ML20235B066
| ML20235B066 | |
| Person / Time | |
|---|---|
| Site: | Davis Besse  |
| Issue date: | 01/20/1989 |
| From: | Boger B Office of Nuclear Reactor Regulation |
| To: | Sterling E ARIZONA PUBLIC SERVICE CO. (FORMERLY ARIZONA NUCLEAR, C-E OPERATING PLANTS OWNERS GROUP |
| Shared Package | |
| ML20235B069 | List: |
| References | |
| NUDOCS 8902090266 | |
| Download: ML20235B066 (362) | |
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SSb Mr. Edward Sterling, Chairman DISTRIBUTION JS 3g gg-Combustion Engineering Owners Group central F11es i c/o Arizona Public Service Company w/o enclosures: l P. O. Box 52034 BABoger BElliot l Phoenix, AZ 85072-2034 GClainas LShao ! AJMendiola AD RI RF 1
Dear Mr. Sterling:
,TChan ]
In our telecon of January 25, 1989, we briefly discussed the NRC staff's concerns on the effect of low-temperature, low flux embrittlement on the structural integrity of reactor vessel supports. The reviews to date have concluded that this issue does not pose an immediate concern to public safety, but it challenges the margins provided by the defense-in-depth concept. Accordingly, we would like to meet with representatives of the_CEOG (as well l i as the Westinghouse Owners Group) in early March 1989 to discuss embrittlement of reactor vessel supports. ; 1 To assist you in your preparation for the meeting, I have enclosed two documents for your review. One is the Oak Ridge National Laboratory Final Draft Report on this issue, and the other is the staff's evaluation. We will be in touch with you in the near future to establish a meeting date.- l Sincerely,
')nginal Slamed 8y:
IsruceA.Soger Bruce A. Boger, Assistant Director i for Region I Reactors Division of Reactor Projects I/II Office of Nuclear Reactor Regulation
Enclosures:
As stated 1 cc w/ enclosures: Mr. , lim Pfiefer Combustion Engineering, Inc. 1000 Prospect Hill Road Windsor, CT 06095-0500 Mr. Charles Brinkman , Combustion Engineering, Inc. 12300 Twinbrook Parkway l Suite 330 Rockville, MD 20852 t DRP: I n la oger:ah
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NUREG/CR-IXXX ORNL/TM-10966 Dist. Category RF FINAL DRAFT December 7,1988 Engineering Technology Division "a3l "d '
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'"F"l IMPACT OF RADIATION EMBRITTLEMENT ON INTEGRITY OF PRESSURE VESSEL SUPPORTS FOR TWO PWR PLANTS R. D. Cheverton W. E. Pennell j G. C. Robinson R. K. Hanstad*
- Metals and Ceramics Division Manuscript Completed -
Date of Issue - l NOTICE: This document contains information of a preliminary , nature. It is subject to revision or correction and therefore I does not represent a final report. Prepared for the , U.S. Nuclear Regulatory Commission Of fice of Nuclear Regulatory Research Washington, DC 20555 under Interagency Agreement DOE 1886-8011-9B NRC FIN No. B0119 Prepared by the OAK RIDGE NATIONAL LABORATORY Oak Ridge, Tennessee 37831 operated by MARTIN MARIETTA ENERGY SYSTEMS, INC. for the U.S. DEPARTMENT OF ENERGY I" under Contract No. DE-AC05-840R21400 L, . i 9
. a.
i-l CONTENTS i Page ACKNOWLEDGMENTS ................................................. iv 1 i
- 1. INTRODUCTION AND
SUMMARY
.............................~....... 1 l Ref e rences .................................................. 6 -
l
- 2. HFIR VESSEL SURVEILLANCE DATA ............................... 11 ;
References .................................................. 14'
- 3. APPLICATION OF HFIR DATA TO VESSEL SUPPORT EVALUATION ....... 22 Re,f e rences .................................................. 25
- 4. SURVEY OF SUPPORT DESIGN FEATURES ........................... 38 1
4.1 Introduction ........................................... 38 l l 4.2 Categorization ......................................... 39 )i Ref e rences .................................................. 44 ) a
- 5. SELECTION OF LWR PLANTS FOR SPECIFIC-PLANT. ANALYSIS ......... 60~ l
- 6. BRITTLE FRACTURE EVALUATION OF TROJAN NUCLEAR PLANT -l REACTOR PRESSURE VESSEL SUPPORTS ...... ..................... 63 ;
i 6.1 Introduction and Summary ............................... 63 6.1 Scope and Objective .................................... 66 6.3 Loading and Operating Environment ...................... 70 6.3.1 Normal plus upset loading ....................... 70 , 6.3.2 Faulted conditions .............................. 70 1 6.3.3 Cyclic loading .................................. 70 6.3.4 Operating environment ........................... 70 6.4 Support Beam Analysis .................................. 71 6.4.1 Analytical model ................................ 71 6.4.2 Model element properties ........................ 72 6.4.2.1 Beam (solid section) ................... 72
^
6.4.2.2 Beam (section with 4 in. diameter i holes) ................................. 72 l 6.4.2.3 Concrete foundation .................... 72 l 6.4.2.4 Flange of the f ront support l pedestal ............................... 74 6.4.2 5 Front support pedestal web ............. 74 6.4.2.6 Flange section of the rear support ..... 74 6.4.2.7 Web section of the rear support ........ 74 6.4.2.8 Foundation modelling sensitivity ass e's sment ............................. 75 6.4.3 Bean analysis input ............................. 75
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\ .E.*& *. -l 6.4.4 Beam analysis output ............................' 76 l 6.4.4.1 Effectiveness of the inner concrete support ....................... 77 6.4.4.2 -Beam bending moments ................... 78 6.4.4.3 Support stiffness ...................... 79 6.5 Material Fracture Toughness ............................ 80 6.5.1 Strain-rate correction .......................... 81 I 6.5.2 Data correlation and extrapolation .............. 83 l 6.5.3 Plane strain-plane stress transition- ............ 86 6.5.4 Fracture toughness design curves ................ 87 6.5.5 A36 nil ductility temperature variations ........ 89 6.5.6 Supporting analysis ............................. 89 6.5.7 Reactor support / bridge structure c o mp a r i s o n : . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 90 6.5.8 Fracture toughness data evolution ............... 90 6.6 Radiation-Induced NDTT Shif t- ........................... 91 6.7 Critical Flaw Depth Analysis ........................... 93 6.7.1 Cavity linear interface ......................... 94 6.7.2 Maximum bending moment location (Node 4) ........ 96 6.7.3 Beam flange grout-hole location . . . . . . . . . . . . . . . . . 98 6.7.3.1 Flaw-tip stress-intensity-f actor equations .............................. 98 6.7.3.2 Flaw tip stress intensity factors ...... 100 6.7.3.3 Lifetime critical flaw depth analyses ............................... 100 l 6.7.3.4 The impact of NDT variations' ........... 102 l 6.8 Evaluation of Results .................................. 103 I l
6.8.1 Flaw growth due to cyclic loading ............... 103 6.8.2 Interface with the reactor cavity liner ......... 104
)
l l 6.8.3 Haximum bending moment location (Node 4) ........ 105 6.8.4 Beam top flange grout hole location ............. 105 6.8.5 Compatibility check on the stress-intensity factor equations ................................ 106 _ 6.8.6 Load sensitivity comparison ..................... 108 6.8.7 Residual stress effects ......................... 110 6.9 Conclusions and Recommendations ........................ 118 References ............................................. 120
