ML20155G439
| ML20155G439 | |
| Person / Time | |
|---|---|
| Site: | FitzPatrick  |
| Issue date: | 04/25/1986 |
| From: | Branlund B, Caine T, Ranganath S GENERAL ELECTRIC CO. |
| To: | |
| Shared Package | |
| ML20155G436 | List: |
| References | |
| DRF-B11-00339, DRF-B11-339, MDE-49-0386, MDE-49-386, NUDOCS 8605060131 | |
| Download: ML20155G439 (112) | |
Text
{{#Wiki_filter:a +b-a ,+&+ 1 m Q MDE-49-0386 O DRF Bil-00339 Class I April 1986 O JAMES A. FITZPATRICK NUCLEAR POWER PLANT O- - REACTOR PRESSURE VESSEL SURVEILLANCE MATERIALS TESTING AND FRACTURE TOUGHNESS ANALYSIS O 1 1 O Prepared by: I b. C 4//25/1e T. A. Caine, Senior Engineer Structural Analysis Services =O Verified by G . % g di_ YZ5/S(e O B. J. Branlund, Engineer Structural Analysis Services Approved by: O S. Ranganath, Manager Structural Analysis Services O lO 8885186818I88l}g3 P
O-
- n. IMPORTANT NOTICE REGARDING v
CONTENTS OF THIS REPORT PLEASE REAL CAREFULLY This report was prepared by General Electric solely for the use of the
1.0 MeV) measured was 1.4x10' n/cm2-sec. Based on the flux wire data, the O surveillance specimens had received a best estimate fluence of 17 2.6x10 n/cm2 at removal.
l j
- e. The vessel peak inner surface and 1/4 T lead factors were lO 4 established analytically using a reference analysis that combines i two-dimensional and one-dimensional finite element computer t
analysis. The flux peak occurs at an azimuthal location 45' to either side of the vessel quadrant references. The lead factors for the surveillance capsule are 0.79 to the peak vessel surf ace and 1.05 to the peak 1/4 T depth location.
- O l
l 2-2 !O l L
g a
- f. The maximum accumulated neutron fluence at the assumed vessel O end-of-life (EOL) of 32 EFPY was determined at the peak 1/4 T location. The maximum 1/4 T vessel EOL fluence is 18 I 1.3x10 n/cm2 (best estimate) and 1.7x10 n/cm2 (upper bound).
o' ' g. The surveillance Charpy V-Notch specimens were impact tested at temperatures selected to define the transition of the fracture toughness curves of the plate, weld, and HAZ materials. Measurements were taken of absorbed energy, lateral expansion and
- n
percentage shear. Fracture surface photographs of each specimen are presented in Appendix A. From absorbed energy and lateral expansion results for the plate and weld materials the following values are extracted: index temperatures for 30 ft-lb, 50 ft-lb, O and 35-mil lateral expansion (MLE) values and USE. The HAZ material data were not curve fit because of the significant scatter in the test results. O h. The irradiated plate impact energy curve is compared to unirradiated data from earlier studies to establish the RT NDT irradiation shift for the surveillance program. The plate material shows a 23*F shift. Decrease in USE was estimated for O the plate material also, showing a 6% decrease.
- i. Rockwell C hardness tests were done on one broken half of each Charpy specimen. The average hardness for the base metal was 0 13.0 HRC. The weld metal specimens had an average hardness of 18.5 HRC. The average for the HAZ material was 18.8 HRC, indicating'that the tests were done on the weld side of the HAZ specimens.
O
- j. The irradiated tensile specimens were tested at room temperature (76*F), reactor operating temperature (550*F), and estimated onset to upper shelf temperature (185'F). The results tabulated n
" for each specimen include yield and ultimate tensile strength uniform and total elongation, and reduction of area. The plate, weld, and HAZ specimens behave similarly for most properties. O- 2-3
.O L
- k. The irradiated plate and weld material tensile test results are compared to unirradiated data from the vessel fabrication test program. The materials show increased strength and decreased
[ ductility, as expected for irradiation embrittlement. The weld i material shows a greater change in properties, which is O consistent with its higher copper content relative to the plate. 1
- 1. As a part of the construction of the updated operating limits curves, the plate metal irradiation shift in RT was compared
} NDT to predictions calculated with Regulatory Guide 1.99, Revision 1 (Reference 5). The surveillance test shift of 23*F in plate material RT ET f r a fluence of 3.25x10 n/cm (upper bound) is
- slightly greater than the predicted shift of 19'F. Accounting for the small underprediction of plate material shift, the EOL adjusted reference temperature (ART = initial RT g plus irradiation shift) of the plate is 86*F. The predicted ART at EOL for the weld metal is 104*F, so the weld is the limiting g; -
beltline material. However, the plate has a higher initial RT 'DT than the weld, so for the first 16 EFPY the plate is limiting. The larger irradiation shift of the weld, with its higher copper content, causes the weld ART to " overtake" the plate ART at about 17 EFPY. This is shown graphically in Figure 7-4 of the report.
- m. The USE at EOL is predicted using the methods in Reference 5.
The weld metal USE is predicted to be 72 ft-lb at EOL. The 'O
- minimum plate USE is 89 f t-lb longitudinal at EOL. Branch Technical Position MTEB 5-2 (Reference 7) recommends 65% of the longitudinal USE as an estimate of transverse USE, so at EOL the plate USE would be 58 ft-lb transverse.
i !O I 2-4
- O I:
i i
,...m_-_ -- , , - . . . , _ - , _ - -.,y. ._,,,v,...- ,- - .-~.,,.c ,,,__ -
O _ n. Operating limits curves were constructed for three reactor conditions: hydrostatic pressure tests, non-nuclear heatup and cooldown, and core critical operation. The curves are valid up to 16 EFPY of operation. The limiting regions of the vessel affecting the curves' shapes are the core beltline (shifted to Q account for irradiation), the feedwater nozzle, and the closure flange region. The bolt preload and minimum permissible operating temporatures on the curves of 90*F provide' some additional margin in the closure flange region where a detectable O ficw size of 0.24 inch is used instead of 1/4 T. The operating limits curves for FitzPatrick are shown in Figures 2-1 through 2-3. lO
2.2 CONCLUSION
S The requirements of Reference 1 deal basically with EOL vessel conditions and with. limits of operation designed to prevent brittle fracture. Based on the evaluation of surveillance testing, the following conclusions are drawn:
- a. The adjusted reference temperature for the weld material of 104*F 0 is the limiting beltline EOL value. This is below the Reference 1 allowable limit of 200*F, above which annealing is required. The EOL values of USE for the plate and weld materials are 58 f t-lb transverse and 72 ft-lb, respectively. These are above the Reference 1 allowable of 50 ft-lb, below which annealing is required. Therefore, provisions for annealing the reactor vessel before completing 32 EFPY of operation need not be considered.
- b. Examination of the normal and upset operating conditions expected for the reactor shows that the worst pressure-temperature conditions expected from unplanned temperature transients are acceptable relative to the limits in Figures 2-1 through 2-3.
Therefore, the only ope ating conditions for which the operating limits are a concern are those involving operator interaction, such as hydrotest and initiation of core criticality. 2-5
b o sooO v Auo TO 16 E"Y A O saco - j 1200
- ADJUSTED BE LTLINE.
1/4T F LAW, ART = 680F D 1000 -
'e E
c 8 SAFE I OPE R ATING O h* - REGION E st IE O ** - 1200F i
/
O * - 312 psi 9 3DLT PRELOAD O 200 - TEMPER ATURE
- 900F fyLANGE REGION RTNOT*3 O ' '
o, 55 200 300 MINIMUM VESSEL METAL TEMPER ATURE 10FI Figure 2-1. Pressure Versus Minimum Temperature for Hydrostatic Pressure Tests O for FitzPatrick 2-6
O O 1800 VALID TO IS EFPY O ,.oo _ ADJUSTE D SE LTLINE. 1/4T FLAW ART = SSOF 0 ,, _ O .,ooo _ Y S i
- O g ,00 _
1 E 3 SAFE C OPE R ATING [, REGION O . _ !O .o0 _ NON BE LTLINE FW NOZZLE LIMITS, 1/4T F LAW, RTNDT = 300F .O y _ SOLT PRELOAD TEMPER ATURE = 900F FLANGE REGION RTNot = 300F 'O O i i O 100 200 300 MINIMUM VESSEL METAL TEMPER ATURE PF) Figure 2-2. Pressure Versus Minimum Temperature for Non-Nuclear Heatup and O Cooldown for FitzPatrick 2-7 - _ - _ _ . _ - - _.-. . . _ _ _ _ .. . - _. _.-_
O
!O~ te00 VALID TO 16 EFPY h 1400 -
ADJUSTED BE LTLINE. 1/4T F LAW, ART
- 680F
.O ,m -
O iom -
?
e w Z iO ! .00 - a NON dELTLINE f W NOZZLE LIMITS j E PLUS 400F.1/4T FLAW !O .00 - aTNoT So F l 0 .x - SAFE OPE R ATING REGION 200 - FLANGE REGtON RTN DT
- 300 F ,
l MINIMUM PERMIS$1BLE TEMPER ATURE = 900F l PE R 10CFR50 ! APPENDIX G O O i ' 0 100 200 300 l MINIMUM VESSEL METAL TEMPER ATURE (OFI Figure 2-3 . Pressure Versus Minimum Temperature for Core Critical Operation jo for FitzPatrick i 2-8
O
- 3. SURVEILLANCE PROGRAM BACKGROUND 0
3.1 CAPSULE RECOVERY The FitzPatrick reactor was shut down in February, 1985 for refueling O and maintenance. The accumulated thermal power output was 5,318,900 mwd or 5.98 EFPY. The reactor pressure vessel (RPV) originally contained three surveillance capsules, at 30', 120* and 300* azimuths at the core midplane. The specimen capsules are held against the RPV inside surface by a spring loaded specimen holder. Each capsule receives equal irradiation because of core symmetry. During the outage, Capsule 1 at 30' was removed. The capsule was cut from the holder assembly and shipped by a 200 Series cask to the General Electric Vallecitos Nuclear Center in Pleasanton, O California. Upon arrival at Vallecitos, the capsule was examined for identification. The reactor code of 29 and the basket code of 13 from O Reference 8 were confirmed on the capsule, as shown in Figure 3-1. The capsule contained three Charpy specimen packets and four tensile specimen tubes. Each Charpy packet contained 12 Charpy specimens and 3 flux wires. The four tensile specimen tubes contained eight specimens. The specimen gage sections were protected by aluminum sleeves, and during removal of the sleeves, the threaded ends of the specimens were slightly damaged. The threads were later chased with a die-hex rethreading tool. The gage sections of the tensile specimens were not damaged during removal. 3.2 RPV MATERIALS AND FABRICATION BACKCROUND 3.2.1 Fabrication Historv v The FitzPatrick RPV is a 218-in. BWR/4. It was constructed by Combustion Engineering to the 1965 ASME Code with Addenda up to and including Winter 1966. The shell and head plates are ASME SA-533 Grade B,
')
Class 1 low alloy steel (LAS). The nozzles and closure flanges are 3-1 O l
;O :
ASTM A508 Class 2 LAS and the closure flange bolting materials are n" ASTM AS40 Grade B24 LAS. The fabrication process employed quench and temper heat treatment immediately after hot forming, then submerged arc welding and post-weld heat treatment. The post-weld heat treatment was typically 6 hours at 1150*F ! 25*F. The arrangement of plates and welds
- relative to the core-beltline and various nozzles is shown in Figure 3-2.
