S P A 7 6 e C C y S 9 R 9 b R 3 N I 6 5 3 S S R i 5 R A E L 0 = T O 2 S l E W f R M y 8 6 S 8 5 R b d e I S O i S l D r C 7 I l a o 7 5 l l t D t 6 l e i n S R - a N o l W M A mrk e co d S
) M 1 i 6 s t l
l l e l N 1 d oel B e C l F D l s s W E ( 0 oo 9 3 4 e = M GC i8IA i1IL 5 Z 2 ! V I W H l l W C i D ll :: l 3 : E P i y N ll egL i J 2 2 I C w H S 1 F L F - 6 Z O O A W W T i H 2 2 = N 2 2 E y W H l H M E C 1 2 W 1 2 H > G N
- L O W 5
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+ 1 A
A R R W V A _ 0 l e 6 8 1 . l i s W ' 1 m S 3 _ o r n
+ 2 E
1 t o e 5 t 2 t T o 1 R i n W 0 0 B t U o 2 2 a G . W H l e I M y l e H F U s _ F C 9 9 s = 1 1 e 9 I l V S oo ilia 7 8 1 8 l W 1 CC jI e L 5 W i t o _ G:::: l l f I l T . u lleJ y 7 7 _ C i It L 1 1 C W H I _ e _
l I 14 Capsule T also was reported to contain the following dosimeters for determining the neutron flux density: i Target Element Form Quantity Copper Bare wire 2 Nickel Bare wire 1 Cobalt (in aluminum) Bare wire 3 i Cobalt (in aluminum) Cd shielded wire 3 Uranium 238 Cd shielded capsule 1 Neptunium 237 Cd shielded capsule 1 In addition, drillings were taken from five Cy specimens to serve as iron dosimeters. . Three eutectic alloy thermal monitors were inserted in hole s 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 capaule) wis 1. 75% Ag, 0. 75% Sn and 97. 5% Pb having a melting point of 590 F.
l
I 1 15 IV. TESTING OF SPECIMENS FROM CAPSULE T l l The capsule shipment, capsule opening, specimen testing and report-ing of results were carried out under a Quality Assurance Plan prepared by Southwest Research Institute (SwRI) and approved by Florida Power and Light Company (FP&L). This procedure is on file at SwRI. Applicable SwRI Nu-clear Project Operating Procedures which were employed in this program include: (I) XI-MS-1, Determination of Specific Activity of Neutron Radiation Detector Specimen. (2) XI-MS-3, Conducting Tension Tests on Metallic Materials. (3) M-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
+
l t Metallic Materials. i ! (6) XI-MS-6, Determination of Specific Activity of Neutron l Radiation Fission Monitor Detector Specimens. l l t Copies of the above documents are on file at SwRI. l Southwest Research Institute prepared a procedure for the removal of the caps 61e from the reactor vessel and the shipment of the capsule to the SwRI laboratorie s. The procedure was reviewed by FP&L, which resulted in several modifications. SwRI contracted with Todd Shipyards - Nuclear i J
i I 16 Division to supply appropriate cutting tools and a licensed shipping cask. Todd personnel severed the capsule from its extension tube, sectioned the l l extension tube into three-foot lengths, supervised the loading of the capsule ! and extension tube materials into the shipping cask and transported the cask i to San Antonio. t 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 i Bridgeport vertical milling machine set up in one hot cell. Before milling i 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 r>o that capsule location was identifiable. After the disassembly had been completed, the specimens were carefully checked for identification and location as listed
- in WCAP 7660.(13)
Each specimen was inspected for identification number, which was checked against the master list in WCAP 7660. No discrepancies were found. j 1 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 exceed 579 F. Two discrepancies were noted concerning the dosimeter materials. The fission monitor materials were reported to be in a 3/8 in. O. D. sealed l 4
17 brass tubes. The tubes were, in fact, 1/4 in. O. D. x 3/8 in. long; one ap-peared to be brass and the other was a soft silvery metal such as aluminum or stainless steel. The " nickel" monitor did not exhibit a 58 Co peak. A qualitative analysis with a microprobe showed that this dosimeter was high in coppe r and contained no nickel. Since it was coated with a silvery-colored material, it was concluded that the chemistry of the dosimeter was not known sufficiently well to report the results. 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 equiprnent was accomplished with appropriate standards and an interlaboratory cross check with two independent counting laboratories on 60Co ,54Mn- and 58 Co-containing dosimeter wires. All a.ctivities were corrseted 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 integral of the energy-dependent microscopic activation cross section and the neutron flux density. The relationship be-tween ATOR and ASAT is given by: i man ATOR = (1. e-ATm) (e-Atm) ASAT m=1 ; where: .A = decay constant for the activation product, day-1; Tm = Equivalent operating days at 2200 MWt for op-erating pe riod m; tm = Decay time after operating period m, days
. e
~
18 The Turkey Point Unit No. 4 operating history up to the 1975 refuelling shut-down is presented in Table II. The specific and 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. The average flux density at full power is given by: A SAT Wi whe re: @ = energy-dependent neutron flux density, n/cm2 _,,c; AISAT = saturated activity of the.ith activation product at full power, dps/ target nucleus; di = th tar-spectrum-averaged cross section for the i get nucleus, em . 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. The neutron flux density was calculated from the 54Fe (n, p) S4Mn reaction because it has a high energy threshold and the energy response is f i well known. The energy spectrum from Capsule T was assumed to be the l same as that previously reported for Capsule T in Turkey Point Unit No. 3. II4) The calcula.ted flux spectral distribution from 1. 74 MeV to 10. 0 MeV was then combined with Helm's(15) differential cross section values for the 54 Fe (n, p) S4Mn reaction to derive the spectrum-averaged cross section i for neutrons with energies greater than 1 MeV as follows: O .
m t e imdo Ti y r 7 1 4 9 5 3 a e 1 1 1 7 9 3 3 2 0 7 5 4 4' 6 3 0 1 5 6 4 0 cP 6 6 6 6 5 5 5 4 3 3 3 2 2 5 1 1 1 8 e r De t f a - m t T n , l es ay 9 8 8 1 3 6 va 1 4 5 4 1 3 2 7 9 6 5 0 0 1 2 8 7 7 1 8 1 1 5 7 6 0 4 iDu 1 1 6 q 1 3 9 2 7 3 6 0 4 2 4 6 9
. 1 2 3 4 6 1 Er e 1 3 7 4 2 2 7 3 4
p O r) et s n wd o w o 9 6 6 6 8 5 4 3 i t PM ( 6 9 2, - 5 , 3 0 - 0 5 0 0 7 0 0 1 6 5 4 6 2 5 0 5 4 1 4 0 5 7 0 7 2 a4 r - 4 - - 9 - 0, - 3, - 1, - 5, 5 - 8 - 7, 2 r t o tu 3 2 3 7 1 8 1 2 3 ei t p 2 5 9 8 8 3 8 9 pn c 2 2 6 7 0 2 7 5 9 5 4 t 6 6 . I OU au 1 1 1 9 ReO I t t E nn L l ai o B PP g L A f y n A T oe is t T yk a y 2 5 - l 1 - 1 O r r r a 1 2 2 59 37 7 0 9 au eD p I 4 4 3 3 7 T mT O m u S n w oys d t a '- 1 4 1 - 1 1 1 uD 1 3 3 2 6 4 - h S 3 4 p 333 3337 4444447444 55 t o 3333777333777/4444777777/777577 7777///777///67777//////7///7// S // 1 //489///3441 2791 1 251 222//3 / / 3351 //7851 1 231 802247/00 1 // / 22/// / 61 3 t s e /// 7777777899991 /////////1 1 / / / / 1 1 244445 5891 1 1 1 1 1 3
//////01 a
D 3 t 3 333 3337 4 4444447444 5 r a 733337773
/ 3 7777///777///67777//////7///7/ 37 77/4444777777/77757 S
t 9////489///34 41 / ///7851 802247/0 1 1 2791 1 251
/////////////1 222//33351 1 231 1 ////61 67777777899991 1 / / ////// //01 22//