- 7. BRITTLE FRACTURE EVALUATION OF TURKEY POINT UNIT 3 REACTOR PRESSURE VESSEL SUPPORTS ............................ 204 7.1 Introduction and Summary ............................... 204 7.2 Description of Supports ................................ 206 ,
.I 7.3 Materials .............................................. .207 1 7.4 Material Fracture - Toughness Properties ............... 207 ;
7.5 Thermal and Radiation Environment ...................... 210 l
iii. i. Page 7.6 Estimates of ANDTT ..................................... 212 7.7 Support. Loadings and Stress Analysis .................... 212 7.8 Stress Intensity Factor, K I, Evaluation ................ 215 7.9 Discussion of Results and Conclusions .................. 217 . I References .................................................. 220
- 8. CONCLUSIONS AND RECOMMENDATIONS ............................. 288 8.1 Conclusions ............................................ 288 8.2 Re commendations .......................................'. 290 Appendix 1: TROJAN SUPPORT BEAM ANALYSIS INPUT AND OUTPUT l
FILES .............................................. 291 j Appendix 2: TROJAN CRITICAL FLAW SIZE ANALYSIS FOR THE BEAM TOP FLANGE AT THE INTERSECTION WITH THE REACTOR CAVITY LINER- ....................................... 301 Appendix 3: TROJAN CRITICAL FLAW SIZE ANALYSIS FOR THE BEAM TOP FLANGE AT THE PEAK BENDING MOMENT LOCATION (NODE 4) ........................................... 313 Appendix 4: TROJAN CRITICAL FLAW SIZE ANALYSIS FOR THE BEAM TOP FLANGE AT THE LOCATION OF THE 4 INCH DIAMETER ', GROUT HOLES ........................................ 335 (. Appendix 5: TROJAN BEAM FOUNDATION MODELLING SENSITIVITY ANALYSIS ........................................... 341 Appendix 6: CALCULATION OF K I VALUES FOR CRACKS EMANATING FROM A HOLE IN THE PLATE (TROJAN) ........ 353 l I t
l l t.. ACKNOWLEDGEMENTS l The authors wish to acknowlege contributions made by J. W. Br'yson .i 1 (fracture analysis), W. R. Corwin (HSST program manager), K. Farrell I (radiation embrittlement), S. K. Iskander (review), F. B. Kam (neutron fluxes), J. C. Merkle (fracture mechanics and review), and C. E. Pugh (PVT section head), all at Oak Ridge National Laboratory, and T. J. Criesbach (Electric Power Research Institute), J. R. Hawthorne (Materi-als Engineering Associates) and C. R. Odette (University of California, Santa Barbara). I i l i e
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( IMPACT OF RADIATION EMBRITTLEMENT 6 I ON INTEGRITY OF PRESSURE VESSEL i SUPPORTS FOR TWO PWR PLANTS l
- 1. INTRODUCTION AND
SUMMARY
j Structural supports for most pressurized-water-reactor (PWR) pres- ]
- \
sure vessels are located in the cavity between the vessel and the bio- .] I logical shield (Fig. 1.1). Within the cavity the fast neutron flux ($) l for energies (E) > 1.0 MeV is s 2 x 109 n/cm2.s, and temperatures are 1
<150*F. The corresponding calculated increase in the nil ductility transition temperature (NDTT) by 32 effective full power years (EFPY),
based on the radiation embrittlement data available from materials test-ing reactors (MTRs) before 1987 (Refs. 1,2), is quite small, if the dif-
, ference in the MTR and the PWR cavity fast neutron energy spectra :is neglected.
Early in 1978 it became apparent that . the fast neutron spectrum above 0.1 MeV was much softer in the PWR cavity than in the MTRs (the result of inelastic scattering in the PWR vessel wall), and thus corre-lating the MTR embrittlement data with fast neutron fluence (4) for E > 1.0 MeV resulted in an underestine.tien of the increase in the nil ductility transition temperature (ANDTT) for supports in the cavity. As a part of the more recent study discussed herein, the MTR data were cor-related with displacements per atom (dpa) for E > 0.1 MeV, and ANDTT values of 570*F were calculated for 32 EFPY for supports located in the cavity at midheight of the core.
2 Several studies pertaining to radiation damage of PWR vessel sup-ports were conducted between 1978 and 1987 (Refs. 3-6), and during this period, presumably there was no reason to believe that low-temperature (<200*F) HTR embrittlement data, correlated with dpa (E > 0.1 MeV), were not appropriate for evaluating embrittlement of PWR vessel supports. However, late in 1986, de.ta from the High Flux Isotope Reactor (HFIR)? vessel surveillance program s ,s indicated that the embrittlement rates of the s'everal vessel materials (A212-B, A350-LF3, A105-II) were substan-tially greater than anticipated on. the basir of MTR data.9 Further evaluation of the HFIR data suggested that a fluence-rate effect was responsible for the apparent discrepancy because the fast neutron flux in the MTR that provided the design datal was -105 times that in the HFIR vessel, while the small differences in the fast neutron energy ( spectra were accounted for by correlating the data with dpa (E > 0.1 HeV), and the irradiation temperatures were nearly the same (120-200'F). As a result of this new information, the Nuclear Regulatory Commission (NRC) requested that the Oak Ridge National Laboratory (ORNL) evaluate the impact of the apparent embrittlement rate effect on the integrity of light-water reactor (LWR) vessel supports. The purpose of this first report is to provide a preliminary indi-cation of whether the integrity of reactor vessel supports is likely to be challenged by radiation-induced embrittlement before 32 EFPY. Because of the diversity in support designs, specific plant evaluations were undertaken; and because of the urgency associated with the prelimi-nary report, only readily available data for these plants were used. This and other factors have resulted in some inconsistencies in the
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)
evaluations of the plants considered. Even so, the report satisfies its r intended purpose. l
)
The scope of the ORNL evaluation included I
- 1. correlation of the HFIR data for application to the evaluation of LWF. vessel supports, !
- 2. a survey and cursory evaluation of all U.S. LWR vessel support designs,
- 3. selection of two plants in accordance with established criteria for l
specific plant evaluation, and l
- 4. a specific plant evaluation of both plants to determine critical flaw sizes for their vessel supports. i
. The two plants selected for specific plant evaluation were Trojan (Portland General Electric) and Turkey Point Unit 3 (Florida Power and
(' Light), both of which are PWR plants and have vessel supports similar to that shown in Fig. 1.1. Westinghouse was the nuclear-steam-supply-sys-tem (NSSS) designer, and Bechtel was the architect engineer for both plants. The utilities and their contractors have been very cooperative in providing, where possible, design data required for the ORNL study. Over the course of several months, two sets of radiation damage trend curves (ANDTT vs dpa), based on the HFIR vessel surveillance data, were developed, and 32-EFPY ANDTT values were calculated for " typical" Ceneral Electric (CE), Babcock and Wilcox (B&W), Westinghouse (W) and Combustion Engineering (CE) plants, assuming that a critical portion of a support existed in the cavity at midheight of the core. The results, presented in Table 1.1, indicate much larger shifts in NDTT based on the HFIR data than on the MTR data.