3.2.2 Material Properties of RPV at Fabrication A search of General Electric Quality Assurance (QA) records was made to determine the chemical properties of the plates and welds in the RPV beltline. Table 3-1 shows the chemistry data obtained for the beltline materials. All data shown for the beltline plates were taken from QA records. Chemical composition of the beltline welds was obtained from Combustion Engineering (Reference 9), but as-welded data for the longitudinal seam welds was not available. Since the surveillance weld metal specimens were fabricated with the same weld procedure as was used in
'O the longitudinal seam welds, samples from the specimens were analyzed to obtain a representative chemistry. The results are presented in Subsection 3.2.3.
D ^ A search of QA records was made to collect results of certification mechanical property tests performed during RPV fabrication, specifically tensile test, Charpy V-Motch and dropweight impact test results. Properties of the beltline materials and other locations of interest are 'O presented in TabJe 3-2. The Charpy data collected were used to establish
- the RT val es fr each vessel component, as described in NDT Subsection 3.2.4.
lO
- O l
l 3-2 O i
1 O l I 3.2.3 Specimen Chemical Composition ,O Samples were taken from tensile specimens SCL and SCM (base) and from SDL and SDM (weld) after they were tested. The tensile specimens were fabricated from the same plates and weld as were the Charpy specimens, as O- detailed in Subsection 3.3, so the samples chosen are representative of all surveillance specimens. Chemical analyses were performed using a plasma l emission spectrometer. Each sample was decomposed and dissolved, and a 2 portion prepared for evaluation by the spectrometer. The spectrometer was O calibrated with a standard solution containing 700 ppm Fe, 8 ppm Mn, 2 ppm Cu, 5 ppm Ni, 5 ppm Mo, 5 ppm Cr, 1 ppm Si, 1 ppm Co, and levels of ! perchloric acid and lithium consistent with the test. The calibration for phosphorus was done by analyzing a series of seven National Bureau of
- O Standards steels with known quantities of phosphorus. The chemical composition results are given in Table 3-3.
! 3.2.4 Initial Reference Temperatures I l
- O The requirements applicable to establishing the RTNDT '# * * "E*#
1966 Edition of the ASME Code can be summarized as follows for the RPV: n N a. Test specimens shall be longitudinally oriented Charpy V-Notch l specimens.
- b. At the RT NDTe n impact test result shall be less than 25 ft-lb, and the average of three test results shall be at least 30 ft-lb.
t
- c. Pressure tests shall be conducted at a temperature at least 60*F above the acceptable RT NDT f r the vessel.
! The current requirements for establishing RT are sign W cantly NDT f different. For plants constructed to the ASME Code after Summer 1972, the i requirements are as follows: lO i
- a. Charpy V-Notch specimens shall be oriented normal to the rolling direction (transverse).
3-3 lO l 1
J g b. RT NDT is defined as the higher of the dropweight NDT or the temperature 60*F above which Charpy V-Notch 50 f t-lb energy and 35 mils lateral expansion are met. g c. Bolt-up in preparation for a pressure test or normal operation v shall be performed at or above the RT r 1 w st service NDT temperature (LST), whichever is greater, Reference 1 states that for vessels constructed to a version of the g ASME Code prior to the Summer 1972 Addendum, fracture toughness data and data analyses must be supplemented in an approved manner. General Electric has developed methods for analytically converting fracture toughness data
,, for vessels constructed before 1972 to comply with current requirements.
v The methods used have been presented to the NRC in about 10 Final Safety Analysis Report updates. These were reviewed on a case by case basis and in each ic'ase approved (e.g., La Salle 1 and 2 and Nine Mile Point 2). These methods and example RT calculations for vessel plate, veld, weld O NDT HAZ, forging, and bolting material are summarized in the remainder of this subsection. Calculated RT NDT va ues for selectd RW locadons are ghen in Table 3-2. O For vessel plate material, the first step in calculating RTNDT '" establish the 50 ft-lb transverse test temperature given longitudinal test specimen data. There are typically three energy values at a given test temperature. The lowest energy Charpy value is adjusted by adding 2*F per O f t-lb energy to 50 f t-lb. 'For example, for the six beltline plates the limiting combination of test temperature and Charpy energy from Table 3-2 is 43 ft-lb at +40*F for Plate G-3415-3. The equivalent 50 ft-lb n longitudinal test temperature is: v T 50L
= 40*F + [(50 - 43) ft-lb
- 2*F/ft-lb] = 54*F The transition from longitudinal data to transverse data is made by adding g
30*F to the test temperature. In this case, the 50 ft-lb transverse Charpy test temperature is T 50T
= 8W. ne RT NDT
" E# # # #
(T 50T
- 60*F). From Table 3-2, the NDT for G-3415-3 is -10*F. Therefore,
*h" "T FDT frt c re wtune is 2W.
O 3-4
O For vessel weld material, the Charpy V-Notch results are usually O The 50 ft-lb test temperature limiting in establishing RT NDT. is established as for the plate material, but the 30*F adjustment to convert longitudinal data to transverse data is not applicable to weld metal. The limiting beltline weld Charpy V-Notch energy from Table 3-2 is 36 ft-lb at
,O +10*F for veld 2-233. The corresponding transverse 50 f t-lb temperature is:
T 50T
=T 50L = 10*F + [(50 - 36) ft-lb
- 2*F/ft-1b] = 38'F As shown in Table 3-2, there are no NDT data available for the veld metal.
As long as the resulting RT is above -50*F, the RT NDT is esta W shed as NDT (T - 60 4), or RT = -22*F for this core beltline weld. 50T NDT For the vessel weld HAZ material, the RT NDT is assumed the same as for the base material since ASME Code weld procedure qualification test requirements and post-weld heat treatment data indicate this assumption is valid. Foi vessel forging material, such as nozzles and closure flanges the method for establishing RT is the same as for vessel plate material. NDT The recirculation inlet nozzle, listed in Table 3-2, has a calculated value I = 40*F. However, the NDT reported in Table 3-2 is +30*F. 50T Therefore, with the RT being the greater of NDT or (T - 60*F), the NDT 50T RT
- NDT For bolting material, the current requirements define the LST as the temperature at which transverse Charpy V-Notch energy of 45 ft-Ib and 25 mils lateral expansion (MLE) are achieved. If the required Charpy results are not met, or are not reported, but the Charpy V-Notch energy reported is above 30 ft-lb, the requirements of the ASME Code at construction are applied. As shown in Table 3-2, the limiting charpy V-Notch energy for bolting material is 39 ft-lb at +10*F. The current requirements are not met, so the construction Code requirements are used.
3-5 O
O The LST is defined as 60'F above the temperature at which 30 ft-lb Charpy O V-Notch energy is achieved. Therefore, the LST of the closure bolting material is 70*F. 3.3 SPECIMEN DESCRIPTION O The surveillance capsule contained 36 Charpy specimens: base metal (12), weld metal (12), and HAZ (12). There were 8 tensile specimens: base metal (3), weld metal (2), and HAZ (3). The 9 flux wires recovered were iron D' (3), nickel (3) and copper (3). The chemistry and fabrication history for the Charpy and tensile specimens.are described in this section. 3.3.1 Charpv Specimens O The fabrication of the Charpy specimens is described in the Surveillance Test Program description given in Reference 10. All materials used for specimens were beltline materials from the lower intermediate ()' shell course. The base metal specimens were cut from plate G-3414-2 from the beltline. The chemical analysis of this heat of low alloy steel is in
'O Table 3-1. The test plate was heat treated for 40 hours at 1150*F 2 25'F to simulate the post-weld heat treatment of the vessel. The method used to machine the specimens from the test plate is shown in Figure 3-3.
i Specimens were machined from the 1/4 T and 3/4 T positions in the plate, in
- O the longitudinal orientation (long axis parallel to the rolling direction).
l The identifications of the base metal Charpy specimens recovered from the I surveillance capsule are shown in Table 3-4. i iO The weld metal and HAZ Charpy specimens were fabricated from trim-off , pieces of plates G-3414-1 and G-3414-2 that were welded together with a l veld identical to longitudinal seam weld 1-233 in the RPV beltline. The l chemical analyses of the plates are given in Table 3-1. As-welded lO l' chemistry data were not reported in Reference 9 for the weld metal. The weld metal chemistry from two specimens is presented in Table 3-3. ! 3-6 !O l
I
!O The welded test plate for the weld and HAZ Charpy specimens received a O' heat treatment of 1150*F 2 25'F for 40 hours to conservatively simulate the fabricated condition of the RPV. The weld specimens and HAZ specimens were fabricated as shown in Figures 3-4 and 3-5. respectively. The base metal orientation in the weld and HAZ specimens was longitudinal. Contained in O Table 3-4 are the identifications of the weld metal and HAZ Charpy specimens from the surveillance capsule.