1 1 2444455891 1 1 1 1 1 r.d oi 1 2 3 4 5 e r 6 7 8 9 0 1 2 3 4 pe 1 1 1 1 1 5 1 6 1 OP
-- . . _ _ - e 20
, TAB LE III Summary of Neutron Dosimetry Results Monitor Activation Measured Activity Saturated Activity JdentificationI") Reaction (dos / mg) (dps /mg)
W- 17 We(n, p)S4Mn 4.27 x 103 7. 52 x 103 W-21 4. 74 x 103 8. 36 x 103 S-67 5. 00 x 103 8. 80 x 103 S-69 4. 89 x 103 8. 61 x 103 S-73 Y 5. 01 x 103 8. 83 x 103 Cu (Top) 63Cu(n )60Co 1. 3 0 x 102 9.18 x 102 Cu (Bottom) 1. 21 x 102 8. 55 x 102 i Ni (Cente r) 58Ni(n, p)58Co (b) (b) Np (Center 237 Np(n, f)l37Cs
- 2. 60 x 103 (c)
U (Center) 238U(n, f)l37Cs 2.19 x 105 (e) Co (Top) 59Co(n, y)60Co 1. 33 x 107 9. 44 x 107 Co-Cd (Top) 5. 08 x 106 3. 59 x 107 Co (Center ) 1. 47 x 107 1. 04 x 108 Co-Cd (Center) 6. 60 x 106 4. 66 x 107 Co (Bettom) 1.44 x 107 1. 02 x 108
- Y Co-Cd (Bottom) 5.16 x 106 3. 64 x 107 j i
l
)
(a) See Figure 3 for location within capsule. { j (b) Monitor material unknown copper alloy. (c) Not calculated. I 1
21 10MeV try,(E)@(E)dE 1.74 f e, (> 1 MeV) = 10 MeV
@(E)dE
. 1.00 where: &c, (> 1 MeV) = the calculated spectrum-averaged cross section for flux > 1 MeV, em2, The resulting value obtained for fast (> 1 MeV) neutron flux density at the Capsule T location was 1. 62 x 1011 neutrons /cm2-s ec. Since Turkey Point Unit No. 4 operated for an equivalent of 431.47 full power days up to the 1975 refuelling octag;e, the total neutron fluence (> 1 Mev) is equal to 6. 05 x 1018 neutrons /em2 . This is about 30% higher than the design value given in the FSAR. However, more recent calculations by Westing-AouseII4) revised the predicted fast fluence to a value of 6.3 x 10 19 neu-trons per em2 (E > 1 MeV) which is within 5% of the value calculated from Capsule T results. Much of the early work published on the radiation-induced embrittle-ment of ferritic steels correlated shifts in ductile-brittle transition tempera-ture with neutron fluence calculated on the assumption that the neutron ener-gica were diatrihoteel accor< ling tn a rianion neutron spectrum. To provide information for reference only, the Capsule T fast flux density based on a fission-spectrum cross section was computed as follows:
22
- 1. 34 x 10- 3 4 Pf , > 1 MeV = 98.26 x 10-27 = 1.36 x 10 31 neutrons /cm 2 .,,c, i
The corresponding value of neutron fluence is 5. 09 x 1018 ne utrons /cm2> 1MeV. . The irradiated Charpy V-notch specimens were tested on an instru-
.mented SATEC impact machine. The test temperatures were selected to
.. develop the ductile-brittle transition and upper shelf regions. The unirra-i,
.diated Charpy V-notch impact data reported by Westinghouse (13) and the data obtained by SwRI on the specimens contained in Capsule T are pre-wented in Tables IV through VH. The Cherpy V-notch transition curves for the forging, weld and HAZ materials and the correlation monitor ma-wrial are presented in Figures 4 through 7. Instrumented impact data are presented in Appendix A. The radiation-induced shift in transition tempera-1:ure is indicated on each plot according to procedures defined in Appendix G tri 10CFR50. III A summary of the shifts in RTNDT and Cy upper shelf ener-gies for each material are presented in Table VHI.
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 autographic recording equipment. The forging tensile specimens were tested at RT and 300 F; the weld specimens were tested at 300 F and 550 F. The results, along with tensile data reported by Westinghouse on the unirradiated materials (13), are presented in Table IX. Testing of the WOL specimens was deferred at the request of Florida Power 8e Light Company. The specimens are in storage at the SwRI radiation laboratory.
* * , P
t 23 TADLE IV Charpy V-notch Impact Properties Shell Forging 122S180VA-1 Late ral Neutron Fluence Test Temp. (") Energy Shear Expansion (E > 1 Mev) (deg F) (ft-lbs ) (%) (mils ) ' . Unirradiated(b) -100 3. 5 0 2
-100 6. 0 0 3
-100 3. 0 0 4
-50 6. 0 0 5
-50 24.0 14 22
-50 12.0 5 10
-20 14.0 3 12
-20 70.0 36 59
-20 36.0 14 29 10 59.5 27 50 10 74.0 33 60 10 73.0 39 60 4 110 133.0 100 92 110 134.5 100 92 110 108.5 73 86 v
210 132.0 100 88 3
, 210 131.0 100 89 210 132.0 100 88 e 6. 05 x 10 18(* } 81 (S74) 97.1 100 74 '
40 (S73) 92.1 60 73 10 (S72) 43.1 10 33
-20 (S71) 35.6 10 27
-50 (S70) 12.5 nil 9 210 (S69) 134.6 100 96 g
25 (S68) 53.0 10 42 10 (S67) 40.5 10 34 (a) Specimen number in parentheses following test temperature. (b) WCAP 7660. (13) (c) Capsule T.
Y 24 TABLE V Charpy V-notch Impact Properties Heat Affected Zone Material Late ral Neutron Fluence Test Temp. (*) Ene rgy Shear Expansion (E > 1 MeV) (deg F) (ft-lbs ) (%) (mils ) Unirradiated(b) -300 15.O' 9 9 i
-300 5. 0 5 4 l
-300 - 4. 0 5 2 .
-250 50.0 20 31
-250 10.0 9 7
-250 2.5 5 5
-200 36.5 14 22
-200.. 80.0 36 52
-200 76. 0 33 49 l
-100 103.0 59 68
-100 85.0 51 54
, -100 90.0 39 60
-50 107.0 68 74
-50 107.0 51 74
-50 161.0 100 89
- 10 171.0 100 87 10 169.0 100 90 110 151. 0 100 94 110 156.0 100 88 llo 140.0 100 91
- 6. 05 x 10 38(C) 81 (H24) 144.8 100 93 10 (H23) 142.6 100 93
-50 (H22) 100.6 50 71 l
-200 (H21) 5. 8 5 2.5
-150 (H2 O) 40.0 10 25
-100 (H19) 15.2 10 7 3 -90 (H18) 61.6 20 41
-100 (H17) 49.5 20 32 I
(a) Specimen number in parentheses following test temperature. (b) WCAP 7660. (13) (c) Capsule T.
l 25 TABLE VI Charpy V-notchImpact Properties
-Weld Metal Lateral Neutron Fluence Test Temp. I") Inergy Shear Expansion (E > 1 Mev) (der F) (ft-lbs ) (%) (mils )
Unirradiated(b) -100 8. 0 13 6
-100 7. 5 13 10
-100 11.0 18 11
-50 15.0 28 17
-50 12. 0 23 12
-50 12.0 28 13 10 29.0 46 30 )
10 39.0 55 40 10 36.0 47 32 1 50 52.5 84 50 50 51.5 90 51 50 .51. 0 77 52 110 58.0 99 60 110 60.0 98 75 110 61.0 100 65 210 66.0 100 68 210 69.0 100 71 210 64.0 100 65
- 6. 05 x 10 18(c) 81 (W24) 13.0 5 10 110 (W23) 15.3 10 13 160 (W22) 21.4 15 18 210 (W21) 27.7 50 25 300 (W20) 43.5 200 43 i 475 (W19) 43.4 100 42 375 (W18) 43.5 100 45 I
300 (W17) 41.75 200 42 (a) Specimen number in parentheses following test temperature. (b) WCAP 7660. (13) j (c) Capsule T. I
26
\
w TAB 1 E VII Charpy V-not.ch Impact Properties A533 Corrulation Monitor 2mte ral Meutros Fluence Tour Temp.I*) Energy Shear Expansion (E > 1 MeV) for F) (It-lbs ) (%) (mile ) Umirradiated(b) -50 5. D 9 3 50 5. 0 9 5
-50 3.0 9 4
-20 6.5 9 6
-20 9. 0 13 10
-20 6. 0 13 . 9
^
10 12. 0 23 19 10 J4. 5 23 14 10 13.5 23 14 40 22,0 33 23 40 36.0 29 32 40 J5. 0 29 32 45 58.5 43 51 85 41.5 41 42 85 52.0 42 45 210 82.5 58 60 llo 85.5 67 71 120 63.5 55 54 160 208.5 84 72 160; 81.0 85 69 160i 1 09. 0 87 79 230 137. 0 98 54 210 115.0 ~ 98 88 210 121. 0 100 87 30G 125.0 100 87 300 217.5 200 83 U B00 127.0 100 84
- 6. 05 x g3 8(c) 41(R64) 12.2 10 10 40 (R63) 8.1 5 5 10 (R62 ') 4.3 ni! 2 110 (R63 ) 23.3 15 17
, 160 (R60) 41.9 20 33 210 (R59) 60.5 50 49
, 300 (R$8; 92. 0 100 83 375 (R57) 91.8 1 00 80 (a) Specimen number in parentheses following test temperature.
(b) WCAP 7660.Il3 3 (c) Capsule T.
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' i _
+ 5100 9..'! t' '
' O Unirradiate d :
{ + T- 4 ' jf '[I ! " U! h -- e Irradiated at 547'F and - N
,. p. .