4 ( Many of the vessel supports are not located at midheight of the core and thus experience smaller shifts in NDTT than those indicated in Table 1.1. For instance, all but one of the CE boiling water reactor (BWR) vessels and all but one of the B&W PWR vessels are supported by a skirt at the bottom of the vessel where the fluxes are much less. How-ever, -25% of the PWR vessel supports are exposed to the peak flux, and i 1 many others are exposed to fluxes within a factor of one-half of the peak. The concern over radiation embrittlement is that it increases the potential for propagation of flaws that might exist in the support structures. In this study the potential for propagation of flaws was evaluated using linear elastic fracture mechanics, which requires know-ledge of the stresses in the structure and the fracture toughness of the material. The output of the fracture-mechanics analysis for this study is the critical flaw size, that is, the size of the smallest flaw that will propagate under a given set of assumed conditions and will result in failure of the support. If the critical flaw size is small enough that the critical flaw is likely to exist, then the frequency of failure is equal to the frequency of application of the assumed load. A deter-mination of the probability of the existence of flaws was not, however, included in the scope of this study. Loading conditions were provided by the utilities and included large- and small-break loss-of-coolant accidents (LBLOCA and SBLOCA), seismic loading, thermal loading, and dead-weight loading. Dynamic loading conditions were available for Trojan but were not available for Turkey Point. k 1 1
9 J
- k. A number of uncertainties exist in the analysis, and those judged to be the most significant are the radiation damage trend curve deduced from the HFIR surveillance data, the fracture toughness of the material, including the initial NDTT, and the norm 41 operating temperature of the structure at the critical location (T). These uncertainties result in a wide variation in calculated critical flaw size. The critical flaw size is, of course, also dependent on load and the location and type of flaw. Critical flaw sizes (depth / surface length) were calculated for a e
reasonable range of conditions, and 32 EFPY "best estimate" values for the most sensitive location and type of flaw and for the most severe credible loading conditions are 0.4/2.5 in. for Trojan and 0.3/0.6 in. for Turkey Point (Table 1.2). Corresponding values for late 1988 are 0.9/16 and 0.3/0.6 in., respectively. i For both plants the flaw location resulting in the smallest criti-cal flaw depth was on the upper flange of the horizontal cantilever beam at a point within the concrete biological shield (Fig. 1.1). The best-estimate values . of T - NDTT at these locations are -13 and -25'F for 32 EFPY and -12 and 28'F for late 1988 for Trojan and Turkey Point, respectively.
~
The propagation of flaws by low-cycle fatigue was calculated to be negligible. Thus, if corrosion is not a viable means of growing flawr, 1 to critical size or larger, flaws of critical size or larger would have , to be present at the time of fabrication. As already mentioned, this report does not address the question regarding the probability of exis-tence of flaws. (
6 l I
- t. Details regarding the HFIR vessel' surveillance date, development of j l
the corresponding embritt knent trend curves, a survey and cursory eval- .] untion of the LWR s up.M.7 7 designs, selection of two LWR plants for specific plant evaluation of the supports, and the specific plant l analyses are included in the following sections of this report. ! References
- 1. J. R. Hawthorne, Studies of Radiation Effects and Recovery of Notch Duetility of Pressure Vessel steels, British Nuclear Energy Confer-ence, Iron and Steel Institute,. London, England, November 30, 1960.
1
- 2. L. E. Steele, J. R. Hawthorne, C. 2. Serpan, Jr. , E. P. Klier, and l U. E. Watson, Irradiated Materials Evaluation and Reactor Pressure l Vessel Surveillance for the Army Nuclear Power Program, MRL Memotan-dum Report 1644, September 1, 1965.
- 3. C. A. Knorovski, R. D. ' Krieg, C. C. Allen, Jr., Practure Toughness of PWR Components Supports, Sandia National Laboratory, NUREG/CR-3009 (SAND 78-2347), February 1983.
l 4. Requirements and Guidelines for Evaluating Component Support Mate-rials Under Unresolved Safety Issue A-12, Electric Power Research Institute, EPRI NP-3528, June 1984
- 5. W. C. Hopkins and W. L. Crove, A Study of the Embrittlement of Reac-tor Vessel Steel Supports, Reactor ~ Dosimetry, Volume 2, J. P. Centhon and H. Rottger (eds.), Dordrecht, Netherlands D. Reidel Publishing Company, 1985, pp. 621-628.
- 6. W. C. Hopkins, Reactor Pressure Vessel Supports for Pressurized
, Water Reactors and Boiling Water Reactors, Residual Life Assessment of Major Light Water Reactor Components ~ Overview, Volume 1, Idaho National Engineering Laboratory, NUREC/CR-4731 (ECC-2469), Vol. 1, June 1987.
- 7. R. D. Cheverton and T. M. Sims, RFIR Core Nuclear Design, ORNL-4621, Union Carbide Corp., Nuclear Div., Oak Ridge National Laboratory, July 1971.
- 8. J. R. McWherter, R. E. Schappel, and J. R. McGuf fey, NFIR Pressure Vessel and Structural Components Naterial Survelliance Program, ORNL/TM-1372, Martin Marietta Energy Systems, Inc., Oak Ridge National Laboratory, January 1966.
7 l
- 9. R. D. Cheverton, J. C. Merkle and R. K. Nanstad, eds., Evaluation of HTIR Pressure-Vessel Integrity Considering Radiation Rabrittlement, ORNL/TH-10444, Martin Marietta Energy Systems, Inc., Oak Ridge l
National Laboratory, April 1988. e 1 (' l l l 1
, v-4 4 8 l I. , Table 1.1. Vessel support ANDTT valves corresponding to 32 EFPY and midheight of core
(" typical" LWR plants) I pggg ANDTT, 'F 6 dpa rate designer i (type (E > 1 MeV) (E > 0.1 MeV) dpa MTR HFIR data n/cm2.s s1 data l reactor) Aa Ba j CE (BWR)b 2.9 x 107 5.8 x 10-14 5.8 x 10-5 0 d d g B&W (PWR)C 2.0 x los 6.1 x 10-13 6.1 x 10-4 20 180 130 W (PWR) 5.9 x 10s 3,9 x 10-12 3.9 x 10-3 50 240 220 CE (PWR) 1.8 x 109 4.5 x 10-12 4.5 x 10-3 70 250 220 ( aANDTT vs dpa (E > 0.1 MeV) correlations A & B. i b Boiling water reactor. l Pressurized water reactor. d ANDTT not estimated for dpa < 10-d. l l
. _ . q
> 4-9 !