3.3.2 Tensile Specimens O Fabrication of the surveillance tensile specimens is described in Reference 10. The chemical composition and heat treatment for the base, weld and HAZ tensiles are the same as those for the corresponding Charpy O specimens. The identifications of the base, weld, and HAZ tensile specimens recovered from the surveillance capsule are given in Table 3-4. A summary of the fabrication methods is presented in the remainder of this section. =Q The base metal specimens were machined from material at the 1/4 T and 3/4 T depth in plate G-3414-2. The specimens, oriented along the plate rolling direction, were machined to the dimensions shown in Figure 3-6. -O The gage section was tapered to a minimum diameter of 0.250 inch at the center. The weld metal tensile specimen material was cut from the welded test plate, as shown in Figure 3-7. The specimens were machined entirely from weld metal, scrapping material that might include base metal. The O fabrication method for the HAZ tensile specimens is illustrated in Figure 3-8. The specimen blanks were cut from the welded test plate such that the gage section minimum diameter was machined at the weld fusion line. The finished HAZ specimens are approximately half weld metal and
-Q half base metal oriented along the plate rolling direction.
!O f 3-7 O
() C) () () () () CJ () () L) (s Tcblo 3-1 CHEMICAL COMPOSITION OF RPV BELTLINE MATERIALS Composition by Weight Percent Identification Heat / Lot No. C Mn P S Si Ni Mo Cu Lower Plates: C-3915-lR C3394-1 0.21 1.32 0.015 0.017 0.26 0.56 0.47 0.11 G-3415-3 C3376-2 0.22 1.33 0.015 0.017 0.22 0.60 0.48 0.13 G-3415-2 C3103-2 0.23 1.36 0.012 0.016 0.26 0.60 0.47 0.14 Lower-Intermediate Plates: G-3413-7 C3368-1 0.21 1.31 0.015 0.018 0.23 0.54 0.46 0.12 G-3414-2 C3278-2 0.20 1.26 0.011 0.016 0.23 0.60 0.48 0.13 G-3414-1 C3301-1 0.20 1.35 0.009 0.015 0.27 0.60 0.48 0.18 Lower Longitudinal Welds: 2-233 A,B,C Heat 27204, Flux not available Y 1092, Lot 3774 Heat 12008, Flux not available 1092 Lot 3774 Heat 8018 0.064 0.81 0.008 0.015 0.17 1.04 0.26 n/a Lot E0AG Lower-Inte rmediat e Longitudinal Welds: 1-233 A,B.C Heat 13253, Flux not available 1092, Lot 3774 Heat 12008, Flux not available 1092, Lot 3774 Heat 8018 see above Lot E0AG Heat 8018 0.081 1.02 0.009 0.011 0.49 0.93 0.21 n/a Lot LODG Lower to Lower-Int. Girth Weld: 1-240 Heat 305414, Flux 0.14 1.45 0.012 0.010 0.18 0.59 0.51 0.33 1092, Lot 3947 Heat 8018 0.080 1.03 0.012 0.012 0.40 0.91 0.24 0.02 Lot B0BJ Heat 8018 0.079 1.04 0.011 0.012 0.42 0.93 0.23 0.03 Lot AOFJ
Tabis 3-2 RESULTS OF FABRICATION TEST PROGRAM FOR SELECTED RPV LOCATIONS 4 Tensile ; } Total Area Test Charpy T -60 RT i Ident. Heat Yield UTS Elong Reduc. Temp. Energy NDT SOT NDT Location Number Number (ksi) (ksi) (%) (%) (*F) (ft-lb) (*F) (*F) (*F) - Lower Shell Plates G-3415-lR C3394-1 67.8 88.5 27.5 69.8 10 53.71,52 -10 -20 -10 l G-3415-3 C3376-2 67.0 88.3 27.0 68.6 40 43,51,49 -10 24 24 G-3415-2 C3103-2 66.1 89.5 26.3 69.3 10 41,48,49 -10 -2 -2 i l Lower-Intermediate G-3413-7 C3368-1 65.3 86.5 27.0 67.7 10 61,55,45 -50 -10 -10 Shell Plates G-3414-2 C3278-2 67.7 89.3 27.0 69.8 10 45,77,58 -30 -10 -10
- G-3414-1 C3301-1 70.9 92.6 26.5 67.9 10 60,63,49 -40 -18 -18
! Longitudinal Weld 2-233 Ht. 27204 78.9 93.5 24.0 66.1 10 49,48,36 n/a -22 -22 l Lot 3774 1 i Girth Weld 1-240 Ht. 305414 72.1 87.7 25.5 66.8 10 82,66,80 n/a -50 -50 Lot 3947 j Upper Shell Plate G-3413-5 C3229-2 68.9 90.5 27.3 68.0 10 50,68,82 -10 -20 -10 j Vessel Flange G-3401 2V595 63.6 85.4 28.5 73.0 10 117,94.117 10 -20 10 , Head Flange G-3402 4P-1885 70.9 92.3 25.5 70.5 10 66,87,96 30 -20 30 l l Top Head Torus G-3411-1 C3055-1 70.6 92.6 26.0 70.6 10 98,73,118 -10 -20 -10 l Recire. Inlet G-3436-9 E21VW-104J5 75.8 93.8 26.0 71.5 10 73,116,118 30 -20 30 Closure Bolts G-3134-1 37385 153.3 168.0 15.0 47.3 10 39,40,39 n/a LST = 70*F ! I
O Table 3-3 O PLASMA EMISSION SPECTROMETRY CHEMICAL ANALYSIS OF RPV SURVEILLANCE PLATE AND WELD MATERIALS () Base Metal Base Metal Weld Metal Weld Metal Element Tensile SCL Tensile SCM Tensile SDL Tensile SDM Mn 1.4 1.3 1.5 1.4
) P 0.011 0.010 0.015 0.014 Cu 0.11 0.12 0.31 0.31 Si
- 0.07 0.06 0.06 0.06 N1 0.62 0.63 0.72 0.72
( Mo 0.48 0.50 0.50 0.51 Cr 0.11 0.11 0.04 0.04 Co 0.011 0.011 0.020 0.019 C) i 'O-l- l l O 0
- Si results may be low, due to precipitation during dissolution heating.
O l 3-10
- O l -- _ _. . . . , _ , _ _ _ _ _ _ . _ _ _ . _ ..__ _ _ _ . _ _ _ _ __ _ _ _ _ _ . _ _ . _ _ . _ _ _ . _ _ ___ _ __
O t Table 3-4 O IDENTIFICATION OF CHARPY AND TENSILE SPECIMENS REMOVED FROM SURVEILLANCE CAPSULE Q Charpy Specimens Base Weld HAZ O 51B 541 572 51C 54B 57K 51D 54D 57T 51L 54L 57Y o 51Y 54P 5A1 52A 552 SAA 52B SSA 5AC 52C 55D 5B4 52K 55J 5BL 52M 55U 5BM 52U SGE SBP 531 5GM SBU C i Tensile Specimens Base Weld HAZ SCL SDL SEB !O SCM SDM SED 5CT SEM l iO l i i 3-11 iO l i
._ _ .._ .. ..._ -_ . . _ . ~. . . , - - . . . . - - - _ _ _ _ _ . ._ , _ _ _ _ _ _ , _ _ , _ _
l O f
;O RE ACTOR CODE 16 e
- 9 8 e o IO
- 1 + 4 + 8 + 16 = 29 2
e l 1 N l lO l
..5 TMi ' . . l 1
.e i ,, h
w, % .y l
.Q l '
f g; g;g 1 A ; @ ::. ., . - i o .] .. O
)L CAPSULE CODE 8 e O 2
- ,.<.8.n 1 e e ,
JL CAPSULE 1 10 l ? Figure 3-1. Surveillance Capsule Recovered from FitzPatrick Reactor i O
O O E E E I y,,,,, [ TOP HE AD ENCLOSURE CLOSURE / ' ' " " ' ' ' '
\
~
FLANGE , E r- REG l0N wmmHmm/wn//MHwwwwwwwn1
' I d I s
UPPER SHELL O h 1 ,, ,, 1,1, , , 3 ( 0 l @ O j UPPER INTERME DI ATE l SHELL l l l. E I 155
- s -
l LOWER l b INTERMEDIATE l G-3414-2 { G3413-7 s SHELL l G34141 ' CORE { l /' BE LTLIN E : LONGITUDINAL l REGION : SE AM WELD l ,0 1 233 { l { l
' jCIRCUMF ERENTI AL GIRTH WELD lf d 1-240 l 2-233 0 _
w l G-34152 LOWER l SHELL l G-34153 G-34151 R l l - f fJ JJJJ Jf II JA Z I II J O BOTTOM HE AD ENCLOSURE ,, , , ,, , 10 ~.,,,,,,,,,,,,:. N 1 I I Figure 3-2. Schematic of the RPV Showing Arrangement of Vessel Flates and Welds !O 3-13
O O gg5gO H 9.gla 'M b POL O 1/2 m-1 g\ s G-3414-2 gg (HE AT C327&2)
= ; \\ g\
o6 0 b s5i m. s s\
, \s
\\
4 SLABS TOTAL g\g\
\ 1/
15 PIECES g\ O PER SLAs \ s s\-y 'qp 3/4T
= ; \s
\b O 9
'O y 90 deg 2 0 deg 10 men + 1.082 + 20 010 M ARKED END g , y UNMARKED END I O h I
= 2.16520 015 =
22-1/2 deg 0 010r 0 001R
~
O + + 0.394 p U
\W - 90 deg 2 0 deg 10 men 0.394 ALL CORNERS 0.315 h h O
Figure 3-3. Fabrication Method for Base Metal Charpy Specimens O 3-14
O ROLLING DIRECTION _- G-34141 (HE AT C33011) U 5/8 in.
.g L
WE LO 1233
# G-34142 O SCRAP ' (HE AT C3278 2) 1 b O N I
6 SLABS i 6-1/2 an.
\
U ~ .O _ c - 21/4 in. -
\
N\ /
\N x N
! x N N to SPECIM S \ PER SLAB MACHINED AS \ \\g IN FIGURE 3-3
\\ \ \ \
\ \
lO
\x\
N x x N N N x N
\\\
5/8 in\ h^ i e N O Figure 3-4. Fabrication Method for Weld Metal Charpy Specimens lO 3-15 l . .. . - . _
- O
- O
- ROLLING DIRECTION - G-34141 (HE AT C3301-1)
JL - WELD 1233 G-34142 p (HE AT C327&2) 6 SLA8S - 1 l l i i 61/2 in.