.!.[; ,iil[. ..)._p h.l i.. ..l;:_. ! ..I.!
L._.
'[6.L05 x 1018 neutrons /cm2 --
. j. g .ep.. t. ,
.7e . .g,
. y,; ,li, ..g. l. [
i. (F > 1 MeV) - 4 ...t7...lT [.....L.{_ ,._ L., @,...
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,,.. '50 10 . Capsule T i .
db. ..... .'.-j~._.r
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l
*- 4--
-t, }! t....:.[-{ f
) . . . .
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y} , .e ,.. 71 . ,. . . . .l . . . l. . ) t.
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Fl.t.i. .--_+. 1
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! ,
- a' t '
1, e- a p t, ..1..F . .t, .t .1._Lp... M r. ,; i ; i i ci
}
~100 0 100 200 300 400 500 Temperature, deg F
** r t,,.
100 .. .. l . 2 . ';t
.i., i i i, l ,; .;,i.j i .I1 3 i ii
. , p i l . p . l l ji i ; (. j i.jI {. ..i j. . } .I L j .f. .' .y }..
llg.. s # ,,,, ,, e ;e
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, . i; ,II! i!.i. i 1... 1
___ {. 6. ; i I.I j i ,; } '. .4
%l% i ' i i ' '-
.l . l.
--y!- Unirradiated
; . i .. - - I _- - ..L
/O!! l.--.-Y-
~
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' - t i 'l 1- t 4-- -
-f_ > ., .1 p.*-. ,.i j ; .:
- .f i ,
ij T l'?.4
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. . +- .
.. p. ; .6.
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, J . . .; . . . l . .g lllll , .... . .,. .
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t- --- ; t-+'-
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.. w 25 / -
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!lii +I . j.!j-
!;i;j!!l 4 ,i i i. e Irradiateo at 547'F and --4 i
l
'. I 6. 05 x 1018 neutrons /cm2 "
i .1 : /; o . i j ll
.I 4 - ; ; . i. .,
It i
>:.ll
. l . .
!!..! i (E > 1 MeV)
I, , i
?
)
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i i;. . i ' I j I, i,, j, i, l ;7 .,-l,.i, . .
.,t..I.... -. d. b. .. ...
-100 0 100 200 300 400 500 Temperature, deg F i
TIGURE 4. EFFECT OF IRRADIATION ON Cy IMPACT PROPERTIES OF TURKEY POINT UNIT 4 LOWER SHELL FORGING 7
l;o [l. 100 28 i.l Tf]!ll! i '2 '; uttu .t a o1 >'ii'i4 +>' j'; I
' l l1 i . ;l . Code: _
l}[l!,lll
~ '
I 0 Unirradiated l l! Unirradiated i e Irradiated at 547'F and
,a, 75
.} '
j, . ,3 j lH
}.lll1ll
; i,,, ;,
i l 6. 05 x 1018 neutrons /cm2, atC Il
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5 ['.- -
~ ~ ~
e 7trr -- r rr , i
;.11 -f '.
.} h.. l}l l! . t , .,-l g ' .g 44-- .i'[{! NE
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- }j -100 0 100 200 300 400 500 i., Temperature, deg F 100.p...p ..
- j q.,I4 t 'i ' i -
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__ . i . 'q . '. .'. '
..pt l -
Unirradiated y r,
, ~
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= 75 i
J. ' -]' i iL.IIT ' ' ' _u ii
.i:H 6. 05 x 1018 neutrons /cm2,h
- (E > 1 Mev) p' g 4-H ri rt T..+FY -l + t : h, .,iiu, t }+ -i A[._ . I.u...L_: s --. qa i+
i ip! I !
. .,a. %.a '
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~
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~
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8 i: i. i i l j. i _it.Hjll
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.' ,e- . ; i ,i ' lli. !'!,:
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.; T,.j1. ..
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.i
.q ~ ., .
~. t- -
- . + .-
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) Temperature, deg F l, FIGURE 5. EFFECT OF IRRADIATION ON Cy IMPACT PROPERTIES OF TURKEY POINT UNIT NO. 4 WELD METAL l j
r . l Ir :
*L -
2 00 1 I 7 t
,.cl. i l
..l ll
,.... j i.',. j, t
'i-
- i
. i '
i-
. c . .:
'l l. l. ;i tt:
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, i . .! -
Unirradiated ' t li ili ' ist i ' , - i' ' '
,li ,..i,
,l,ij i 'l .
ie i, , i . .., ,I
.' ; e i !je. , , !. :. i' .L.i. L'.J Z t
i a ll
,i i .
l Jlii. . .
'Q.
.i : -
. s a i,.
i . ,t f .i , lr +- y..-
.. r -
'I,,! . .
9 . i r.t t 150 - _fl - I . it.. . e !l .t, j a 3 i ' j . 8 .,l ,
. ! i ;.1 ; .
.. i
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p e!5 . . . ~ .
...,lIl;j...l.
i .
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.o,..... fl'... ..-
a j il;t.t. f
), f2 ri.,, ;rt ...l.
' - i, ,.i { .r --
, ,,i ,y '! . :' .; :
d l l } ! r, ! t l lI I -
- /: . Code.
l ' l: I rt t
.}.,IiI i
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o .i ? .i.
. a 1iii F; l -
0 Unirradiated Iii i ! l I me 100 w a !i >! -
,, y ,i ,!; ;l ii (I 8..
e ' Irradiated at 547'F and
,i:;i , n . . . / . .'i j I-'
. s c J. L.. .p.j ... .. .. . l
; ie w p.; .$.. 5-
>,4 -. .i Il 6. 05 x 10 18 neutrons /cm2
.i p-i . .; -4 ; d. .iM...;:; .I l . i . ., .i .i 4
.I
., t>
1 /i (E > 1 MeV) 2
-p;;..L--..l,.!. .+ i .l lI .. . ...
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+t
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, Fl
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. ;.'.3 1 .- . . . , . . .. . j .
.' N. .,
. Capsule T '
4 r
. . .iL 1.L.1
_[.;..t:. 9 105'- /.' 1 ' ' , >, -> - + ,! H, '- -.:J-w1.
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8
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. {. r . s .e 4.9
); -300 -200 -100 0 100 200 300 Temperature, deg F
' 100 I
-.. ..i..., , , - l .i..i. .!
, ; i l !.l .p' .a
. . .a.
ii
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Unirradiated - I
i ii ' e .,.- .- -e .
'.r.. i.. . ,H q d --:+i.. r- F H..i i
--.,t+- -l '-- ! .
- . . . . oi #L e .f:-c -.
-- r ,
1 e- . . .; . , - - m_ ; , T.
.,, 4.e 4 C t l., .1 ,.,.- .
94.. aa !iL 7 1 l ,.4 ; y 3 -
..,.y...
p 4 .i , l e ; 7
/- .
.-y : ,a
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. . p;. + pr u . .. 7r.. . -. ~t
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e 75
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. .. .. .o.
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c I~{.4 ..e. F !r; .i.., . . fj , . . g,.i a
. .-4.
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o m.. m.. . . i ,i .."r .l ' -
'+ *-
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..;i,iI:. i v.Ii j.. t- .'.- i W i ; p .j .
. .; i
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.h ..;-a- +! '.-,'
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. i.. , i . . .
-i - 85* M'/
e
' ;o. .
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. ' Code: +
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g 4 . 7.g,. I ii. i .,i.. ; . . r. l (,.lj,!, 1 . . . .. . , 2 l
.: 14 -r
. .. i ] r . rr.. . .f r N- t-
./
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.. .r.'
. . .-.!---.1.J . A .,L.. r-Ly -
4 l
-300 -200 -100 0 100 1 200 300 l
Temperature, deg F
, FIGURE 6. EFFECT OF IRRADIATION ON Cy IMPACT PROPERTIES l ' t i
OF TURKEY POINT UNIT NO. 4 HAZ MATERIAL ; l
200 ,, ;,r i. ri,, . i N. lj;l . i l ll,'I t.i .',;.lil ! I , iCode:
. i ,i i i .i, i. .
l ll l ; i i , o Unirradiated ) 7.
' PI'*Ii ;I',...
l i l j ,' l' il l i l. : l r'l et e Irradiated at 547'F and 1 flt,: ! ll t ;l '! i'
'.l.; 6. 05 x 1018 ne utrons /cm2 I ISO !
Unir radiate d [~ (E > 1 MeV)
. i.,l;ij i tj !, !, ;
4;. i.' lit i i
',;..;g..,
, i li;. ..
. !-j _.
o g.. . .
, . 1 , .
- j. . .
^',.ll
.,t!-ll I':
e.. '. 'h lIi -l i.i i t- - a . . i ,. ,l
._p j.i.. l,il . . .
g {;i 14 .J.6, 6l t .. , jl j..
.. b.l t4j. .
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g il. !l.. l{.$ ',-, 'D. .ii, ] Of .9 .). !.
.t .il .