l f Table 1.2. Summary of best-estimate minimum critical flaw sizes and values of T-NDTT for Trojan and Turkey Point vessel supports (cantilever beam) Critical flaw size T-NDTT (depth x surface length) Plant EFPY
- D * '
LBLOCA SBLOCA (ggE)C Troja,n 7.5d -12 -12 0.9 x 16.0 1.1 x 16.0 l 32 -13 -118 0.4 x 2.5 1.2.x 2.5 l Turkey Point 11.8d 28 10 0.3 x 0.6 1.0 x 2.0 l 32 -25 -25 0.3 x 0.6 0.9 x 1.8 aAt location of minimum-depth critical flaw. bat inner surface of biological shield (cavity interface). C Safe shutdown earthquake. dLate 1988. ( 1 1 l i s 4 l
t i 10 t: ( . l l ORNL-DWG I 884837 ETD I ps it a
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r; 2 Y?' i,.9i 3 fd BIOLOGICAL - A,,id SHIELD g lI M W9 if l iv:
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Fig. 1.1. A PWR vessel support located in the cavity (- between the vessel and the biological shield. u
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- 2. HFIR VESSEL SURVEILLANCE DATA The High Flux Isotope Reactor (HFIR)1 is a high performance light-water-cooled, low-temperature (120-160*F) research reactor at ORNL that began operation in 1965. Its stainless-steel-clad carbon-steel pressure vessel (Fig. 2.1) was designed for 20 EFPY, and a surveillance program was maintained to monitor the actual radiation-induced embrittlement.2 Late in 1986, a reevaluation of the integrity of the vessel was com-menced in an effort to extend the permissible life.3 The 6 surveillance I data, which had not been carefully examined since 1974, indicated that the embrittlement rate was significantly greater than had been antici-pated on the basis of data obtained in the early 1960s from materials testing reactors (MTRs).* The neutron energy spectra and the irradia-( tion temperatures for the HFIR surveillance specimens and for specimens in the MTRs were believed to be essentially the same, and the materials l
l were very similari however, the fast neutron flux ($) in the MTRs was about 105 times that in the HFIR specimens. Thus, it appea-ed that the lower flux in HFIR was responsible for the relatively large amount of embrittlement per neutron; that is, there appeared to be a negative flu-ence-rate effect. The portions of the HFIR vessel that are subjected to the highest fast-neutron fluxes are close to the beam tubes (Fig. 2.1) because the beam tubes displace beryllium (reflector) and water that otherwise con-stitute shielding for the vessel wall. Thus, both shell material (A212-B) and beam-tube nozzle materials (A105-II and A350-LF3) were { included in the HFIR vessel materials surveillance program, and the sur-( veillance capsules, each containing three Charpy V-notch (CVN) specimens l
x . . l s e 12 - i y
'd,- and a flux monitor, were located close to the beam tubes (keys 1-7, Fig. 2.1)..
Surveillance specimens of A212-8 were removed for testing after 15.0 and 17.5 EFPY, and A105-II and A350-LF3 specimens were removed after 2.3, 6.5, 15.0, and 17.5 EFPY. The corresponding ANDTT data are presented in Table 2.1;" and in Figs. 2.2 and 2.3, plots of- ANDTT vs fast neutron fluence (9) (E > 1.0 MeV) are compared with the MTR data 5 i l l that were available at. the time the. vessel was designed. If it is l l_ assumed that spractrum and chemistry effects are not responsible for the l L incongruity of the several sets - of data, the comparison indicates _ a fluence-rate effect. l To evaluate the effects of possible differences in chemistry and fast spectrum, HFIR archive A212-B material was recently irradiated in
~
the Oak Ridge Research Reactor - (ORR), a. typical'MTR, and the HFIR and ORR ' A212-B data were plotted as a function of dpa for E > 0.1 MeV (Fig. 2.4)* as well as a function of 6 for E > 1.0 MeV (Figs. 2.2 and ] 1 2.3). (The ORR data and values of dpa corresponding to both the HFIR and ORR data are included in Table 2.1.) Figure 2.2 shows the A212-B archive material (irradiated in the ORR) to be' consistent with the MTR data ,- implying that the chemistry of the HFIR A212-B material is not significantly different than that corresponding to the MTR data (assum-ing that a difference in spectrum does not compensate for a difference in chemistry). Furthermore, surveillance data from Army reactors PM-1 and PM-1A' (Fig. 2.5) indicate no significant difference in sensitivity. I h *To obtain the MTR curve, it was assumed that the calculated spectrum for the ORR was appropriate for the "MTR" data.
\
h M 4 13 J k., between A212-B and A350-LF3, at least for the specific exposure condi-tions. These latter data also indicate that for the relatively high fluzes in the PM-1 reactors, the increase in NDTT for A212-8 at a flu-ence of 1-2 x 1017 n/cm2 is much less than that obtained from HFIR, ' indicating that the HFIR data are not simply a part of the tail of the MTR data. Figure 2.4 shows that when the HFIR and MTR data are plotted as a function of dpa (E > 0.1 MeV) there is still evidence of a significant rate effect. Thus, the small differences in the HFIR and NTR fast spec-tra are not responsible for the incongruity in Figs. 2.2 and 2.3. ) If an embrittlement rate effect exists, the total reaction rate associated with embrittlement should probably be accounted for when applying the HFIR data to a situation involving a different fast spec- l trum. This can be accomplished by correlating NDTT with dpa rate rather than e, integrating the dpa rate over all energies that contribute sig-nificantly to embrittlement. 1 In the previous discussions, the only spectral variations consid- I ered were those above 0.1 MeV. However, Hanstad et al.7 have suggested l that lower energy neutrons may be making a significant contribution, and
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that there might be significant differences in the low-energy spectra of HFIR and the MTRs. Multigroup transport calculations performed at ORNLs indicate that the HFIR spectrum above 1.0 MeV is somewhat harder than that in the ORR; the epithermal flux is a smaller fraction of 'the total non-thermal flux; and 4th *f, the ratio of the thermal to fast (E > 1.0 MeV), is greater in HFIR (-50) than in ORR (-8). Thus, the trend in Figs. 2.2 and 2.3 is not the result of differences in the epithermal
m 14 ; i th/*f sus 8est8 a Possible thermal-flux l .' fluxes, but the comparison of $ I effect. In addition, Alborman, et-al.8 irradiated A537 at 140*F in a spectrum with $ h/ 'f - 1000 and found that > 70% . of the embrittlement was due to thermal neutrons. Thus, perhaps thermal-neutron embrittle-l ment is at least in part responsible for what was otherwise believed.to be a rate effect.no At this time, however, there are not enough data to l l support this contention. ! References l
- 1. R. D. Cheverton and 7. H. Sims, NFIR Core Nuclear Design, ORNL -
4621, July 1971.
- 2. J. R. McWherter, R. E. Schappel,'and J. R. McCuffey, NFIR Pressure Vessel and Structural Components Material Surveillance Program, ORNL/TM-1372, January 1966.
- 3. R. D. Cheverton, J. C. Merkle and R. K. Nanstad, eds., Evaluation i
( of NFIR Pressure-Vessel Integrity Considering Radiation Rabrittle-ment, ORNL/TM-10444,' April 1988. I
- 4. R. K. Nanstad, S. K. Iskander, A. F. Roweliffe, W. R. Corwin, and C. R. Odctte, Effects of 50*C Survelliance and Test Reactor Irradi- '
\
ation on Ferritic Pressure Vessel' Steel Rabrittlement, presented at l 14th Int. Symp. on the Effects of Radiation on Materials, Andover, l MA, June 27, 1988, to be published in ASTM STP.