\
SCRAP j jf I O ; + i r .n. J, i A \ \ %
- \ \\
N N N ~
- o x N N
! N N \ x NN N N N
\ \\N N
- D \ \
5/8 in
~
~g l lI
{ 10 SPECIMENS PER SLAB MACHINED M2-1/4 en. 1 AS IN FIGURE 3-3 WITH NOTCH l Q CENTERED ON WELD FUSION LINE 1 Figure 3-5. Fabrication Method for HAZ Charpy Specimens O 3-16 _ m.- , .- . ,-.--,_,.m,.,,,... ,-..-.,--e.-.-- --... ,,_ . - ..~,._,.,,... --,__ . _ . . . - - , , . - . - - - - -
O O ROLLING DIRECTION O
\ \
G.34142
\ (HE AT C327&2) x x 0 N N
\ \
10 SLABS \ \ \ O N \ \\
\ \5/8 n.
p __ N s N o
\ h 1/2 in %_ _ _ _ _ 1/4 T II Q
r s
\ --- -
q-3 /4 T x\ 61/2 '" s N
- s. _ _ _ _ " N Q (\ v Q :_ - @ ~_}
3 5/8 in. O C 1.000r0 005 GAGE LENGTH 7/16 - 14 UNC - 2A 3/8R ~ ~ BOTH ENDS 1/2- l (TYP) Q ,.... b. j
} AGE MARKS -....s I ' D' 30 DEG O D' l D TYP p T- l 3M H -
1/16 + g3 fp 9j4 ,1 1/2.in. RE DUCED , 2 1/7 = g SECTION d 3s t /16 = NOTES
- 1. D = 0 0250:0 001 DIAM AT CENTER OF REDUCED SECTION
- 2. D' = ACTUAL "D" Ol AM + 0.002 TO O 005 AT ENDS OF REDUCED SECTION, TAPERING TO "D" AT CENTER O
Figure 3-6. Fabrication Method for Base Metal Tensile Specimens O 3-17
b O l O 1 1 \ l
^
I o . O T ST PLATE o WELD 1233 iO s ! (%, 6 SPECIMENS s PER SLA8 h5 y) 13/4 en. ': DIMENSIONS ( ASIN FIGURE 3 6 3 SLASS \qg h SCRAP i 6-1/2 in. ! JL 1/2 in 1j h O Figure 3-7. Fabrication Method for Weld Metal Tensile Specimens 'O 3-18
- t. _ - . -
_- . . . . - - - - . . - - - -.. .. -.. _ _ _ ~ . =.._~__ . ~ . _ ~ _ _ _ _ = _ - . . - ~ . Q T 1 lo ROLLING DIRECTION C r G 34141 (HE AT C3301 11 G 34142 (HE AT C327&2) [) WE LD 1233
- 0 x sx c x- ,
l 6-1/2 en f 3 SLABS
- 4 { ;
1
= 10 m. =
0
- O c 31/8 m. %
g 1/2o g
'- 6 SPECIMENS PER
\ ) SLAB DIMENSIONS [
t AS IN FIGURE 3-6 I SC"^" O l , T O Figure 3-8. Fabrication Method for HAZ Tensile Specimens O 3-19 I- - . - . _ , _ _ , . . . _ . . . . . . . . . _ _ _ . , _ , _ , . . , . _ _ . , . _ , _ _ , _ _ , , _ _ _ _ _ __ _ _ _
\
l O3
- 4. PEAK RPV FLUENCE EVALUATION O
Flux wires were analyzed to determine flux and fluence received by the surveillance capsule. A reference analysis combining two-dimensional and one-dimensional flux distribution computer calculations was evaluated to O establish the location of peak vessel fluence and the lead factors of the surveillance capsule relative to the peak vessel location. 4.1 FLUX WIRE ANALYSIS O 4.1.1 Procedure The surveillance capsule contained nine flux wires: three each of O iron, copper, and nickel. Each wire was removed from the capsule, cleaned with dilute acid, weighed, mounted on a counting card, and analyzed for its radioactivity content by gamma spectrometry. Each iron wire was analyzed for Mn-54 content, each nickel wire for Co-58, and each copper wire for O co-60 at a calibrated 4- or 10-cm source-to-detector distance with an 80-ce Ge(Li) detector system. The gamma spectrometer was calibrated using NBS material.
- O- To properly predict the flux and fluence at the surveillance capsule from the activity of the flux wires, the periods of full and partial power irradiation and the zero power decay periods were considered. Operating days for each fuel cycle and the reactor average power fraction are shown O in Table 4-1. Zero power days between fuel cycles are listed as well.
O O 4-1 O
N From the flux vire activity measurements and power history, reaction rates for Fe-54 (n.p) Mn-54, Cu-63 (n,a) Co-60, and Ni-58 (n.p) Co-58 were calculated. The >l MeV fast flux reaction cross sections for the iron, copper, and nickel wires were estimated to be 0.177 barn, 0.0031 barn and 0.230 barn, respectively. These values were obtained from measured cross O section functions determined at Vallecitos from more than 65 spectral determinations for BkTs and for the General Electric Test Reactor using activation monitor and spectral unfolding techniques. These data functions are applied to BkT pressure vessel locations based on water gap (fuel to O' vessel wall) distances. The cross sections for >0.1 MeV flux were determined from the measured 1-to-0.1 MeV cross section ratio of 1.6. 4.1.2 Results O The measured activity, reaction rate and determined full-power flux results for the surveillance capsule are given in Table 4-2. The >l MeV 9 9 and >0.1 MeV flux values of 1.4x10 and 2.2x10 n/cma -sec from the flux n
- ' monitors were calculated by dividing the reaction rate measurement data by the appropriate cross sections. The corresponding fluence results, 17 2 2.6x10 and 4.2x10 n/cm for >l MeV and >0.1 MeV, respectively, were obtained by multiplying the full-power flux density values by the product O 8 of the total seconds irradiated (2.75x10 sec) and the full-power fraction (0.686).
Generally, for long-term irradiations, dosimetry results f rom copper "a flux wires are considered the most accurate because of Co-60's long half-life (5.27 years). The secondary nickel and iron flux monitor reactions yielding Co-58 and Mn-54 gave results fairly consistent with the copper reaction despite the shorter half-lives of 70.8 days for Co-58 and 312.5 days for Mn-54. Consistency in results indicates an accurate power-history evaluation and a consistent core radial power shape. O 4-2 0
p$ 1 1 l 1 The accuracies of the values in Table 4-2 for a 2c deviation are O estimated to be: 5% for dps/g (disintegrations per second per gram) 12% for dps/ nucleus (saturated) i O 25% for flux and fluence >l Mev 35% for flux and fluence >0.1 MeV A set of flux wires from FitzPatrick was evaluated by General Electric 2 .O in 1982. The >l MeV flux was 1.5x10' n/cm -sec. The result from this 2 study of 1.4x10 n/cm -sec is very close. The flux wires from each analysis are from the same location, so consistent results are expected. !O 4.2 DETERMINATION OF LEAD FACTORS The flux wires detect flux at a single location. The wires will therefore reflect the power fluctuations associated with the operation of O the plant. However, the flux wires are not necessarily at the location of peak vessel flux. Lead factors are required to relate the flux at the wires' location to the the peak flux. These lead factors are a function of the core and vessel geometry and of the distribution of bundles in the O core. Lead factors were generated for a 218 inch BWR/4 geometrically identical to FitzPatrick and with a similar bundle distribution (Reference 11). The Reference 11 analysis was reviewed and it was concluded that the lead factors are appropriate for FitzPatrick
- O 8PPlication. The methods used to calculate the lead factors are discussed below.
!O O 4-3 ^O
O 4.2.1 Procedure O Determination of the lead factors for the RPV inside wall and at 1/4 T depth was done using a combination of one-dimensional and two-dimensional finite element computer analysis. The two-dimensional analysis established O the relative fluence in the azimuthal direction at the vessel surface and 1/4 T depth. A series of one-dimensional analyses were done to determine the core height of the axial flux peak and its relationship to the surveillance capsule height. The combination of azimuthal and axial O distribution results provides the ratio of flux, or the lead factor, between the surveillance capsule location and the peak flux locations. The two-dimensional DOT computer program was used to solve the Boltzman O transport equation using the discrete ordinate method on an (R,0) geometry, assuming a fixed source. Quarter core symmetry was used with periodic boundary conditions at 0* degree and 90*. Neutron cross sections were determined for 26 energy groups, with angular scattering approximated by a O third-order Legendre expansion. A schematic of the two-dimensional vessel model is shown in Figure 4-1. A total of 99 radial elements and 90 azimuthal elements were used. The model consists of an inner and outer core region, the shroud, water regions inside and outside the shroud, the vessel wall, and an 0 air region representing the drywell. Flux as a function of azimuth was calculated, establishing the azimuth of the peak flux and its magnitude relative to the flux at the wires' location of 30*.
- O The one-dimensional. computer code (SNID) was used to calculate radial flux distribution at several core elevations at the azimuth angle of 45*,
where the azimuthal peak was determined to exist. The elevation of the peak flux was determined, as well as its magnitude relative to the flux at IO the surveillance capsule elevation. O 4-4 .O
O 4.2.2 Results
'O The one-dimensional flux calculations established the elevation of peak flux at 106 inches above the bottom of active fuel, or 34 inches higher than the capsule. The two-dimensional calculation indicated the flux to be a O maximum 45 degrees on either side of the RPV quadrant references (0*, 90*,
etc.). The peak closest to the 30* location of Capsule 1 is at 45*. The distribution calculations establish the lead factor between the ~O surveillance capsule location and the peak location at the inner vessel wall. This lead factor is 0.79. The fracture toughness analysis done is based on a 1/4 T depth flaw in the beltline region, so the attenuation of the flux to that depth is considered. The resulting lead factor from the
- O capsule to the 1/4 T depth at the peak location is 1.05.
4.3 ESTIMATE OF END-OF-LIFE FLUENCE "O 'The fluence at end-of-life (EOL) is estimated by taking the upper bound of the measured flux from Table 4-2 and the 1/4 T lead factor. The period assumed to represent EOL is 32 EFPY, or 1.01x10 seconds. The resulting EOL fluence is: ..o 9 2 4 18 2 (1.4x10 n/cm -s)(1.25)(1.01x10' s)/1.05 = 1.7x10 n/cm . O O. O 4-5 0
O Table 4-1
- O l
SUMMARY
OF DAILY POWER HISTORY l~ Operating Percent of Days Between ! Cycle Cycle Dates Days Full Power Cycles 'O 1 1/26/75 - 6/21/77 877 0.523 93 2 9/22/77 - 9/16/78 359 0.796 81 3 12/6/78 - 5/6/80 517 0.513 C) 95 4 8/9/80 - 10/31/81 448 0.825 128 5 3/8/82 - 6/3/83 452 0.863 91
! 6 9/2/83 - 2/15/85 532 0.781 'o.