7
.. .r r- . l-L! : ..
r -l ' l' 3" -'- ' Y'-"7' f- *-tht-t-I. no 100 'T'' :' l'" ' l '- ;b T7+-* "t-t "Y y '
- i -l-f- -f*-- I'
! *: *r ~L! .
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.u. t.I . . . ...... . .. o l. ;:
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. . > .6 i .,7, .i
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0
-100 0 100 200 300 400 500 l Temperature, deg F 10 0 ;i...
.q-i . . . 3..- . . p .. . ;
,. j. j.4 ;+ ;
, i,s i: .;
.l .. ,ri. ;. r i.,.4..rq- i F-
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- i. , .. i . .
. ..r.'.. L,. .,. .;.
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q .. .l.pg ... 3 Unirradiate d - t ' : i,! M- ' t' - A ' 0 ' ' F 'L t' l i ' -^ J-'l ' : '
..'t. l' H.
'! j 't'*'
'..g.p :*' L . ,i .--
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, t
., 1 ? ., $. i;e -'i , i,,,,,',,,.~L.4.. .
... . .. ..l . L. . 9 ., . . l. 4 . , fs. , . . . . e j .
- . .. + . L ,.4 .p. ..9 . ,. . ; . .. .
..g........... ,
j . .. .7... ... r ~ e 75 '
+gi
_i ,
. j ,o. l im . : L ;;..,".,.;gq+;ty_'PT' E
...l.;..,!,...e .'.'. ',,' .. i.o .,
4 ' /' 1.'./, i Code: o , 1
.i . .ae.. .
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1 /. g
. i . .
o -
.- o Unirradiated
'4c'.' .
, /, .
,i . !
g . . Ii '
;. :: o i e Irradiated at 547'F and g, 50
.u . . . .
11' .L -
. L /, ! ..:' 6. 05 x 1018 neutrons /em2 1
.1. ... . . i y
i
.L ; (E > 1 MeV) i.
M.. p ;4st.< .h..mp-,!. i.,. . i
- n. i. .F
.n. ., _. 7 i .r,y ;,;
. . . .o .f /. . . , . , . .
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25
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-100 0 100 200 300 400 500 Temperature, deg F 1
FIGURE 7. EFFECT OF IRRADIATION ON Cy IMPACT PROPERTIES l OF TURKEY POINT UNIT NO. 4 CORRELATION MATERIAL F
- i j
31-TAB 12 VIII
)
Notch Toughness Properties of Capsule T Specimens Turkey Point Unit No. 4 Forging Weld HAZ Correlation i 122S180VA-1 Metal Mate rial Monitor
, 50 ft-lb Cy Temp. (deg F)
Irradiated '10 300(*) -100 185 Unirradiated J 2 5(a ) -180 85 AT 10 275 80 100
' 35 mil Cv Temp. (deg F)
Irradiated 10 255 - 95 165 ! Unirradiated -15 15 -180 55 AT 25 240 85 110 Upper Shelf Energy (ft-lbs ) Unirradiate d 132 66.3 170 123 Irradiated 134 43.5 143 _%2, AE - 22.8 27 31
\
i (a) 42 ft-lbs Cy Temperature (deg F).
- 11 .
N n oa i t e ) 422844 02 O5 5893 36 c r% . . . . . . 877788 duA( 87 5 44281 72 666666 66 66665 5 55 e in R - n i o l t a a) 227366 33 5I 7305 85 t . . . . . . . . o g% 441 224 99 34231 0 89 n To( l 222222 1 1 222222 1 1 E s h t l g a i r t e en 00 0000 00 000000 00 t e a r) t i 00 0500 02 055005 0 0 1 9981 0 03 a mS s , , , , , , , , 33403 0 02 M i t l l e (p 1 1 4 436 60 741 1 1 3 5 4 998888 99 898888 00 e Ui s 1 1 c n n e a l l T i X e I v d r l h E u et 000000 00 000000 00 S i g)i 00 0 000 90 000505 5 4 L f Yn s 700570 80 321 577 68 B o % re (p A 01 5320 38 006491 91 s 2t 776666 76 776656 88 T e .S i t 0 r e p o . r t s p) mm0000 m0 mm0000 00 t P e mF oo0000 oo3336 o0 o3 oo0000 oo3366 05 35 l e Te* T( RR R RR i s n e n 1 l / o T ni A t a et e ii mca 0 V p S6 I 1 M I 56 cfi et pn 8 1 S SS d l 1 WW 2 e S e d I 2 1 W ) 3 I I e ) c ) 0 n) a ) a ) 6 ev ( d b ( ( d b ( 6 T l u e e 8 e 8 7 e FM i t a 1 0 t i a 1 0 Pu As l n1 d~ 1 i d ' 1f Cp a o> a x a x r r r WC t r 5 r 5 ue (E i n 0 i n 0 ) ) N U 6 U ab 6 ( ( e
? , i - l 33 V. ANALYSIS OF RESULTS The analysis of data obtained from surveillance program specimens has the following goals: (1) Istimate the period of time over which the properties of the ) vessel beltline materials will meet the fracture toughness requirements of i Appendix G of 10CFR50. This requires a projection of the measured reduc-
- tion in C yupper shelf energy to the vessel wall using knowledge of the energy -
.and spatial distribution of the neutron flux and the dependence of Cy upper l
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 RT NDT to the vessel wall using knowledge of the dependence
.of the shift in RTNDT on the neutron fluence and the energy and spacial distri- !
bution of the neutron flux. The energy and spacial distribution of the neutron flux for Turkey Point Unit No. 4 has been taken to be the same as thr.t for Turkey Point Unit No. 3 l and described in WCAP-8631. (14) The calculated lead factors for Capsule T ! are 2.48 for the vessel 1. D. surface, 4.17 for the vessel 1/4 T, and 17. 4 for
\
the vessel 3/4 T. ' A method for estimating the reduction in Cy upper shelf energy as a function of neutron fluence is given in Regulatory Guide 1. 99. (7) Howe ve r, it has been suggested $9) that a square root of fluence dependence fits existing l
r __ l i 34 1 (
, j i
data better. The results from Capsule T are compared to a portion of ' Figure 2 of Regulatory Guide 1. 99 in Figure 8. The Turkey Point Unit No. 4 weld metal data point lies just under the 0.25% Cu weld metal curve. Two supplemental curves have been added to provide a eneans for estimating the Charpy shelf reduction at lower fluences. The dotted curve through the data point is parallel to the Regulatory Guide 1. 99 curves (slope approxi-mately proportional to the fourth root of fluence). The dashed curve through the data point has a slope proportional to the square root of fluence. Since l i the initial shelf energy for the Turkey Point Unit No. 4 weld metal was 66.3 it-lbs, a shelf energy of 50 ft-lbs would be represented by a decrease in shelf energy of 25%. Therefore, the predicted fluences* for reducing the weld metal
- shelf energy to 50 ft-lbs are 1. 6 x 1018 and 3. 3 x 1018 for the Regulatory Guide
- 1. 99 and square root of fluence methods, respectively. Since the neutron flux is proportional to power generation, the neutron fluence and weld metal shelf
+
energy projections through five EFPY of operation of Turkey Point Unit No. 4 are presented in Table X. A similar approach can be taken to estimate the increase in RT NDT as a function of reactor power generation. Again, a method for estimating shifts in RT NDT is given in Regulatory Guide 1. 99. An alternate family of radiation damage curves developed by Westinghouse (14), shown in Figure 9, is more conservative than the Regulatory Guide 1. 99 curves up to a fluence
- All values of fluence cited in this section are in neutrons per em2 (E > 1 MeV).
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38-of nearly 1 x 10I9 Therefore, the projections of shift in RT NDT have been based on the Westinghouse curves. The result obtained from Capsule T has been added to Figure 9, and a normalized response curve has been drawn through the data point parallel to the Westinghouse curves. The predicted shifts in RT NDT for the Turkey Point Unit No. 4 reactor pressure vessel obtained from Figure 9 are summarized in Table XI. The values predicted at the 1/4 T and 3/4 T are used to develop heatup and cooldown limit curves to meet the requirements of Appendix G to Section III of the ASME Code. These projections for Cy shelf energy reductions and RTNDT shifts l
.are based on one data point, Capsule T, and trend curves for like materials.
i
.It is anticipated that the reliability of the trend curves will be improved as !
snore surveillance data becornes available and a better understanding of the factors affecting radiation embrittlement has been achieved. As an example of the latter, Mr. E. C. Biemiller of Combustion Engineering, in a paper given at the ASTM E10 Symposium on Effects of Radiation on Structural Materials in St. Louis, May 4-6, indicated that a parameter of (% Ni + % Si) l
+ (% Mo + % Cr + % Mn) may explain the variation in radiation embrittlement l observed in ferritic materials of nominally the same copper content. Also, the Metal Properties Council is developing new radiation damage curves that will be based on more data than those currently in use. !