- 5. J. R. Hawthorne Studies of Radiation Effects and Recovery of Notch Ductility of Pressure Vessel Steels, British Nuclear Energy Confer-ence, Iron and Steel Institute, London, England, November 30, 1960.
- 6. L. E. Steele, J. R. Hawthorne, C. Z. Serpan, Jr., E. P. Klier and H. E. Watson, Irradiated Materials Evaluation and Reactor Pressure Vessel Surveillance for the Army Nuclear Power Program, NRL Memo-randum Report 1644, September 1, 1965.
- 7. R. K. Nanstad, K. Farrell, D. N. Broski, and W. R. Corwin, " Accel-ersted Neutron Embrittlement of Ferritic Steels at Low Fluence Flux and Spectrum Effects," submitted to J. of Nuc. Maths., April 1988.
- 8. R. L. Childs, personal communication to R. K. Nanstad,. August 10, 1988.
l
e a 15' Is.
- 9. A. Alborman, et al, " Influence of Thermal Neutrons on the Brittle-ness of High-Temperature Gas-Cooled Reactor Liner . Steel," Nuclear Technology, Vol. 66, p. 639-646, Sept. 1984.
- 10. L. K. Hansur and N. K. Farrell, private communication to R. D.
Cheverton and R. K. Hanstad, September 1968. l l l l l 4 l I (
v l 16 Table 2.1. Summary of irradiation effect for selected HFIR pressure vessel esterials Unirradiated $ 4,2 Material NolT EFPY M TT ,f,,D*s dpa/s' n/cm dpa j I C, (*F)] l'C, ('F)] (E > 1 MeV) (E > 1 Nev) { l ORR 1 irradiations l A2128 (LT) -21 (-5) 56 (101) 9.59 x 10 12 1.34 x 10-8 p,43 , jo 18 3,39 , jo-3 l A2128 (TL) -12 (10) 47 (85) 9.59 x 10 12 g,34 ,go-8 2.43 x 10 18 3,39 , jo-3 1 A2129 (TS) -21 (-5) 56 (101) 9.59 x 10 12 1.34 x 10-8 2.43 x 10 18 3,39 , go-3 j A2128 (EGCR)e 0 (32) 10 (18) I3 1.46 x 10~8 l7 2.14 x 10-4 l A2128 (EGCR)* O (32) 103 (185) 1.05x30 2x10 I 4.8 x 10'9 1.54xif8 9.8 x 10 2.4 x 10-2 l HFIR survelliance l A2128 (LT) -21 (-5) 15.01 11 (20) 4 x 10 7 1.89 x 10 16 15.01 29 (52) 2.43 x 10 8 3.66 x 10-I3 1.15 x 10 I7 1.73 x 10-4 17.53 42 (75) 2.43 x 10 8 3.66 x 10*I3 1.34 x 10 I7 2.02 x 10-4 A1051I -62 (-80) 2.34 10 (18) 4.66 x 10 8 6.92 x 10*I3 3.44 x 1 16 5.12 x 10-5 6.45 17 (30) 4.89 x 10 8 7.26 x 10*I3 9.9 x 10 6 i,43 , jo-4 15.01 33 (60) 4.89 x 10 8 7.26 x 10*I3 2.31 x 10 II 3.44 x 10-4 15.01 33 (60) 3.35 x 10 8 4,gg , jo-13 1.85 x 10 17 2.70 x 10~4 17.53 35 (63) 7.27 x 10 8 1.08 x 10-12 4.01 x 10 I7 5.97 x 10'4 A350LF3 (key 2) -79 (-110) 2.34 14 (26) 1.1 x 109 1.58 x 10-12 8.20 x 10 16 1.16 x 10~4 6.45 29 (52) 1.1 x 10 9 1.58 x 10-12 2.26 x 10 17 3.21 x 10'4 15.01 56 (100) 1.1 x 10 9 1.58 x 10-12 5.26 x 10 I7 7.47 x 10~d l 17.53 66 (118) 1.1 x 10 9 1.58 x 10-12 6.14 x 10 17 8.73 x 10'4 l l A350LF3 (key 3) -62 (-80) 2.34 19 (34) 1.29 x 10 9 1.85 x 10-12 9.55 x 10 16 1.36 x 10-4 9 6.45 33 (60) 1.40 x 10 2.01 x 10*I2 2.84 x 10 I7 4.09 x 10 15.01 54 (97) 1.03 x 10 9 1.48 x 10-12 4.88 x 10 17 7.01 x 10~d
'7.53 64 (115) 1.29 x 10 9 1.85 x 10-12 7.12 x 10 17 1.02 x 10'3 l
- Material from previous ORNL study on Experimental Ges-Cooled Reactor.
(
17 I I f. e oRNL-DWG 37 4472 gTD, SEAM WELD -
//////
REACTOR PRES $URE VESSEL
/
REACTOR COOLANT DISCHARGE HB*i 6 . . HB-4 o
- ~WE/g' / i x [THRU ATust n\ M K )
p !) \ Q tl l- ) l c M i8' ' nj-, -; g/> '
= a g .-
INNER FUEL ELEMENT ,
/,
] P y ;,:,5lllr tg_; 5 J
SEMI PE*M ANENT R E F LECTO R - O_
'1
s \ v , , ., j
,, g 4. REACTOR REFLECTOR LINER s
CONTROL PLATES v ( CONTA PLATE % TANGENTIAL TUBE p OO ' o ;
- l EXPERIMENT ACCESS 3' OUTER FUEL ELEMENT
\
- k ISOTOPE IRRADIATION ACCESS W HB*3 n RACIAL TUBE 4 \ ,
\
ozzLE WELD I ) Fig. 2.1. Cross section of HFIR vessel and core at midheight of core, indicating locations of vessel surveillance specimens (Keys 1-7). l
18 A. I I I lllIlj IRRADIATED l l I i i il 3l l l i l i Ill
~
~ A A201 STE E L '
O A212B 1960's ""
~ O A302B "
TEST-0 SA336 I . REACTOR
~
~ E WELD A DATA 9 WELD B . o 300 - TEMPERATURE < 200'F -
DATA OF HAWTHORNE masemammes
$ $ A2128 - ORR,120'F
~
w g -A A2128 - HFIR SURV.120'F - E, - - E /C 3 200 - aj - l 3 1 o - j _ (E > 1.0 MeV) : i E 1 X 1013 n/cm2.s 1 e _ a , k- 5 100 - l
-A _
4% = 2.4 X 108 _ 0 t i I IItlll I I t i t 11I! I t 1 i t t 11 1017 1 18 10 s 10 1020 HEUTRON FLUENCE (E > 1.0 MeV) (n/cm2) l Fig. 2.2. Increase in NDTT with fluence (E > 1.0 MeV) for l A212B irradiated in HFIR (vessel surveillance positions) and ORR, l and for several similar materials irradiated in MTR (Hawthorne). 1 - b 1
19
,=
6 I
. - . ~
O ORNL DWG 88-4785 ETO 350 l l I MTR DATA (HAWTHORNE annammmer) [o (E > 1 WV).1 X 1013 rvem2,3)
* ~ HFIR SURVEILLANCE:
O A2128 (o 2.4 X 108) v A105!!(c 3 7 X 108) - 150 9 O A350 LF3 (1.01.4 X 10 ) 250 -
'200 -
E 100 E ; g - ! [* k $' 150 - Oo 100 - g
- 50 0
vvO V 1 5c - 0 O l O V
. .e ' o ! ! O j 1016 1017 1010 1019 102 i 2 !