3185 0.686 (average) O O O
'o O
4-6 'O .
- ~_ .- . . -. _ - - _ . . .- .. . -. .. ..
Table 4-7 4 SURVEILLANCE CAPSULE LOCATION FLUX AND FLUENCE FOR IRRADIATION FROM 1/26/75 To 2/15/85 4 Wire dps/g Element Reaction Rate Full Power Flux
- Fluence Wire Weight (at end of [dps/ nucleus (n/cm2 -s) (n/cm2 )
(Element) (g) Irradiation) (saturated)] >l MeV >0.1 MeV >l MeV >0.1 MeV 4 -18 Copper 64742 0.4582 'l.35x10 4.49x10
-18 Copper 64743 0.4029 1.29x10 4.29x10 4 -18 Copper 64744 0.4181 1.26x10 4.20x10
-18 9 I7 17 Average = 4.33x10 1.4x10 2.2x10 2.6x10 4.2x10 5 -16 Iron 64742 0.1438 1.06x10 2.33x10 5 -16 Iron 64743 0.1471 1.05x10 2.31x10
-16 1ron 64744 0.1042 1.04x10 2.29x10 l Average = 2.31x10 -16 1.3x10 9
i 6 -16 I Nickel 64742 0.3113 1.44x10 2.63x10 6 -6 I Nickel 64743 0.2849 1.44x10 2.64x10 6 -6 Nickel 64744 0.3039 1.49x10 2.73x10 ! Average = 2.67x10 -16 1.2x10 9 l i
- Full power of 2436 MW .
I i
O 90" O g a 2 2 2 1 i i , i i i i i i h 2 2 2 2 1 1 1k 1 1 1 1 1 1 2 2 2 2 1 1 1 1 (1 1 1 1 1 2 2 2 2 1 1 1 1 1k 1 1 1 1 Q 2 2 2 2 1 1 1 1 N 1 1 1 1 2 2 2 2 1 1 1 1 1 h 1 1 1 REFLECTIVE BOUNDARY 2 2 2 2 1 1 1 1 1 1 1 : 7 2 2 2 2 1 1 1 1 1 1 1 , CORE INTERIOR; 4 ELEMENTS 2 2 2 2 1 1 1 1 1 O 1 1 O.
- 4 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 CORE EXTERIOR; 2 2 2 2 2 2 2 22 ELEMENTS 2 2 2 2 2 2 2
- WATER REGION; 90 ELEMENTS "
27 tiLEMENTS IN AZIMUTHAL DIRECTION f HROUD;4S ELEMENTS CAPSULE e , WATER RFGION; O 26 ELEMENTS X# k VESSELWALL; 4 f 14 ELEMENTS pORYWELL: 2 ELEMENT 3 10 O o o 1 = CORE INTERIOR FUEL 2 = CORE EXTERIOR FUEL O O Figure 4-1. Schematic of Model for Two-Dimensional Flux Distribution Analysis O 4-8
O
- 5. CHARPY V-NOTCH IMPACT AND HARDNESS TESTING
.O The 36 Charpy specimens recovered from the surveillance capsule were impact tested at temperatures selected to establish the toughness transition and upper shelf of the irradicted RPV materials. Testing was .O conducted in accordance with ASTM E23-82 (Reference 12). After impact testing, Rockwell C hardness testing was performed on the broken specimen halves per ASTM E18-79 (Reference 13).
O 5.1 IMPACT TEST PROCEDURE The testing machine used was a Riehle Model PL-2 impact machine, serial number R-89916. The pendulum has a maximum velocity of 15.44 ft/sec O and a maximum available hammer energy of 240 ft-lb. The test apparatus and operator were qualified using U.S. Army Watertown standard specimens. The standards are designed to fail at 74.1 ft-lb and 13.9 ft-lb at a test temperature of -40*F. According to Reference 12, the test apparatus O averaged results must reproduce the watertown design values within an accuracy of !5% or 21.0 ft-lb, whichever is greater. The successful qualification of the Riehle machine and operator is summarized in Table 5-1. 10 Charpy V-Notch tests were conducted at temperatures between -60*F and 400*F. For tests between 32*F and 212*F, the temperature conditioning fluid was water. Dichloromethane was used at temperatures below 32*F. O Above 212*F, a silicone oil was used. Cooling of the conditioning fluids was done with liquid nitrogen, and heating by an immersion heater. The fluids were mechanically stirred to maintain uniform temperatures. The fluid temperature was measured by a chromel-alumel thermocouple and a O copper-constantan thermocouple. These were calibrated with boiling water (212*F), and ice water (32*F). Once at test temperature, the specimens were manually transferred with centering tongs to the Riehle machine and impacted within 5 seconds. O 5-1 O i
O For each Charpy V-Notch specimen tested, test temperature, energy 'O absorbed, lateral expansion, and percent shear were evaluated. Lateral expansion and percent shear were measured according to Reference 12 methods. Percent shear was determined with method two of Subsection 11.2.4.3 of Reference 12, which is a comparison of the fracture O surface appearance with the reference fracture surfaces in Figure 15 of Reference 12, 5.2 IMPACT TEST RESULTS
- O Twelve Charpy V-Notch specimens each of base metal, weld metal and HAZ were tested at temperatures selected to define the toughness transition and upper shelf portions of the fracture toughness cut . Absorbed energy.
O lateral expansion and percent shear data are listed in Table 5-2 for each material. Plots of absorbed energy data for base, veld, and HAZ metal are presented in Figures 5-1, 5-2 and 5-3, respectively. Lateral expansion plots for base, weld and HAZ metal are given in Figures 5-4, 5-5 and 5-6,
- O respecti.ely.
The data sets are freehand fit with best-estimate S-shaped curves characteristic of fracture toughness transition curves. The HAZ data in .O Figures 5-3 and 5-6 are not fit with a curve. There is too much scatter in
- the data for a meaningful curve to be drawn. The HAZ data probably show I
the greatest scatter because the HAZ has been uniquely heat treated by the welding process. In addition, the uncertainty of the specimen notch
- O location relative to the weld fusion line causes scatter in the HAZ results.
Photographs were taken of the fracture surfaces for each specimen. O The fracture surface photographs were used to evaluate percent shear. The photographs and a summary of test results for each specimen are contained in Appendix A. O O 5-2
O , 5.3 IRRADIATED VERSUS UNIRRADIATED CHARPY V-NOTCH PROPERTIES O As a part of the RPV fabrication test program, Charpy V-Notch testing was done at various temperatures on the unirradiated RPV plate materials. Data for the beltline plate from which the base metal specimens were .O fabricated, (plate G-3414-2) were recovered from QA records. The impact energy and lateral expansion data are plotted in Figures 5-7 and 5-8, respectively. The corresponding irradiated data from the base metal surveillance specimens are plotted in Figures 5-7 and 5-8 as well. O The curves of irradiated and unirradiated Charpy V-Notch properties are used to estimate the values in Table 5-3: 30 ft-lb, 50 ft-lb and 35 MLE index temperatures and USEs, The RT s W t values are dete M ned NDT
- O as the change in the temperature at which 30 ft-lb impact energy is ;
achieved, as required in Referet.ce 4. In previous experience, the shift in the transition curve has been approximately equal at the 30 ft-lb and 50 f t-lb levels. Table 5-3 shows a slight difference (23*F versus 26*F, O respectively). The shift in lateral expansion at the 35 MLE level of 13*F supports a lower shift value, so the 30 ft-lb shift of 23*F is used as the RT shift. This shift from initial to adjusted RT f r the base metal g ET is compared to analytical values calculated in Section 7 according to
- O Regulatory Guide 1.99, Revision 1. A similar comparison cannot be made for the weld metal because unirradiated weld metal Charpy test data are not available.
O 5.4 ROCKWELL HARDNESS TESTING After the Charpy specimens were tested, one broken half of each specimen was subjected to Rockwell hardness testing, according to ~O ASTM E18-79 (Reference 13). The test used for the surveillance materials was the Rockwell C test, which employs a diamond sphero-conical penetrator with a minor load of 10 kgf and a major load of 150 kgf.
- O 5-3 0
4O The machine used was a calibrated Wilson Rockwell Hardness Tester. As 10 a further calibration before testing, a test block with a reference hardness of 35.0 1 1.0 HRC was tested. The three values taken were 34.0, 34.2 and 36.0 HRC, for an average of 34.7 HRC, which is acceptable.
.O Three indentations were made on each specimen half and the results were averaged to develop the hardness values in Table 5-4. Each piece tested was consistently the half with the specimen identification stamped in the end. The indentations were made on the same side of each specimen O in a group approximately 3/8 inches from the fracture surface.
The results show a definite difference between weld and base metal. The weld metal grouping averages 5 HRC above the base metal average. As
.O discussed in section 3, the HAZ specimens are half weld and half base metal. The results of Table 5-4 indicate that the weld side of the HAZ specimens was tested.
'O O O i O !O 5-4
.O , --r - , , , , - - . - . + - - - . , ,-. _ .---- - --- . - - . - - - . , . - - . - - . - . - - - - - - __ - - - - . - . , -
O Table 5-1
- O QUALIFICATION TEST RESULTS USING U.S. ARMY WATERTOWN SPECIMENS (TESTED IN FEBRUARY 1986)
.O Energy Absorbed Qualification Test Test Temperature Mechanical Gage Specimen Identification (*F) (ft-lb) lO EE30427 -40 72.0 EE30075 74.5 .O EE30233 80.0 EE30956 75.8 EE30365 74.2 .O Average 75.3 1 Allowable -40 74.1 ! 3.7 Acceptable DD50442 -40 15.0
" 13.8 DD50905 lO DD50486 14.0 DD50977 14.0 DD50831 14.0
- Average 14.2
.O 1 Allowable -40 13.9 ! 1.0 Acceptable
- O 5-5
-O l J
O Table 5-2 O-CHARPY V-NOTCH IMPACT TEST RESULTS FOR IRRADIATED RPV MATERIALS Test Fracture Lateral Percent Shear
.O Specimen Temperature Energy Expansion (Method 2)
Identification (*F) (ft-lb) (mils) (%) Base: 52K -60 14.3 15 0 52C -20 21.0 26 0 0 52B 0 46.5 46 0 531 10 61.0 57 10 52A 20 67.5 61 20 51B 40 44.5 51 20 51L 80 83.6 70 50 Sic 120 100.0 68 60
- O SiY 160 125.0 86 50 51D 200 136.5 92 90 52M- 300 136.5 94 100 52U 400 121.3 92 80 Weld:
O 55J -60 7.2 11' O 55D -20 24.0 26 20 55A 0 15.0 20 10 56M 10 14.0 23 0 552 20 20.0 25 30 541 40 31.4 33 20 O 54L 80 42.0 39 40 . 54B 120 54.1 52 70 54P 160 75.0 62 85 54D 200 82.1 53 70 55U 300 86.0 65 60 56E 400 79.3 70 100 0 ! HAZ: 5BL -60 13.2 16 10 SB4 -20 38.3 35 20 SAC 0 78.5 70 50 5BU 10 26.5 28 40 0 SAA 20 86.7 64 60 572 40 60.1 51 50 57Y 80 61.5 46 60 57K 120 110.6 65 50 sal 160 81.2 66 100 57T. 200 107.5 70 100 'O SBM 300 87.2 71 60 SBP 400 72.0 70 100 5-6
- O
.. ,..,,m m, ,-v_ _ .e. -- . , , . .,. - , . .-._.~,._..__._y- ,e,_ _ y -.