l Because of the potential of reaching a low Cy shelf energy condition in (a the Turkey Point Unit No. 4 weld metal in the next few years, it is advisable s l 1 to obtain another data point in the not too distant future. I i! I (
, B J
T n u 4 7 2 $ T 4 5 6 1 Y R 4 3 P A F ) 9 9 Eb 1 3 8 1 2 0 0 0 3 (e 1 1 1 c n x x x e 2 3 3 l u 6 9 4 F 6 3 9 T n u 4 9 2 0 T 4 3 3 3 2 Y R 4 P A
. F 9 9 o E )
b 1 3 8 1 _ N 0 ( 0 0 0 t 1 e 1 1 1 i c x n n x x U e 7 3 5 t l u 0 2 9 n F . . . i 2 1 2 o P y T e D k N 5 1 8 r T 2 8 8 u 3 2 1 I T Y A R X E r P L o F ) 9 8 8 B f E b I 0 1 0 1 T 5 ( e 0 _ A D c 1 1 1 T N n x x x T e 3 5 7 u 0 R l 1 4 n F 1 6 1 i s t f T i D h N 2 2 2 S T 8 4 6 d 2 2 1 e I R t I A c Y e ) 8 j P b 1 8 1 7 o F ( 0 0 I 0 r E ec 1 1 1 V P e 3 n x x x M e 0 u 9 4 t 1 2 6 8 D> l F 6 3 8 W ME
. T T 0 D 4 4 0 2 I
. /
1
/
3 0, me l l l 3 e e e 0 r n s s s 8 ep o s s s = i t e e e s a c V V v Yn o o e r e r e r Pr Ft L u s u s u s Eue s s s 1 N e e e r r r ) ) P P P ab ( (
. t
40 i I The proposed Turkey Point Unit No. 4 reactor vessel surveillance . program schedule, summarized in Table X11, has been organized to satisfy Appendix H of 10CFR50 as closely as possible. There are seven additional
'l capsules in the vessel, - but only two of them, V and X, contain weld metal specimens. Capsules V and X are currently located in low flux positions (1/3 and 1/8 of that for Capsule T, respectively). The estimated fluence h
for Capsulev is 4.15 x 1018 at the end of Core Cycle III (April 1977) and --
- 5. 8 x 1018 at the end of Core Cycle IV (April 1978).
There are two ways to achieve higher fluences in a shorter time period. If archival weld metal is 'available, a replacement capsule should be prepared for insertion into the Capsule T alot during the 1977 refuelling outage. Such a capsule would receive a fluence of about 4 x 1018- pe r year at an 80% load factor. As an alternative, FP&L should investigate the pos-sibility of moving Capsule V to the T position during the 1977 refuelling outage. If this can be done, it is estimated that Capsule V will have received a fluence of 8. 25 x 1018 (4.15 x 1018 through Core Cycle III in the V position plus 4.10 x 1018 during Core Cycle IV in the T position) by the 1978 refuelling outage. ! In the event that the latter course is chosen, it is still advisable to install a replacement weld metal capsule if archival material is available to provide additional irradiated material for fracture toughness and annealing studies should they be required. Such a capsule could be installed in the 3 slot i which is scheduled to become available during the 1978 refuelling outage.
. 3 .e
4 41 TABLE XH Proposed Reactor Vessel Surveillance Capsule Schedule Turkey Point Unit No. 4 Approx. Predicte d Capsule I.D. Lead 1/4T Lead Caps ule Removal Capsule Fluence Ident. Factor Factor Type (") Date(b) (n/cm 2, E > 1 MeV) T 2.48 4.17 H Tes ted 6. 05 x 1018 S 1.60 3.60 I Apr.1978 1.3 x 10(19)I'I V 0.79 1.50 H (c) - W 0.34 0.61 1 Apr.1983 5. 6 x 1018(e) Y 0.49 0.88 I Apr. 2 003 2. 4 x 1019(*) U 0.49 0.88 1 (d) - X 0.34 0.61 II Apr.1993 1.1 x 10 19I* I Z 0.34 0.61 I (d) -
, (a) Type H contain weld metal specimens.(13)
(b) At time of refuelling outage. (c) Move Capsule V into Capsule T alot during 1977 refuelling outage, if feasible. Remove for testing during 1978 refuelling outage. (d) Spare capsule. - (e) Based on 80% load factor. ! l l l
42 VI. . 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 E208-69, " Standard Method for Conducting Drop-Weight Test to Determine Nil-Ductility Transition Temperature of Fer-zitic 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 Pres-sure Vessel Steels," International Atomic Energy Agency, Technical l Reports Series No. 163, 1975.
1
- 6. ASME Boiler and Pressure vessel Code, Section XI, " Rules for In- I seriice Inspection of Nuclear Power Plant Components," 1974 Idition.
- 7. Regulatory Guide 1. 99, Office of Standards Development, U. S.
Nuclear Regulatory Commission, July 1975.
- 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 E185-73, " Standard Recommended Practice for Surveillance Tests for Nuclear Reactor Vessels," 1975 Annual Book of ASTM i Standar ds .
- 11. ASTM E399-74, " Standard Method of Test for Plane-Strain Frac-ture Toughness of Metallic Materials," 1975 Annual Book of ASTM
.Standa r ds .
t 43 ; 1 i
- 12. Witt, F. J. , and Mage r, T. R. , "A Procedure for Determining Bounding Values of Fracture Toughness Kyc at Any Temperature,"
ORNL-TM-3894, October 1972.
- 13. Yanichko, S. E. , " Florida Power & Light Company Turkey Point Unit No. 4 Reactor Vessel Radiation Surveillance Program, "
WCAP 7660, May 1971.
- 14. Yanichko, S. E. , Phillips, J. H. , Anders on, S. L. , " Analysis
' of Capsule T from the Florida Power & Light Company Turkey Point Unit No. 3 Reactor Vessel Radiation Surveillance Program,"
j- WCAP 8631, December 1975. l )
- 15. Helm, J. W. , "High Temperature Graphite Irradiations: 800 to 12.00 *C: Interim Report No.1," BNWL-112, September 1965.
l I l 1 1 ! l l L ) l f l l l l l t l
. * ~
. t
r t f e a O APPENDIX A INSTRUMENTED CHARPY V-NOTCH IMPACT TESTS b e l l l
*
- w . o
A-1 INSTRUMENTED CHARPY V-NOTCH IMPACT TESTS The Charpy V-notch tests were carried out on a SATEC impact tester equipped with an instrumented tup and a sine-cosine transducer. The cali-bration of the tup signal was carried out statically in a universal testing machine. A dual-beam oscilloscope and a two-channel oscillographic recording system was employed to obtain the following information from each test: (1) The force-time history (2) The energy-time history The energy-time history was obtained by integrating the force signal with respect to velocity. 4 The force-time and energy-time records were used to obtain the fol-lowing information: (1) Load at general yielding, P y (2); Maxirnum load, P max (3) Load at onset of fast fracture, Pg l (4) Energy prior to crack initiation, E i i (5) Energy prior to onset of fast fracture, E f (6) Energy of shear lip formation, E, The methods for obtaining the above information are illustrated in Figures A-1, A-2 and A-3. Figure A-1 illustrates a case where unstable crack propagation initiates shortly after general yielding in the lower portion e9 0 0 t p
't 0,s
>b
_ i k, _ I _ O
~
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4 4 . T
. S E
t . T N E O I - G E R N O I T I S N k I A .
- R _
T h R t - b 5 , E . W sl a '5 9 3 O L p3m2 1 4 cE; - 1 I i A b E l R o U
.o y G PoB .
NE g1 L I F s et i t _ i' i 1 1 2 2 4 7-F ! 2 o 0 S i
+
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+
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- ' i I l '
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e I b l O x o 3 1 s t E - T S E T F L E H S R E P P t U 3 A E R U G I F i f o h t E 4 o
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t 0 s 2
+
A.$ed
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t 8 e
l A-5 of the energy transition region. Figure A-2 represents the behavior in the upper portion of the transition region, and Figure A-3 shows a typical upper shelf response. In the latter case, the fracture propagates immediately in shear as evidenced by the lack of a sudden drop in load caused by unstable crack propagation. The instrumented Cy data obtained are summarized in Tables A-1 through A-4 for the shell forging, weld metal, heat affected zone, and reference materials, respectively. 4
~
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. Lit t f 0. 5 5 0 r a(
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- no a vin g r 1 yrt C
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d L Cen a hi o B dSo eL A e t r P l ) 5 0 5 0 0 5 T ney t t es 2 0 2 5 0 7 i rb 8 - ewe mok r uI B tc ( 3 3 4 3 4 1 4 0 3 8 2 uLr r u a t T r t s F I n m m o f r u ma d )s 5 2 0 0 5 2 0 0 0 5 5 2 0 0 s iob I 8 3 - 3 2 1 0 8 t xL( a 3 4 4 4 4 4 3 l u M _ s e R d l a _ ao r L )s e 5 7 5 7 5 0 0 0 0 5 0 0 0 5 0 5 ndI b 3 3 1 0 9 9 7 4 el ( 3 3 3 3 2 2 2 2 Gi e Y _ p _
)
0 0 0 0 mF e* 5 2 1 1 5 2 0 4 1 8 0 1 T( 2 _ c . 0 1 7 2 8 3 4 9 eo 7 7 6 7 6 7 7 6 pN - - - - - - - S S S S S S S - S S
. m.