F1.UENCE (E > 1 WV) (neutronsern ) J Fig. 2.3. Increase in NDTT with fluence (E > 1 MeV) for , irradiations in HFIR (vessel surveillance positions) and MTR j (Hawthorne). i 1 l (. ! l,
20 t[ 300 , j , , , , , , , , i
, , ,igig HFIR SURVEILLANCE: ,
4 A212B ($ = 2.4 X 108 n/cm2.s) y A*.0511 ($ = 3 7 X 108) l
-200 - 9 A10511, A350 LF3 ($ = 1.01.4 X 109) -
l C l
' ORR:
o O HFIR R212 B (4 =9.6 v10) l z i
- Og 100 -
p O - MTR
- T < 2000F
(. 4 Y gY# Y
$(E > 1.0 MeV) = ,
1 X 1013 n/cm2.s g I l 1 l l l Ill 1 l I I lllll l 104 10 3 10 2 DPA (E > 0.1 MeV)
- Fig. 2.4. Increase in NDTT with DPA for irradiations in HFIR (vessel surveillance positions), ORR, and MTR (Hawthorne).
t o______--_---_---- - - _ - - - _
m , j 21
'\
ORNL DWG 88 4459 ETD , 600 333
. . i;6 ,
t ! 1; 'Iji n i l liiii i i .l11 7 i- n 2 ARMY REACTOR PRESSURE VESSEL STEELS ~ A/240 -
, c , _ e 3 5/8. in. A 350 LFI(MOD) PLATE /
l i 9 - (SM 1 A) , 300 / l -
$ 500 Z-e 2 5'8.in. A 350 LFI(MOD) FORGING A./ / I d ~~
278 (SM 1 A1
$ Y 4.in. A 212 GRADE B PLATE / ' -
, j ,
e - (SMI) , 250 - E : A 2.4. in. A 350 LF3 FORGING O'
-# 8'
$ 2 (PM 2A) , ! I g 400 ,
222 5 E o = 1010 , j en nedrons'em2.s i III , E C - 200,! 500l -
$ [ 3 E
- 430[ /
f [ 510 - ; 167 { ' I
- h. 303
~
1 490 2
~ ('C)
$ .I TRENOFOR -
j 3 240 {, / <450' F IRRADIAllONS l l ( '00 -
~
111 8 : I SM 1 A - SURVEILLANCE - i 2 ] 5 gg l SMI 445 475cp
-~
l
}
_,_ g SURVEILLANCE;l* O '
= 4750F 2 1 + 2
- , fy -
2 t ' ' li t!'I i ! i liin- ! i I llin- ' I I llitt 0 i i f lite 0 10" 10H 10" lon INTEGRATED NEUTRON EXPOSURE (neutionstm 2 > 3 yey) Fig. 2.5. Effects of neutron radiation on the Charpy-V transition temperature behavior of Army Reacter Pressure
'essel steels. Numbers adjacent to data points indicate , exposure temperatures.
v 22 k 3. APPLICATION OF HFIR DATA TO VESSEL SUPPORT EVALUATION Soon after the HFIR surveillance data were evaluated in late 1986, it became apparent that the indicated embrittlement rate effect might apply to the supports of some LWR' vessels because f ast neutron fluxes, irradiation temperatures, and materials were thought to be similar. Temperatures of the supports range from -500*F, at the point where they contact the vessel, to < 150*F at a point where they contact the bio-logical shield (150*F is the normal maximum permissible operating tem-perature of the concrete biological shield). The temperature of the HFIR vessel and surveillance specimens is -120*F. Thus, presumably a portion of the support operates at a temperature close to that of the HFIR vessel. ( Hultigroup neutron transport calculations were performed recently ; for the vessel wall and the cavity of one BWR and three PWRs, and the results are provided in Tables 3.1-3.4.1,2 Table 3.5 summarizes the fast fluxes (E > 1.0 MeV) (as well as dpa rate (E > 0.1 MeV)] for the LWR cavities and the HFIR surveillance specimens (also, see Table - 2.1). It is apparent that 4 (E > 1.0 MeV) for the PWR cavities is similar to those for the HFIR surveillance specimens (10a.los n/cm2's), while that for the BWR is much less. l Table 3.6 summarizes LWR f ast-flux data for E > 1.0 MeV (Croup A) j i and 0.1 5 E s 1.0 MeV (Croup B). These data indicate that the ratio of l 1 group A to group B fluxes is much less in the cavity than it is at the inner surface of the vessel wall (the result of inelastic scattering in the vessel wall). Thus, the fast flux (E > 0.1 MeV) in the LWR cavity ( is much softer than that at the location of the HFIR surveillance ) 1
, .~ a
~ l 23 i
< \
( specimens. As suggested in Sect. 2.0, to account for this difference in i energy spectrum when applying the HFIR data to the evaluation of the i supports in the cavity, the ANDTT data can be correlated with dpa rate l and dpa for E > 0.1 HeV instead of $ and 9 for E > 1.C MeV, the sesump-tion being made that most of the neutrons contributing to embrittlement have energies above 0.1 HeV. A comparison of dpa rate (E > 0.1 MeV) for HFIR and the LWR cavities (Table 3.5) indicates that the maximum cavity dpa rate (5.0 x 10 12) is about twice the maximum HFIR dpa rate t e (2.0 = 10 12), while the maximum fast flux values (E > 1.0 MeV) are about the same. This indicates that some extrapolation of the HFIR data l is necessary. Application of the HFIR data to the LWR vessel supports requires extrapolation with regard to both dpa rate and dps. Thus, correlations I between ANDTT, dpa rate, and dpa are required. Two different correla- t tions between ANDTT and dpa are proposed and are shown in Figs. 3.1 and 3.2. Figure 3.1 was created by drawing separate straight-line curves through the maximum ANDTT data points, for the A212-B and A350-LF3 HFIR surveillance materials, parallel to the MTR curve, which Hawthorne con-structed as a straight line on semilog paper.3 The other correlation (Fig. 3.2) was obtained by first plotting the HFIR A350-LF3 data on log-log paper and cons'tructing a best-fit straight-line curve. Next, the ORR A212-B data obtained in connection with the recent HFIR vessel study" and a much earlier study of the ves-sel for the Experimental Cas-Cooled Reactor (ECCR)5 (Table 2.1) were plotted and a straight line constructed. And, finally the MTR data 3 above ANDTT = 100*F were plotted and a straight line constructed. The
24
.. ' indication is that the relatively high fast-flux data (ORR and MTR) are .
essentially parallel to the HFIR A350-LF3 data. This was used as justification for constructing a curve through the HFIR A212-B data points parallel to the A350-LF3 curve. The upper A212-B point was used as a conservative measure.