O Table 5-3 O SIGNIFICANT RESULTS OF IRRADIATED AND UNIRRADIATED CHARPY V-NOTCH DATA ()~ Upper # Shelf Index Temperature (*F) Energy (ft-lb) () Material E=30 ft-lb E=50 ft-lb MLE=35 mil L/T Unitradiated Plate -32 1 -13 130/85 O Irradiated Plate -9 27 0 122/79 Difference 23 26 13 8/5 (6%) O Irradiated Weld 48 100 61 82/82 O O () Longitudinal (L) USE is read directly from Figures 5-1 and 5-3. Transverse (T) USE is taken as 65% of the longitudinal USE, according to Reference 7. L/T USE values are equal for veld metal, which has no orientation effect. O 5-7 O
0-Table 5-4 O ROCKWELL C HARDNESS TEST RESULTS Specimen Rockwell C Identification Type Hardness (HRC) O 52K Base 14.3 52E 13.9 52A 13.7 531 13.4 52M 13.3 C) 52B 12.8 51Y 12.6 51B 12.5 52U 12.4 51D 12.3 51C 12.2 () 51L " 12.0 Average = 13.0 54L Weld 21.3 54D 20.1 54B 18.8 C) 552 18.6 54P 18.4 55A 18.3 55U 18.2 55J 18.1 541 18.1 C)' 55D 17.6 56M 17.4 56E 16.7 Average = 18.5 57Y HAZ 21.1 () 57T " 20.7 57K 19.3 5AA 19.3 572 18.9 SBM 18.9 5A1 18.6 C) SBL 18.3 SBU 18.3 5BP 18.1 SAC 17.5 SB4 17.0 Average = 18.8
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- 6. TENSILE TESTING O
Eight round bar tensile specimens were recovered from the surveillance capsule. Uniaxial tensile tests were conducted in air at room temperature, RPV operating temperature, and onset of upper shelf temperature. Tests C) were conducted in accordance with ASTM E8-81 (Reference 14). 6.1 PROCEDURE I) All tests were conducted using a screw-driven Instron test frame equipped with a 20-kip load cell and special pull bars and grips. Heating was done with a Satec resistance clamshell furnace centered around the specimen load train. Test temperature was monitored and controlled by a C) chromel-alumel thermocouple spot-welded to an Inconel clip that was friction-clipped to the surface of the specimen at its midline. Before the elevated temperature tests, a profile of the furnace was conducted at the test temperature of interest using an unirradiated steel specimen of the () same geometry. Thermocouples were spot-welded to the top, middle, and bottom of a central 1 inch gage of this specimen. In addition, the clip-on thermocouple was attached to the midline of the specimen. When the target temperatures of the three thermocouples were within 5'F of each other, the i) temperature of the clip-on thermocouple was noted and subsequently used as the target temperature for the irradiated specimens. All tests were conducted at a calibrated crosshead speed of
$) 0.005 inch / min until well past yield, at which time the speed was increased to 0.05 inch / min until fracture. A 1 inch span knife edge extensometer was attached directly to each specimen's central gage region and was used to monitor gage extension during test. .O The test specimens were machined with a minimum diameter of 0.250 inch
- I at the center of the gage length. The three specimens each of base metal and HAZ were tested at room temperature (RT = 76*F), onset of upper shelf j)> temperature (estimated at 185'F), and RPV operating temperature (550*F).
6-1
-O
O The two weld metal specimens were tested at room temperature and 550*F.
'O - The yield strength (YS) and ultimate tensile strength (UTS) were calculated by dividing the nominal area (0.0491 in.2) into the 0.2% offset load and into the maximum test load, respectively. The values listed for the uniform and total elongations were obtained from plota that recorded load
- O versus specimen extension and are based on a 1 inch gage length. Reduction of area (RA) values were determined from post-test measurements of the necked specimen diameters using a calibrated blade micrometer and employing the formula
- O l
RA = 100% * (A 9
- Ag )/A9 After testing, each broken specimen was photographed end-on showing the 0 fracture surface and lengthwise showing fracture location and local necking behavior.
6.2 RESULTS .;O Tensile test properties of YS, UTS, RA, uniform elongation (UE) and total elongation (TE) are presented in Table 6-1. Shown in Figure 6-1 is a stress-strain curve for a 550*F base metal specimen typical of the .O stress-strain characteristics of all the specimens tested. Shown graphically in Figures 6-2 and 6-3 are the data in Table 6-1. Photographs of fracture surfaces and necking behavior are given in Figures 6-4, 6-5 and 6-6 for base, weld and HAZ specimens respectively. The base, weld, and HAZ !O materials generally follow the trend of decreasing properties with increasing temperature.
- O d
10 6-2
.O
. _ - _ _ - _ _ _ . . , _ , , ,_ .- , . - - - _ . - --.---,y -+ _ _y. _..-----_y, _, . - ,-.
O , 6.3 IRRADIATED VERSUS UNIRRADIATED TENSILE PROPERTIES O Unitradiated tensile test data were recovered from QA records for the surveillance specimen plate (G-3414-2) and the beltline weld (1-233). Data for 0.505 inch diameter gage tensile specimens from the fabrication test
- O program were used to get unirradiated room temperature YS, UTS, RA and TE properties. These are compared in Table 6-2 to the irradiated base metal and weld metal specimen RT data to determine the degree of irradiation effect. The trends of increasing YS and UTS and of decreasing TE and RA O are characteristic of irradiation embrittlement. The weld metal shows a larger effect, which is in agreement with the larger RT shift predicted ET because of the weld's higher Cu content.
O O O 1 O O O i T 6-3 0
i l l () i i Table 6-1 O TENSILE TEST RESULTS FOR IRRADIATED RPV MATERIALS Test Yield Ultimate Uniform Total Reduction C)' Specimen Temp Strength Strength Elongation Elongation of Area Number Material (*F) (ksi) (ksi) (%) (%) (%) SCL Base 76 71.4. 93.6 10.0 20.6 68.7 C) SCT Base 185 68.5 88.7 9.1 20.7 72.5 SCM Base 550 65.1 89.1 8.9 17.4 65.9 SDL Weld 76 88.6 105.0 9.4 18.5 64.5 {) SDM Weld 550 76.3 96.2 8.5 14.4 44.7 SED HAZ 76 77.2 99.6 7.7 17.3 68.2 SEB ERAZ 185 73.2 94.0 6.5 16.3 69.1 C) SEM HAZ 550 74.4 98.0 8.2 13.9 54.8 o O O O 6-4 O
0_ Table 6-2 O COMPARISON OF UNIRRADIATED AND IRRADIATED TENSILE PROPERTIES AT ROOM TEMPERATURE Yield Ultimate Total Reduction Strength Strength Elongation of Area (ksi) (ksi) (%) (%) h Base (G-3414-2): Unirradiated
- 67.7 89.3 27.0 69.8 O b Irradiated 71.4 93.6 20.6 68.6 Difference ' 5.2% 4.6% -31.1% -1.8%
~,
d Weld (1-233): Unirradiated
- 72.9 85.4 26.0 68.5
') Irradiated b 88.6 105.0 18.5 64.5 Difference " 17.8% 18.7% -40.5% -6.2% D
?)
8 Specimens have 0.505 inch gage diameter. Specimens have 0.250 inch gage diameter. Difference = [(Irradiated - Unirradiated)/ Irradiated]
- 100%
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B-Hee 5DM t 550"F Figure 6-3. Fracture 1.oeat ion , !;eck ing liehavio r , and l'rac t ure Appearance for Irradiated Weld :letal Tensile Specimens
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O SE D 760F l b 4 b Q SES 1850F
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Aw. _ . :,?gys V. N'" ' j' ,1 SE M 55@F Figure 6-6 Fracrure Locat ion, Neck ing Behavior, and Fract ure Appearance for I r radiated IIAZ !!etal Tens ile Specimens l b
- 7. DEVELOPMENT OF OPERATING LIMITS CURVES
7.1 BACKGROUND
Operating limits for pressure and temperature are required for three categories of operation: (a) hydrostatic pressure tests and leak tests, referred to as Curve A; (b) non-nuclear heatup/cooldown and low-level physics tests, referred to as Curve B; and (c) core critical operation, referred to as Curve C. There are three vessel regions that affect the operating limits: the closure flange region, the core beltline region, and the remainder of the vessel, or non-beltline regions. The closure flange region limits are controlling at lower pressures primarily because of Reference 1 requirements. The non-beltline and beltline region operating limits are evaluated according to procedures in References 1 and 2, with the beltline region minimum temperature limits increasing as the vessel is irradiated, m 7.2 NON-BELTLINE REGIONS Non-beltline regions are those locations that receive too little fluence to cause any RT increase. Non-beltline components include the NDT nozzles, the closure flanges, some shell plates, top and bottom head plates and the control rod drive (CRD) penetrations. Detailed stress analyses of the non-beltline components were performed for the BWR/6. The analyses took into account all mechanical loadings and thermal transients anticipated. Detailed stresses were used according to Reference 2 to develop plots of allowable pressure (P) versus temperature relative to the reference temperature (T - RTNDT). These results were applied to the FitzPatrick vessel components, since the geometries are not significantly different from BWR/6 configurations, and the mechanical and thermal loadings anticipated are comparable. N 7-1
O The non-beltline region results were established by adding the O highest Rryg, for the non-beit11ne discontinuities to the F versus (T - RTNDT) curves for the most limiting BWR/6 components. The CRD penetration and feedwater nozzle results bound the results for all other non-beltiine BWR/6 components. Shown in Figures 7-1 through 7-3 are the FitzPatrick O unique caicu1atea non-beltiine operating limits for Curves A, B, and C, respectively. Curve A limiting values are set by the BWR/6 CRD penetration results in conjunction with the 30*F RT f the recircu1ation inlet ET nozzle. The BWR/6 feedwater nozzle analysis, combined with the O recirculation inlet RT NDT f '3 0
- F , establishes the limits for Curves B and C.