) >L s
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)
s e ol-nb l i t t b a 0 5 0 t a g (f . . . s ay 9 4 4 - - - - - s npo g 1 1 t s Ur re e Pn T E t c a p n i o m t I i ay) h t ge 0 5 0 5 0 0 0 c irb n I t e - 9 4 4 - 7 6 7 6 o 4 I nt 1 1 1 1 1 1 N . kc E (f
- o a
V N r 2 l
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C A n reU E haM t d L Cdl ni a B _o A d e o eL ) 5 0 T teW Py n l t t es 7 0 5 7 irb 7 8 8 - - - - - e e ruI ( 3 3 3 k Bt c m r u u a r r t s T F I n m m o ud) r s 5 0 5 5 0 5 0, f ma iob 7 7 0 8 7 8 - 2 0 2 5 s l 6 8 5 3 t l xL( a 3 3 3 3 3 3 3 u M s e R d l a ao re L )s b 0 5 0 0 5 2 0 0 5 2 0 0 0 5 ndI 1 8 9 - 8 9 6 6 el ( 3 2 2 2 2 2 2 Gi e Y - p) 1 0 0 0 0 0 5 5 m*F e 8 1 6 1 0 0 7 7 1 2 T( 1 3 3 3 4 c . 4 3 2 1 0 7 8 9 eo 2 2 2 2 2 1 1 1 pN - - - - - - - - S V V V V V V T ) ) W T T T W -
4
)
>5 .
s pnb i Lit o-tl 0 0 O 5 5 0 0 0 r a (f . . . amy 2 8 9 1 6 8 6 5 erg 1 9 9 h or SFe n E
)
s nb l e ol-i t b t 0 0 0 a ga (f 0 0 0 t s ay 4 1 0 8 5 2 - - s npo g 3 3 3 5 9 t s Ur re e Pn T E t c a n p o i m t I i a yg )s hl t 0 0 0 0 0 ca i rb 0 0 0 In e - l . . . . . . t i4 4 1 0 8 0 8 5 o re nt 3 3 3 4 8 Nt . 3 4 4 o kc E (f VM aN a r t 3 y i C
- p e n n A r oU E h aZt n d L Cdi a eo o B dt P eL A ec l t
es
) 0 0
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)
>E s
pnb _ i ol - _ Lit ft 5 O 5 5 5 O 0 0 r a( amy 2 3 4 4 4 5 5 7 er g 1 5 5 h or SFe n E
)
s nb ol le it - b t 0 0 a ga (f 5 0 0 0 t s ay 2 5 7 8 6 2 - - s npo g 1 1 4 t s Ur re e Pn T E t c a n p o i m t I ay) h i t gs 0 0 5 c i rbl 0 0 0 0 0 t o 4 I n ent - 2 5 7 8 6 8 7 5 1 Nl-ao kc E (f 1 3 3 3 4 VriNet a r yt C i
- pa n A r U E aM t d -
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)
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. i
SOUTHWEST RESEARCH IN S TIT U T E Post Office Drower 28510, 8500 Culebro Rood- )
. San. Antonio, Te xa s 7828 4 kPRESSURE-TEMPERATURE LIMITATIONS . l
.FO R THE )
. T.URKEY POINT UNIT NOS. 3 AND.4 i
NUCLEAR POWER PLANTS 1 bY E. B. Norris J. F. Unruh SwRI Project No. 02-4383 039 to Florida Power and Light Company Miami, Florida June 301976 Approved: 7 . U. S. Lindholm, Director Department of Materials Sciences 4 4 9
?
TABLE OF CONTENTS Page ' l
' LIST OF FIGURES iii-I. PROCEDURE FOR THE GENERATION OF ALLOW-ABLE PRESSURE-TEMPERATURE LIMIT CURVES FOR NUCLEAR POWER PLANT REACTOR VESSEI.S 1 A. Introduction 1
- 15. Background. 1 C. Pressure-Temperature Relationships 2 l 1.- Inservice' Leak and Hydrostatic Test 2'
- 2. Heatup and Cooldown Operations 6
- 3. Core Operation 10 D. Thermal Stress Analysis 10 E. Example Calculations 16 IL HEATUP AND COOLDOWN LIMIT CURVES FOR e NORMA 1 OPERATION OF TURKEY PO)NT UNIT NOS. 3 AND 4 25 A. Introduction 25 B. Input Information 26
- 1. Fracture Toughness Properties 26
- 2. Ves sel Constants 27 C. Heatup and Cooldown Limit Curves 27 j i
1
'I ii ,
l l i
'r , j
__g s
I l LIST OF FIGURES Ficure Page 1 Stress Correction Factor 3 2 Reference Stress Intensity Factor 4
.3 Heatup and Cooldown Stress. Distribution 7
- 4. Typical Normalized Temperature and Stress Distribution During Heatup 17 5 Reference Stress Intensity Factor as a Function
, of Temperature Indexed to RTNDT (1/4T) 19 6 Vessel Temperature at 1/4T and 3/4T Locations as a Function of Coolant Temperature 20 7 Thermal Stress Intensity Factor at 3/4T and 1/4T Locations as a Function of Coolant Temperature 21 8 Pressure-Temperature Curves for 100*F/Hr Heatup 23 9 Pressure-Temperature Curves for 2 00 *F/Hr Cooldown - 24 10 Turkey Point Unit 3 Reactor Coolant Heatup Limitations Applicable for Periods up to 5 Effective Full Power Years 29 11 Turkey Point Unit 3 Reactor Coolant Cooldown Limitations Applicable for Periods up to 5 Effective Full Power Years 30 12 Turkey Point Unit 3 Reactor Coolant Heatup Limitations Applicable for Periods from 5 to 10 Effective Full Power Years 31 ,
i 13 Turkey Point Unit 3 Reactor Coolant Cooldown Limitations Applicable for Periods from 5 to 10 Effective Full Power Years 32 iii
,l
)
1
. . s l
LIST OF FIGURES Cont'd. Figure Page 14 Turkey Point Unit 4 Reactor Coolant Heatup ( Limitations Applicable for Periods up to 5 ; Effective Full Power Years 33 15 Turkey Point Unit 4 Reactor Coolant Cooldown ' Limitations Applicable for Periods up to 5 Effective Full Power Years 34 16 Turkey Point Unit 4 Reactor Coolant Heatup Limitations Applicable for Periods from 5 to
- 10. Effective Full Power Years 35 17 Turkey Point Unit 4 Reactor Coolant Cooldown Limitations Applicable for Periods from 5 to
.10 Effective Full Power Years 36 e
1 IV t j
1 I. PROCEDURE FOR THE GENERATION OF ALLOWABLE l PRESSilRE-TEMPMRATllRE LIMIT CtIRVES FOR 'l NilC LEAR l'OW MR Pl. ANT REACTOR VMSSMIS ? r A. Introduetion -
.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 llI, Appendix G, " Protection Against Nonductile m.
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 KI corresponding to membrane tension is given by KIm = M m *F rn (1) where M m is the membrane stress correction factor for the postulated
' flaw and e m the membrane stress. Likewise, K Icorresponding to bend-ing is given by 6
KIb = Mb * *b (2) where Mb is the bending stress correction factor and ab is the bending
^
stress. For vessel section thickness of 4 to 12 inches, the maximum , i i 4
)
. . r
,l '
\ .. i 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 c
J 1ength 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 I as taken from the Pressure. Vessel Code (ref. Figure G-2114.1). These curves are a function of the stress ratio parameter r/r y, .where e 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 I 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 Powe r Plant Components, " Article A-3000. The Code specifies the minimum Ky that can cause failure as a func-tion of material temperature, T, and its reference nil ductility temperature, RT NDT. This minimum Ky is defined as the reference stress intensity fac-- tor, K Ig, and is given by KIR = 26777. + 1223. exp
- 0. 014493(T - RTNDT + 160) (3 ) -
where all temperatures are in degrees Fahrenheit. A plot of this expre ssion. is given in Figure 2 taken from the Code (ref. Figure G-2010.1). C. Pressure-Temperature Relationships f
- 1. . Inservice Ieak and Hydrostatic Test '-
During performance of inservice leak and hydrostatic tests, the reference stress intensity factor, KIR, must always be greater than l 1 e j
l 3 3.0 l l l { l l #/ry ' 3.0 MEMDRAHEg K ,= M ,x r , - g,o BEN 0!NG Kib ' Mb * 'b / , p N a2/3M m b
/ ,
o'f 3.0 ! 2.0 E ; si 2A /'/ - i
~
/
/ .