- A comparison of the correlations represe.mted by Figs. 3.1 -and 3.2 (upper two curves) shows that Fig. 3.2. predicts less of a rate effect between 4(E > 1.0 HeV) = 2.4 x 10s and 1.2 x 108 n/cm2.s. In this study, the authors have considered the correlation implied by Fig. 3.2 1 to provide the more accurate crimate of ANDTT for the LWR vessel sup-1 ports.
One might argue that the small difference in fast fluxes corre-sponding to the HFIR A212-B and A350 LF3 irradiations (2.4 x 108 and 1.2 = 109 n/cm# s) relative to the factor of -105 between HFIR and the HTRs (-los and 1013 n/cm2.s) would not permit distinguishing between the I l two HFIR fluxes with regard to establishing a rate effect. However, Hamilton 6 recently presented data which indicate that at -200*F there was essentially no rate effect in the fast flux (E > 1.0 HeV) range of 1 = 1010 to 3 x 1013 n/cm2.s. This indicates that the ANDTT differences
-4' observed between HFIR and the MTRs is associated with a rate effect l
below a fluence rate of -1 x 1010 n/cm2.s. Thus, for this study, the observed difference in ANDTT for fluxes of 2.4 x los and 1.2 109 n/cm2.s were considered to be real. There is, however, an inconsistency with regard to the A105-II datas although these data correspond to an intermediate flus level (3-7 x 108 n/ cat.s), they tend to coincide with the A350-LF3 data, which correspond to a higher flux (1.2 x 108
j 25 { l (- n/cm2.s). Perhaps this implies that a rate effect is not discernible ! within the flux range of 2.4 x los to 1.2 x 108 n/cm2.s. However, for this study the A105-II data were discounted insofar as establishing a rate effect. Extrapolation and interpolation of the HFIR data for application to the support study was accomplished by assuming that dpa a (dpa rate)" , 1 for a given value of ANDTT. Corresponding values of dpa, dpa rate, and - ANDTT were taken from Figs. 3.1 and 3.2 to obtain the log-log plots in ! 1 Figs. 3.3 and 3.4, respectively. As mentioned above, Fig. 3.2 and thus j i Fig. 3.4 are believed to be the more accurate representation of the trends. However, it must also be emphasized that in either case the values of ANDTT > 100*F and values of dpa rate outside the range 3.7 10 13 - 1.7 x 10 12 s'1 represent extrapolations of the HFIR data, g L. References
- 1. Nicolas Tsoulfanidis et al., " Neutron Energy Spectrum Calculations in Three PWR's," Proceedings of the Fifth ASTM-EURATONS Symposium on Reactor Dosimetry, p. 693-701 (1985).
- 2. M. L. Williams, ORNL, personal communication.
- 3. J. R. Hawthorne, Studies of Radiation Effects and Recovery of Notch
, Ductility of Pressure vessel Steels, British Nuclear Energy Confer-ence, Iron and Steel Institute, London, England, November 30,.1960.
- 4. R. D. Cheverton, J. C. Merkle and R. K. Nanstad, eds., Evaluation of HFIR Pressure-Vessel Integrity Considering Radiation Embrittlement, ORNL/TM-10444, Martin Marietta Energy Systems, Inc., Oak Ridge Natl.
Lab., April 1988.
- 5. M. S. Wechsler, R. G. Berggren, N. E. Hinkle, W. J. Stelsman, Radia-tion Hardening and Embrittlement in a Low-Carbon Pressure Vessel Steel. Irradiation Etfects in Structural Alloys for Thermal and P.sst Reactors, STP 457, American Society for Testing and Material, Phil.,
1969, pp. 242-260. ( i J
s . l 26 l 1 l (' 6. . M. L. , Hamilton, H. L. Heinesch, Pacific Northwest Laboratories,
" Tensile Properties of Wautron Irradiated A212-B Pressure Vessel Material," ASTM 14th International 3ymposium on Effects of Radiation on Materials, Andover, Massachusetts, June 27-29, 1988.
4 e
27 ( Table 3.la. Calculated neutron energy spectra for i B&W (ANO-1) Reactor 3-D DOT FLUX (n/cm2.s)' 1 l
"PP'* I" # "E A Group In cavity I (HeV) of PV of PV i 1 1.733+01 2.699+07 1.258+07 5.175+05 2 1.221+01 6.543+07 2.940+07 1.061+06 3 1.000+01 2.743+08 1.205+08 3.694+06 4 7.408+00 7.879+08 3.242+08 7.908+06' I
$ 4.966+00 1.133+09 4.882+08 1.348+07.
6 3.012+00 7.621+08 3.822+08- 1.244+07 7 2.466+00 3.448+08 1.794+08 5.967+06 8 2.307+00 1.376+09 9.010+08 3.893+07 l 9 1.653+00 2.000+09 1.688+09 1.176+08 l 10 1.003+00 1.108+09 9.956+08. 1.036+08 l 11 7.427-01 1.863+09 2.314+09 3.413+08 l 12 4.979-01 2.052+09 3.036+09 5.289+08 I 13 2.972 1.171+09 1.214+09 2.908+08 ! 14 1.832-01 1.235+09 1.643+09 3.327+08 ' 15 1.111-01 8.378+08 1.303+09 2.556+08 16 6.738-02 3.851+08 6.444+08 1.953+08 17 3.183-02 1.820+08 7.730+07 2.832+07 2.606-02 (' 18 19 2.418-02 3.043+08 4.703+08 6.040+08 6.469+08 1.094+08 2.004+08 20 1.503-02 6.938+08 3.110+08 1.652+08 21 7.102-03 2.697+09 1.528+09 4.918+08 I 22 4.540-04 1.489+09 8.032+08 2.364+08 23 1.013-04 3.968+09 1.807+09 4.823+08 24 1.855-06 1.272+09 3.777+08 1.309+08 25 4.140-07 9.277+09 2.232+08 5.184+08 26 1.000-11 I Total 3.641+10 2.152+10 4.613+09 Table 3.lb. 4-Croup fluxesa for B&W (ANO-1) PWR Croup ",'#8Y OT 1/4T Cavity 1 1.0 MeV - 17 MeV 0.186 0.192 0.044 2 0.111 MeV - 1.0 MeV 0.204 0.428 0.346 3 0.4 eV - 0.111 MeV 0.355 0.370 0.498 4 10 5 - 0.4 eV 0.255 0.010 0.112 dpa/s (E > 10 5 eV) 1.041E-11 7.232E-12 6.829E-13 (- aWormalized to unity.