7.3 CORE BELTLINE REGION O The pressure-tempefature (P-T) limits for the unirradiated beltline region are shown in Figures 7-1 through 7-3. As the beltline fluence increases during operation, these curves shift to the right by an amount ) discussed in Subsection 7.6. Eventually, the beltline curves shift to become more limiting than the non-beltline curves. The stress intensity factors calculated for the beltline region according to Reference 2 pro <.edures are based on a combination of pressure and thermal stresses. 3 The pressure stresses were calculated using thin-walled cylinder equations. Thermal stresses were calculated assuming the through-wall temperature distribution of a flat plate subjected to a 100*F/hr thermal gradient. The initial RT NDT f 24*F for the beltline was used to adjust the U (T - RTNDT) values from Figure G-2210-1 of Reference 2. 7.4 CLOSURE FLANGE REGION b Reference 1 sets several minimum requirements for pressurs and temperature in the closure flange region in addition to those outlined in Reference 2. In some cases, the results of analysis for other regions exceed these requirements and they do not affect the shape of the P-T D curves. However, some closure flange requirements from Reference 1 do impact the curves. In addition, Cencral Electric recommends 60*F margin on the required bolt preload temperature. ) 7-2
a As stated in Faragraph C-2222(c) of Reference 2. for application of C full bolt preload and reactor pressure up to 20% of hydrostatic test pressure, the RPV metal temperature must be at RT # E##^ '#* NDT practice is to require (RT DT + 60*F) for bolt preload, for two reasons: o o a. The original ASME Code of construction requires (RT.g + 60*F); and
- b. The highest stressed region during boltup is the closure flange N
region, and the flaw size assumed in that reglen (0.24 inches) is less than 1/4 T. This flaw size is detectable using ultrasonic testing (l"') techniques. In fact, References 15 and 16 report that a flaw in the closure flange region of 0.09 inch 3 can be reliably detected using UT. (RT ET + 60*F) is not a current ASME Code requirement; it provides extra margin for Curves A and B. However, (RT.DT + 60*F) is 4. uirement for
-J Curve C, as described in paragraph IV.A.3 of Reference 1.
Reference 1, paragraph IV.A.2, sets temperature minimum requirerents for pressure above 20% hydrotest pressure. Curve A temperature must be r) no less than (RT ET + 90*F) and Curve B temperature no less than (RT g + 120*F). The Curve A requirement causes a 30*F shift at 20% hydrotest pressure (312 psig) as shown in Figure 7-1. The Cu rve B requirecent has no impact on Figure 7-2 because the analytical results for
-) the feedwater nozzle require that temperature be greater than (RTNDT + 120*F) at 312 psig.
7.5 CORE CRITICAL OPERATION REQUIREMENTS OF 10CFR50, APPENDIN G Curve C, the core operation cu rve shown in Figure 7-3 is generated from Figure 7-2, accounting for the requirements of Reference 1, paragraph IV.A.3. Essentially paragraph IV.A.3 requires that core I critical P-T limits be 40*F above any Curve A or B limits. Curve B is more limiting than Curve A, so Curve C is Curve B plus 40*F. The (RT NDT 7-3
O minimum permissible temperature mentioned in Subsection 7.4 for O Curye C is an exception for 3wRs, a11owing critica1 operation at temperatures below the hydrostatic test temperature. 7.6 EVALUATION OF RADIATION EFFECTS O The shift in fracture toughness properties in the beltline materials is a function of neutron fluence and the presence of certain elements, such as copper (Cu) and phosphorus (P). The specific relationship from n v Reference 5 is: SHIFT (*F) = [40 + 1000(%Cu-0.08) + 5000(%P-0.008)]*(f/10 ') (7-1) O where:
%Cu = wt % of Cu present,
%P = wt % of P present, O f = fluence (n/cm 2) at selected EFPY.
The limiting beltline plate and weld are determined based on the Cu-P content and initial RT f the materials. Calculations based on the ET information in Tables 3-1 and 3-2 show the following: Limiting plate: 0.13% Cu, 0.015% P. Initial RT ~ NDT Limiting weld: 0.31% Cu, 0.015% P Initial RT ~~ NDT O Surveillance plate: 0.13% Cu, 0.011% P This material information is used to evaluate irradiation shift versus fluence. 3 7-4 )
'~)
7.6.1 Measured Versus Predicted Surveillance Shift n J Table 5-3 presents a measured shif t for the base metal of 23*F. No measured shift is available for the weld metal. The predicted shift of the surveillance plate, calculated according to Equation 7-1, assumes 0.13% Cu,
) 0.011% P and an upper bound fluence of:
17 f - (2.6x10 n/cm2 for capsule)(1.25 uncertainty) = 3.25x10 n/cm . I) The predicted shift is 19'F, versus the measured shift of 23*F. 7.6.2 Modification of the Shift Relationship
$) Since the measured shift exceeds the predicted shift, using Reference 5 methods to predict the limiting beltline plate shift may be non-cons e rva tive , and therefore is modified. Using Reference 5, the shift calculated is proportional to the material characteristics and to the $) square root of the fluence. Assuming that the fluence relationship is correct means that the coefficient representing the materials in Equation 7-1 must be increased by the factor (23/19) or 1.21.
'3 7.6.3 Radiation Shift Versus EFPY Equation 7-1 can be simplified and expressed as a function of EFPY for the base metal and weld metal. Subsection 4.3 concludes that the EOL 3 (32 EFPY) 1/4 T fluence is 1.7x10 n/cm2 Therefore, in te rms of EFPY, the fluence is f = 5.31x10 6
- EFPY (7-2) fD Equation 7-2 is used in Equation 7-1 with the appropriate Cu, P values.
For the weld metal: SHIFT w = 22.23 * (EFPY)
. (7-3) 7-5 3)
For the base metal: (including 1.21 factor) wr' SHIFT3
= 11.02 * (EFPY)'. (7-4)
The adjusted reference temperature (ART) is defined as the 3 initial RTg plus the irradiation shift. The initial RT NDT values of the limiting plate and weld materials are 24*F and -22*F, respectively. Figure 7-4 shows the ART for each material based on these initial RT NDT values and the shifts of Equations 7-3 and 7-4. As shown in the figure,
, the plate is initially limiting, because of its higher initial RT .
However, the larger shift associated with the weld, because of its higher Cu content, causes the weld ART to exceed the plate ART at about 17 EFPY. The ART values of Figure 7-4 are used to develop the operating limits
- 1) curves. The higher of the plate or veld ART is used for a given EFPY.
7.6.4 End-Of-Life Conditions Paragraph IV.B of Reference I sets limits on the ART and on the upper shelf energy (USE) of the beltline caterials. The ART must be less than 200*F, and the USE must be above 50 ft-lb. Based on Figure 7-4, the ART values at 32 EFPY of 104*F for the weld and 86*F for the plate are j acceptable. Calculations of USE, using Reference 5, are summarized in Table 7-1. The equivalent transverse USE of the plate caterial is taken as 65% of the % longitudinal USE, according to Reference 7. The veld metal USE is not adjusted because weld metal has no orientation effect. Surveillance results for the base metal in Table 5-3 show USE decrease of 6%. The calculated value, using Reference 5, is 10%, so Reference 5 appears to be % conservative. Original fabrication testing done by Combustion Engineering developed Charpy data for the beltline pintes up to 160*F. The values at 160*F are averaged to estimate USE for the unirradiated beltline plates. The veld USE is established for an intermediate fluence by the surveillance % test results. This intermediate value is used to predict the original unirradiated USE, and an EOL value is then calculated. The minimum E01 plate and weld USE values are estinated as 58 ft-lb and 72 ft-lb, % 7-6
O respectively, which are above the minimum limit. Therefore, irradiation O effects are not severe enough to necessitate RPV annealing before 32 EFPY. 7.7 OPERATING LIMITS CURVES VALID TO 16 EFPY O The ART selected for the core beltline curves depends on the amount of operation for which the curves will be valid. Sixteen EFPY was selected because Reference 4 recommends withdrawal of the second surveillance capsule at 15 EFPY. The beltline ART estimated with Figure 7-4 is 68'F, O based on the plate material. Adjusting the unitradiated beltline curves in Figures 7-1 through 7-3, with their initial RT of 24*F, and considering the non-beltline curve s , gives the operating limits valid to 16 EFPY, as shown on Curves A, B and C in Figure 7-5, 7-6 and 7-7, respectively. The O values plotted on Figures 7-5 through 7-7 are tabulated in Table 7-2 for Curve A and Table 7-3 for Curves B and C. 7.8 REACTOR OPERATIuN VERSUS OPERATING LIMITS D For most reactor operating conditions, pressure and temperature are at saturation conditions, which are in the operating zone - of the limits curves. The most severe unplanned transient is an upset condition 3 -consisting of several transients which result in a SCRAM. The worst combination of pressure and temperature is 1180 psig with temperatures in the lower head at 250*F. At the same time, the steam space coolant temperature is still nearly 550*F. Steam space coolant temperature is used ') to identify the appropriate curve to be applied. In this case, the core is not critical and, according to the steam space coolant temperature, there is no significant cooldown occurring, so the hydrostatic pressure curve applies (Curve A). As seen for Curve A in Figure 7-5, at 1180 psi the % minimum transient temperature in the vessel of 250*F lies in the safe operating zone. Therefore, violation of the operating limits curves is only a concern in cases where operator interaction occurs, such as hydrostatic pressure testing and initiation of criticality, r) 7-7 ')
Table 7-1 ESTIMATE OF UPPER SilELF ENERGY FOR BELTLINE MATERIALS Upper Shelf (ft-lb)
'_ ; Longitudinal / Transverse Identification % Cu f=0 f=3.3x10 f=1.7x10 Lower Shell:
G-3415-1R 0.11 132" 120 115/75 G-3415-2 0.14 127* 113 108/70 G-3415-3 0.13 119" 107 102/66
) Low-Int Shell:
G-3413-7 0.12 103a ,b 93 89/58 G-3414-1 0.18 127* 112 104/68 G-3414-2 0.13 130* 117 112/73 Longitudinal Weld: 1-233 0.31 104 82" 72/72 a
.)