3 2.2 7 ~ 2.0 - -- b_.J
-- ^
I '
. l.B =
, I l.G l.4 l.2 1.0 ID 1.2 1.4 1.6 1.0 2.0 2.2.2.4 2.G 2.B 3.0 3.2 3.4 3.6 3.8 4.0 YTHICKNESS ( LN.)
I 1
)
FIGURE 1. STRESS CORRECTION FACTOR
1
. l- , !
.q .- ' f 4 i . . j I- 17 0 , 10 0 -
; 15 0 - (g m -2G.777) a l.223 eo oi449dT-Int,,,-Icod
,j 140 V.'HERE 130 - K ig
- REFEREllCE STRESG INTEllsiTY FACTOR I
12 0 - - - - - - --- T = TEMPERATURE AT V/HICH K,7, . IS PERMITTE(),'F 11 0 -- RTc7 = REFEREr:CE I.'tL-0VC11LITY !
, _W 100 -- --
' TEl.tFERATURE !
,_ 90 - - _ . .
- /'
.dw
- g c CO E
.x 70 60 ----. - - - - - - -.. .. . _
c $0 - 40 - - - - - .--- 30 " 7
~~
20 to -- o; ' f , I 8 I !
, - 24 0 -200 -100 -12 0 -80 -40 0 40 80 .12 0 100 200 240. .,
TEMPERATURE RELATIVE TO Rig p7,(T- ATuoT), FAHRENHEIT DEGREES i m ew 9 1 FIGURE 2. REFERENCE STRESS INTENSIT'Y FACTOR f
4 1
]
5 l i 1.5 times the Ky caused by pressure, thus ' l l 1.5 Kyp < KIR (4) or
- 1. 5 M m Fm"KIR . .(5)
For a cylinder with inner radius ri and outer radius ro, the stress distribution due to internal pressure is given by r o2 + r2 r(r)=[\ r ozri- riZ r2 * (6) With 1/4T flaws possible at both inner and outer radial locations, i.e. , at ryf4 = ri + 1/4(ro - ri) and r3 /4 = ri + 3/4(ro - ri), the maximum stress will occur at the inner flaw location, thus 4 2 ri r o2 + (1/4ro + 3 /4ri)2 i
' max
- P o r o2 - ri2 (1/4ro + 3 /4ri)2 j
' ~
With the operation pressure known, i.e. , Po , we deter- l mine the minimum coolant temperature that will satisfy Equation (4) by I evaluating KIR = 1.5 Mm ' max (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 iny26777.- , g
-j 6 )
e !' The inservice curves are generated for an operating pres- ! mure range of .96 P o to J.14 Po , where P ois the design operating pre s s ur e.
- 2. Heatup and Cooldown Operations At all times during heatup and cooldown operations, the ref-erence stress intensity factor, Kg, must always be greater than the sum of 2 times the Kyp caused by pressure and the kit caused by thermal gra-dients, thus
- 2. O Kyp + 1. O kit < Em (10) l or
- 2. O Mm Fmax
- Km - kit (11) where ' max is the maximum allowable stress due to internal pressure, and Ki g 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. l 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 Iccation, we see from Figure 3a 1 that the thermal stress tends to alleviate the pressure stress at this point in the vessel wall and, therefore, the steady state pres sure stre ss l would represent the maximum stress condition.at the 1/4T location. At l
, 1
.l
i
- 7 I i I
OUTER RADIUS - l 1 l 3 3/4T i' 3 3 __ 1/4T ,
-lNNER RADIUS Pressure stress distribution Thermal stress distribution i (a) Heatup '
}
OUTER RADIUS 3 3/4T 3
} __
1/4T 3 3 INNER RADIUS Pressure stress distribution Thermal stress distribution (b) Cooldown ! i Figure 3. Heatup and Cooldown Stress Distribution r >
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 KIR P***(1/4T ) = (12) [ ri 3 fr,2 + (1/4ro + 3 /4r i)2) 2 i m r o2 - rt2 (1/4ro + 3 /4ri)2
)
where M m is determined from the curves in Figure 1 and K IR is obtained from Equation (3) using the coolant temperature and RTNDT at the 1/4T location. ' Here we may note that M m must be iterated for since it is a
= functicn of the final stress ratio to yield strength (r/ry).
At the 3/4T location, the maximum pressure is determined from Equation (11) as - KIR - kit i Pmax(3/4T) = 2 (13) [ ri )[r,2 + (1/4ri + 3 /4ro )h m o r Z - ri2j (1/4ri + 3 /4ro)2 ) j where KIR,is obtained from Equation (2) using the material temperature and RTNDT at the 3/4T location and Kyg is determined from the analysis procedure outlined in Section D. M m is determined from Figure 1. q= 6 6
9 The minimum of these maximum allowable pressures at ; the given coolant temperature determines the maximum operation ' l 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
'Kg-kit Pmax(1/4T) =
[ ri 2
) fr o2 +(3/4ri+ 1/4r o)2)
(14) m o 2-rij( 2 r (3/4ri+ 1/4ro) ; where K IR is obtained from Equation (3) using the material temperature and RTNDT at the 1/4T location. K3 e 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 pres sure. When the thermal gradients increase, the stresses do likewise, and, finally, the transient analysis governs the maximum allowable pressure. Hence a point-by-point
10 t-comparison must be made between the maximum allowable pressures pro-q 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 re sults.
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 pcTg-K (1/r) Tr+T rr =0 (15)
, subject to initial condition T(r, 0) = T o, (16) l l and boundary conditions
-KTr(ri, t) = h(Te(t) - T(ri, t) _
(17) 4 B
- n. . _ _ .
11 and T r (r o, t) = 0 (18) whe re - Te=To + Rt. (19) p is the material density, e 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 or apart, across the vessel are: for I < n < N t r at K - Tn+At * [,1 - pc(ar)2 ( 2 +{Ar ) T _n At K - e
+ ( 1 + Ar ) Tn+1+Tn-1 (20) pe(Ar)2 { ,
for n = 1 6t X Ath Tt3+6t = 1 pc(or)2 ( 1 + ar )- pc(Ar) .TI t r1
+
atK - or t-pc(Ar)Z _( +-)T 2+Arh g T e, , (21)
12 and for n = N ! t
. At X -
g Kat ! T N+ At
- I~
_ oc(Ar)2 , TN + pc(Ar)2 T N.1 . (22) For stability in the finite difference operation, we must choose At for a given Ar such that both At K pc(Ar)2 (2 + Ar ) $ 1 (23) j and ! l At K At h oc(Ar}2 (3 r] < I
+~Ar ) + pc(Ar) l (24) are satisfied.
These conditions assure us that heat will not flow in the d'irection of increasing temperature, which, of course, would violate the second law of thermodynamics.
.Since a large variation in coolant temperature is considered, the i
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. I The dependence of (K/pe) 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: t i! K(T) = 38. 211 - 0. 01673
- T (BTU /HR-FT *F) (25) {
1 O O
13 and k(T) = (K/p c) = 0. 6942 - O. 000432
- T (FT2 /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 , i 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 + 10-2
- T 5
- T2 ) yo. 8 /DO.2 (27) where v is in it/sec, D in inches, the temperature is in *F, and h is in Btu /hr-ft2 . .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 (ft2). 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 = Q/(3600
- p(T)
- A) .
(29) Glasstone, S. , Principles of Nuclear Reactor Engineering, D. Van Nostrand Co. , Inc. , New Jerse y, pp. 667-668, 1960. 4 I l
4 14 l The thermal stress distribution is calculated from cE - 1 r2 + ri2 (r o T(r, p, 1 ' rT(r, t) = g T(r, t) rdr'- T(r, t)+g (r0 2. q2 )J t) rdr (30) rt ri 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 a and E are evaluated as a function of the average ternperature across the vessel 2 r# 0 T ,yg = (r 2-ri)J 2 M*- M 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. 9342 + 2. 5782 x 10-4
- T - 6. 5723 x 10-6
- T2. (33) l 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 I linear; however, an equivalent linear stress distribution is determined from the resulting moment. The moment produced by the nonlinear I a j T 1 1
[' .__ . 15 stress distribution is given by
.ro i M(t) = b. 'T (r, t) rdr (34) ri
, where b is a unit depth of the vessel. Here we note that the moment is a i
Junction of time, i. e. , coolant temperature via Te = To + Rt. For a lin-ear stress distribution we have that Mc
' max
- I (35) where r max is the maximum outer fiber stress, e the distance from the neutral axis, taken to be (ro - ri)/2, and I the section area moment of inertia which is given by y , bh3 ,b(re - ri)3
- 12' 12 (36)
Combining these expressions results in the equivalent linear stress due
- to thermal gradients 6 '*o ,
' max * 'bt * (ro - ri)Z .ri "T (r, t) rdr (37) j The thermal stress intensity factor kit is then defined as Kyg = Mb 'bt (38) l l
where Mb is determined from the curves given in Figure 1 wherem i Mb = 2/3 M m. It is of interest to note that a sign change occurs in the l i stress calculations during a cooldown analysis since the thermal gradients l
'l 16 produce compressive stresses at the vessel outer radius. This sign change must then be reflected in the Kyg 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, T avg, by
~
T=g avg) 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 1/4T location becoming the critical point. E. Example Calculations The following example is based on a reactor vessel with the follow-ing characteristics: Inn'er Radius = 82. 00 in. (ri) Outer Radius = 90. 0 0 in. (ro) Operating Pressure = 2250 psig (Po) i i
17
-OUTER WALL l 1.0 -
f
- /
0.8 - 0.6 -/ r 0.4 I -
- 0. 2 0 /' ' '
- 1. 0 0 1.0 - 1. 0 0 1.0 lNNER WALL Normalized temperature Normalized stress
. distribution ( AT/ ATmax ) distribution ( al amax I i
Figure 4. Typical Normalized Temperature and Stress Distribution During Heatup l r i
18 Initial Temperature = 70*F (To) Final Temperature = 550'F (Tf) Effective Coolant Flow Rate = 100 x 106 Lbm/br (Q) Effective Flow Area = 20. 00 ft2 (A) Effective Hydraulic Diameter = 10. 00 in. (D). RTNDT (1/4T) = 200*F 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, Km, as a
. function of temperature indexed to RTNDT (1/4T). For the steady state analysis, Km 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 100*F/hr heatup and cooldown analyses are given in Figure 6. These temperatures allow computation of the corresponding reference stress intensity factors, Km (3/4T) and Kg (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.
- . . ~
ii 1 0 )
- - - 0 T 4
4 I I ( T _ D T N i R 0 o _ 8 5 t 3 - d e x e d I n . e r 0 t u
- 0 a 3 r e
p m e T f. o
)
0 F n 5* i o t 2 ( c E n R u F
. U a T
A s R a
. E r 0 P o 0
2 M E t c a T F y F t i s 0 n . 0 e - 2 t
- i n -
) 0 5 s T 1 s 4 e
/
1 t r ( S T e D c N n T e r -
, R e
. 0 f 0 e 1
R 5 _ e r _ u g
~ - -
i
- 0 F 0
0 2 l 0 2 1 0 8 0 4 0 5
.E. ;> 3 _ $
e - '
~ - - - 3 e
r t u
/ a r
e
' p I
0 m s ' 5 3 T e - t
/
n a
/,,
l o o C
, f I
o
/ '
0 n 0 o 3 i
' t c
n
/ u F
/ a s
a I
/ '
5 0 ) s
)
, 2 F* n o
n i . ( t -
) o a i
t / E c - n a R L o _ o c U i t a L o ' T A T 4 c / 4 I l o T 4
/
s ' 0 E 0 P R 3 d - T 1 2 n _ 4
/
( / M E a 3 N T T ( W / T / 4 P O N 1 U D / A t T L L a A O O e I E H O C
/ 0 O r u
i 5 C 1 t a R H R H
/ r e
/ / p F
F
/ m e
0 0 T 0 0 /
/
l 1 1 e I s 0 s
- / i 0
1 e - V
- / 6 e
r - u g -
~ -
i F - IY 2 0 0 0 0 1 O5 0
' i u_* " h < E ' 5 n > U W > -
e r t u a ~
- - - _ - r e
p m e
/ T t
n
' a 8
0 l 5 o 3 o C
- ~
)
f o
- n n
)
o o n i i t c i o t a n t c 0 . a o i 0 u W c L Lo 4 / 3 F a p T 1 ( s a 4
/
s
/ 3 N
( W n o
' P U
T L A D O 0 i 0 5
)
2 F* i t L a c o p R E H C R 0 . ( E 4 I T I H H R I / U d F F T n A a
, 0 0 R T 0 0 ' 0 E 4 1 1 0 P 2
1 3
- M t
- E a T
T N A t r o c a L F O y
' 0 O i t
5 C 1 s n
' t I
e n s s e 0 0 t S r l 1 a m r _ e _ T h 7
- - _ - 0 2 0 8 e 1 1 6 4 2 0 5 r u
g i F _ E. P x s e
22 Figures 8 and 9 demonstrate the construction of the allowable com-posite pressure and temperature curves for the 200*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.
l l l
- _ - ~ - - _
Y U _ T i l 0 - l I 0 - A 4 _ C _ I T - TI _- M I - R Cl l i 0 I 5 g l 3 p t u l
/ I a
e
/ H T
/ H r -
I / / i M / 0
- F I I 0
L T
/ 3 0 0 _
S E / ) 1 r o T K
)
r
/
- F f s
A r o ( e v E r E r L e i
' 0 R 5 UT C u
S I t n 2 e A e E r m T R E u Pu A P t a Ur t T r T s S M e An E p Ei Y T m Hr D e i o A 0 D T Rf E I 0 E Hg T 2 T - i S A e I s C r F p I u 0 D s s 0 0 1
~ n 6-d
N I P e r i E a 0 . VF , I 5 8
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U *0 1 e r
- u C+ g f - i E o P F T
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~
i /
/ t 0
5
/ 3
/ -
f[ / n w o d i
/ l o
o
/ e 0
0 3 C
/ H r
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/
F 0 0 1
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) r W i 5 F f o
O 2 *( s D e L E v r O O C
)
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T A C u e r R r R i H o E t u
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- e 2 ME e 0 t T p 0
1 d e n m
/, D e
~n a m E T
T E r u N A e V *F t W C r i R0 s O I u U1 n D 0 D s C+i r L e 5 N 1 s e f 0 I r E T ofo 0 P I sg C S ni E . Oisgp T A
- 9 D r Ma0 T
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0 0 2 0 6 1 0 2 0 8 0 g - 1 ~
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25' II. HEATUP AND COOLDOWN LIMIT CURVES FOR NORMAL OPERATION OF TU'RKEY POINT UNIT NOS. 3 AND 4 A. Introduction Turkey Point Unit Nos. 3 and 4 are twin 2200 Mw e pressurized
' water reactors located south of Miami, Florida.
Each unit has been provided with a reactor vessel material surveillance program as re-quired by 10CFR50, Appendix H. The first surveillance capsule (Capsule T) was/ removed from Unit No. 3 during the 1974 refuelling outage. This capsule was evalu-ated by Westinghouse, and the results have been reported.
- In summary, this analysis indicated that:
(1) Of the beltline materials in Unit No. 3, the weld metal is the limiting material and should control the RT NDT over the life of the plant. (2) The RT NDT of the weld metal in Capsule T in-creased 190*F as a result of exposure to a nen-tron fluence of 5.68 x 1018 neutrons /cm2 (E > ' 1 MeV). (3) The RT NDT after lo effective full power years ! of operation was predicted to be 170'F and 112 *F at the 1/4T and 3/4T vessel wall locations, respectively. Yanichko, S. E. , Phillips, J. H. , and Anders on, S. L. , " Analysis of Capsule T from the Turkey Point Unit No. 3 Reactor Vessel Radiation Surveillance Program, " WCAP-8631, December 1975. J
n ., m - nl;j _
.j s
(; ' 26 ! 1
- The first surveillance capsule was removed from Unit No. 4 du. ring - i
1 MeV). A summary of these values is as follows:
Unit Operating RTNDT ' RTNDT No. Period
- at I/4T at 3/4T 4
3 5EFPY 194 *F 131*F-3 10 EFPY 236'F. 159'F 4 5 EFPY 281*F 188'F 4 10 EFPY 342*F 230*F
- EFPY = Effective Full Power Year
- 2. Vessel Constants The following input data were employed in this analysis:
Inner. Radius, ri = 77. 75 in. Outer Radius, r o = 85. 78 in. Operating Pressure, Po = 2235 psig Initial Temperature, To = 70*F Final Temperature Tf = 550*F Effective Coolant Flow Rate, O = 97 x 106 lbm/hr 3 Effective Flow Area, A = 19.15 ft2 Effective Hydraulic Diameter, D = 11. 9 in. C. Heatup and Cooldown Limit Curves Heatup curves were computed for a heatup rate of 100 *F/hr. Since lower rates tend to raise the curve in the central region (see Figure 8), these curves apply to all h%ating rates up to 100'F/hr. Cooldown curves were computed for cooldown rates of O'F/hr (steady state), 20'F/hr, 1 t
28 60 *F/hr and 100 *F/hr. The 20'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 2 00*F/hr curve would apply.to rates from 60*F/hr to 200 *F/hr. -
. The Unit No. 3 heatup and cooldown curves for up to 5 EFPY are given in Figures 10 and 11. Unit No. 3 curves covering 5 to 10 EFPY are given in Figures 12 and 13. Corresponding curves for Unit No. 4 are given in Figures 14 through 17. -
4 1 h. l l
/
4
I _. 0 eIu,n ;= L ., _.;
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- u. . u...::n ._u.. =_:
5
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2"n := = : r dn E.
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