- USE values taken from test data. Other USE values are calculated using Reference 5.
Based on Charpy data at only 110*F. Actual USE should be somewhat higher. s / 7-8
/
Table 7-2 PRESSURE-TEMPERATURE VALUES FOR FIGURE 7-5, CURVE A Pressure (psi) Temperature (*F) Remarks 0 90 Boltup Temperature o 312 90 'J 120 RT +90 per 10CFR50 App. G 312 755 120 BeklIinebecomeslimiting 760 121 770 122.5 780 124 m 790 126 'J 800 127.5 810 129 820 130.5 830 132 840 133.5 3 850 135 860 136.5 870 138 880 139 890 140.5 900 142 s 910 143 920 144.5 930 145.5 940 147 950 148 960 149 s 970 150.5 980 151.5 990 152.5 1000 153.5 Pressure increment of 20 psi 1020 156 1040 158 g 1060 162.5 1080 164.5 1100 166.5
20 168.5 1140 170 1160 172 g 1180 173.5 1200 175 1220 177 1240 178.5 1260 180 1280 181.5 g 1300 183 1320 184.5 1340 186 1360 187 1380 188.5 1400 190 7-9
l
.a Table 7-3 {
PRESSURE-TEMPERATURE VALUES FOR FIGURE 7-6 (CURVE B) AND 7-7 (CURVE C) Pressure Curve B Curve C (psi) Temp. (*F) Temp. (*F) Remarks 0 90 90 Boltup Temperature 57 90 FW Nozzle limiting Curve C 60 94 70 102 80 110 90 117 100 124
]' 110 129 112 90 130 FW Nozzle limiting Curve B 120 94 134 130 99 139 140 104 144
'N 150 108 148 160 112 152 170 115.5 155.5 180 119 159 190 122 162 200 125 165
~)
210 127.5 167.5 220 130 170 230 132.5 172.5 240 135 175 250 137 177 260 139 179
; 270 141 181 280 143 183 290 145 185 300 147 187 Increcent pressure 20 psi 320 150 190 340 153 193
~1 360 157 197 380 160 200 400 162.5 202.5 420 165 205 440 168 208 460 170 210 'N 480 172 212 500 174.5 214.5 520 176.5 216.5 540 178 218 560 180 220 580 182 222 s
,.v 7-10
dD Table 7-3 (continued) Pressure Curve B Curve C (psi) Temp. (*F) Temp. (*F) Remarks 600 184 224 Increment pressure 50 psi 650 188 228
; 700 192 232 ~'
750 194 234 800 196 236 850 198 238 900 200 240 950 202 242
'^> 1000 204 244 1050 205.5 245.5 1100 207 247 1150 209 249 1200 211 251 1250 212.5 252.5 Beltline becomes limiting
'l 1300 215.5 255.5 1350 218.5 258.5 1400 221 261
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m l 10CF R50. APPENDIX G. IV A 2 l REQUIREVENT (RTNOT*3M3 1 4w - 1 I
") i ~
r l l 200 - GE RECOMMENDED BOLT i PRELOAD TEMPER ATURE OF ]) l 90cF (INCLUDE 5 600F M ARG6N) l I l I i 1 o 100 200 300 0 J MINIMUM VESSE L ME T AL TE MPER ATURE (OF ) F'gure 7-1. Components of Operating Limits Curve for Pressure Tests (Curve A) for FitzPatrick 7-12 ]
1400 0 1400 - h UNIRR ADI ATE D SELTLINE REGION (R TNot = 240FI f 1200 - 0 -
,a -
l f O l O 5
== -
I 8 E D E a E a00 - f l 3 I g NON BELTLINF REGION tRTNOT= 300Fi 10CFR50. APPENDIX G. IV. A.2 REQUIREMENT tRT l t Not = 30DF)X. _ . _ _ 3 I I l
/
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O 100 200 300 ] Figure 7-2. MINIMUM VESSEL METAL TEMPER ATURE toF) Components of Operating Limits Curve for Non-Nuclear } Heatup/Cooldown (Curve B) for FitzPatrick i
] 7-13
.l 0 1 00 1400 - UNIRR ADI ATE D SELTLINE REGION (RTwoy - 240F) f 1200 - 0 I I 1000 - O i s I l r 800 - 2 Y 3 N E. a 10CFR50. APPENDIX G
~ IV A.2 AND IV.A.3 REQUIRE MENTS (RTNDT
- 300F) I O I I
I 400 - l/ l c. r-------- 1 NON eELTLINE REGION l (RTsoy = WF) 200 - g 3 l I -
/
O 300 100 200 O b MINIMUM VESSEL METAL TEMPER ATURE (OF) Figure 7-3. Components of Operating Limits Curve for Core Critical Operation (Curve C) for FitzPatrick 7-14
llll.lllllllll . o 0 4 o 8 3 a 6 d 3 l e W o e 4 3 d n ak c ei 2 t r
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g - 1200 - ADJUSTED BE LTLINE. 1,4T F LAW, ART = 680F G a 1000 - 3 b c b SAFE I OPER ATING m h, 800 - R E GION a - i! w 5 m E a
^
, 600 - 1200F
/
3 a - 312 psig BOLT PHE LO AD 200 - TE MPER ATUH[ = 900F r
.']
[F LANGE REGION RTNOT = 30 F f f
's 0 J 0 100 200 300 MINIMUM VESSEL ME T AL TEMPER ATURE 10F )
Figure 7-5. Pressure Versus Minimun Temperature for flydrostatic Pressure Tests N for FitzPatrick u 7-16
2 3
,J 1600 V ALID To tS EFPY 1400 -
ADJUSTE D BE LTLINE, li4T F LAW ART
- 640F q 1200 -
v , 1000 - s. _? e w 1
, b
- sw -
E
$ $AFE w OPE R ATING f REGION , 600 -
J 3 NON BELTLINE FW NOZ2LE LIMITS, 1/4T F L AW RTNot*300F ) 1 BOLT PRELOAD TEMPER ATURE
- 90nr F LANGE FtEGION RTNOT
- 300F f I Q O ,w a 3w WINIMUM VESSE L MET AL TEMPER ATURE (OF )
Figure 7-6. Pressure Versus Minimum Temperature for Non-Nuclear lleatup and ^3 Cooldown for FitzPatrick 7-17
O O 2000 V ALID TO IS EFPY C ] 1400 - ADNSTED BELTLINE, It4T F LAW, ART = 680F ) . 1200 - 3 iOm - 5 e b z EL D ?W ~ E E D w NON BELTLINE FW NOZZLE LIMITS E PLUS 40 F,1/4T F LAW ) 800 - RTNot
- WF r] 400 -
SAFE OPE R A TING RE GION
# ~
FLANGE REGtON RTNOT* WF. MINIMUM PERMIS$18LE TEMPER ATURE
- 900F PE R 10CF RSO APPEN0lx G
) 0 O 100 200 300 MINIMUM VESSEL ME TAL TEMPER ATURE ("F I Figure 7-7. Pressure Versus Minimum Temperature for Core Critical Operation ] for FitzPatrick 7-18
O
- 8. REFERENCES
)
- 1. " Fracture Toughness Requirements " Appendix C to Part 50 of Title 10 of the Code of Federal Regulations, July 1983 (24FR24008).
3 2. " Protection Against Non-Ductile Failure," Appendix G to Section III of the ASME Boiler & Pressure Vessel Code, Addenda to and including Winter 1984.
) 3. " Reactor Vessel Material Surveillance Progran Requirements,"
Appendix H to Part 50 of Title 10 of the Code of Federal Regulations, July 1983 (48FR24008). J 4 " Conducting Surveillance Tests for Light Water Cooled Nuclear Power Reactor Vessels," Annual Book of ASTM Standards, E185-82, July 1982.
- 5. " Effects of Residual Elements on Predicted Radiation Damage to
_i Reactor Vessel Materials," l'SNRC Regulatory Guide 1.99, Revision 1 April 1977.
- 6. Deleted.
m Y
- 7. " Fracture Toughness Requirements " USNRC Branch Technical Position KIEB 5-2, Revision 1, July 1981.
) 8. " Capsule Basket," Generr.1 Electric Drawing Il7C4349 Revision 1, October 1969.
- 9. Letter, R. A. Hillis of Combustion Engineering, Inc. to C. A. Barry of O General Electric Company, " Welding Material Certification Data for FitzPatrick Beltline," January 1986.
- 10. " Surveillance Test Program for Nine Mile Point III Feactor Vessel,"
) Combustion Engineering Specification (CE VPF 1980-234-1), May 1969.
8-1
O 4
- 11. Caine, T. A., " Hatch 1 RPV Surveillance Materials Testing and Fracture j Toughness Analysis," General Electric Company (NEDC-30997),
October 1985.
- 12. " Standard Methods for Notched Bar Impact Testing of Metallic O Materials," Annual Book of ASTM Standards, E23-82, March 1982.
- 13. " Standard Test Methods for Rockwell Hardness and Superficial Rockwell Hardness of Metallic Materials," Annual Book of ASTM Standards,
[) E18-79.
- 14. " Standard Methods of Tension Testing of Metallic Materials," Annual Book of ASTM Standards, E8-81.
- 15. "Citrasonic Examination for Cracks in the Top Head Flange," CBl Nuclear. Development Report 74-9047, December 1975.
?) 16. " Ultrasonic Examination for Cracks in the Shell Flange," CBI Nuclear, Development Report 74-9056, November 1975.
m
._.)
8-2
O APPENDIX A O CHARPY V-NOTCH FRACTURE SURFACE PHOTOGRAPHS Photographs of each Charpy specimen fracture surface were taken 3 to facilitate the determination of percent shear, and to comply with the requirements of ASTM E185-82. The pages following show the fracture surface photographs along with a summary of the Charpy test results for each specimen. The pictures are arranged by increasing 3 test temperature for each material, with the materials in the order of base, weld and HAZ. w.e g s. ~s, w. A-1/A-2 S
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