. . 1 28
. Table 3.2a. Calculated neutron energy spectra for combustion Engineering.(ANO-2) Reactor
{ 3-D DOT FLUX (n/cm2.s) Croup E upper 'In front- At T/4 3, ,,ygty (HeV) of PV of PV 1 1.733+01 7.373+07 3.189+07 1.622+06 2 1.221+01 2.330+08- 9.728+07- 4.384+06 3 1.000+01 1.265+09 5.148+08 1.990+07 4 7.408+00. 4.565+09 1.724+09 5.359+07 5 4.966+00 7.486+09 2.910+09 9.773+07
'6 3.012+00 5.313+09 2.383+09 9.108+07 7 '2.466+00 2.438+09 1.129+09 4.406+08 8 2.307+00 9.928+09- 5.725+09 2.754+08 9 1.653+00 '1.492+10: 1.104+10 8.080+08 10 1.003+00 8.237+09 6.600+09 7.108+08 11 7.427-01 1.460+10 1.602+10 2.258+09 12 4'.979-01 1.567+10 2.041+10 3.328+09 13 2.972-01 8.603+09- 8.400+09 1.932+09 14 1.832-01 9.282+09 1.138+10 2.164.09 15 1.111-01 6.773+09 8.073+09L 1.585+09 16 6.738-02 6.212+09 4.538+09 1.320+09 17 3.183-02 9.359+08 4.206+08 2.006+08 18 2.606-02 2.570+09 3.218+09 7.005+08 19 2.418-02 4.319+09 4.496+09 1.320+09 20 1.503-02 4.580+09 2.194+09- 1.119+09 g 21 7.102-03 1.829+10 9.711+09 3.355+09
% 22 4.540-04 9.975+09 4.705+09 1.626+09 23 1.013.-4 2.731+10 1.172+10 3.458.09 24 1.855-06 8.374+09 2.297+09 9.823+08 25 4.140-07 5.962+10 1.753+09 4.648+09 26 1.000-11 27 Total 2.516+11 1.415+11 3.254+10 Table 3.2b. 4-Croup fluxes
- for CE (ANO-2) Reactor I
~
Croup ",'#87
, OT 1/4T Cavity 1 1.0 MeV'- 17 WeV 0.184 0.181 0.055' 2 0.111 MeV - 1.0 MeV- 0.224 0.444 0.319 3 0.4 eV - 0.111 MeV 0.355 0.363 0.483 4 10 5 eV - 0.4 eV 0.237 0.012 0.143 dpa/s (E > 10 5 eV) 7.041E-11 4.549E-11 4.975E-12 aNormalised to unity.
29 Table 3.3a. Calculated neutron energy spectra for Westinghouse Reactor McQuire Unit 1, Cycle 1
'{ 3-D DOT FLUX E upper In front At T/4 In cavity
# "P (MeV) of PV of PV R=323.83 cm 1 1.733+01 5.062+07 2.105+07 3.020+05 2 1.221+01 1.567+0S 6.249+07 7.764+05 3 1.000+01 8.316+08 3.236+08 3.293+06 4 7.408+00- 2.893+09 -1.048+09 '8.135+06
'5 4.966+00 4.424+09 1.662+09- 2.014+07-6 3.012+00 3.001+09 1.331+09 2.193+07 7 2.466+00 -1.360+09' 6.255+08 1.135+07 ,
8 .2.307+00 5.295+09 3.126+09 9.912+07-9 1.653+00 7.603+09 5.901+09 4.448+08 10 1.003+00 4.142+09 3.448+09 4.510+08 11 7.427-01 6.947+09 28.317+09 2.546+09-12 '4.979-01 -7.404+09 1.087+10 4.565+09
'13 2.972-01' 4.298+09 4.316+09- 2.449+09 14 1.832-01 4.468+09 6.069+09 3.419+09 15 1.111-01. 3.380+09- 4.375+09 2.604+09 16 6.738-02 3.362+09 2.360+09 2.024+09 17 3.183-02 6.466+08 2.369+08 3.793+08 18 2.606-02 1.153+09 -1.697+09 1.490+09 1 19 2.418-02 2.209+09 2.410+09 2.812+09 )
20 1.503-02 2.657+09 1.170+09 1.946+09 { 21 22 7.102-03 4.540-04 1.016+10 5.600+19 5.142+09 5.942+09 2.547+09- 2.867+09 23 1.013-04 1.473+10 -5.865+09 5.993+09 24 1.855-06 4.670+09 1.151+09 1.680+09 25 4.140-07 3.382+10 7.968+08 7.164+09 26 1.000-11 l 27 Total 1.353+11 7.487+10 4.894+10 Table 3.3b. 4-Croup fluxes
- for W Reactor Croup ", nge
,' 8 OT 1/4T Cavity 1 1.0 Mev - 17 MeV- 0.184 0.188 0.012 2 0.111 MeV - 1.0 MeV 0.202 0.441 0.274 3 0.4 eV - 0.111-MeV 0.359 0.360 0.567 4 10~5 eV - 0.4 eV 0.250 0.011 0.146 .
i dpa/s E > 10-5 eV) 3.874E-11 2.484E 4.712E-12 aNormalised to unity.
30 Table 3.4a. calculated neutron energy spectra for CE Reactor
-{.~
Croup E upper In front At T/4 gg,y)
,g py ,g py In cavity 1 1.733E+01 2.78532E+05- 1.19476E+05 1.26309E+04 2 1.419E+01 1.17743E+06 5.10065E+05 5.31195E+04 3 1.221E+01 4.20585E+06 -1.73173E+06 1.64842E+05 4 1.000E+01 7.86872E+06- 3.22077E+06' 2.92429E+05 5 8.607E+00 1.27925E+07 5.10243E+06 4.31247E+05 6 7.408E+00 2.97338E+07 1.15308E+07 8.93362E+05 7 6.065E+00 3.82946E+07 1.42439E+07 1.06782E+06-8 4.966E+00 5.75978E+07 2.15559E+07 1.67350E+06 9 3.679E+00 3.59718E+07- 1.51213E+07 1.29333E+06 10 3.012E+00 2.41009E+07 1.12497E+07 9.89848E+05 11 2.725E+00 2.59448E+07 1.28555E+07 1.19784E+06 12 2.466E+00 1.25811E+07 6.32145E+06- 6.21796E+05 13 2.365E+00 3.22105E+06 1.78435E+06 1.76455E+05' j 14 2.346E+00 1.50974E+07 8.68747E+06 8.79155E+05 15 2.231E+00 3.58901E+07 2.23694E+07 2.14032E+06 16 1.920E+00 3.56560E+07 2.62479E+07 2.89304E+06 17 1.653E+00 4.59384E+07 3.65276E+07- 4.35308E+06 18 1.353E+00 6.38427E+07 6.42605E+07 9.28453E+06 19 1.003E+00 4.05210E+07- 4.46711E+07 7.60235E+06 20 8.208E+01 2.11290E+07 1.92272E+07 4.04310E+06 21 7.427E+01- 4.78865E+07 6.94558E+07 1.34120E+07
( - 22 23 24 6.081E+01 4.979E+01 3.688E+01 4.06699E+07 4.48645E+07
-4.04880E+07 5.56024E+07 6.27020E+07 6.98931E+07' 1.30706E+07 1.33038E+07 1.48399E+07 25 2.972E+01 6.22632E+07 7.57791E+07 2.24843E+07 26 1.832E+01 5.40647E+07 7.47653E+07 2.27497E+07 27 1.111E-01 4.23644E+07 4.61578E+07 1.65013E+07 28 6.738E-02 3.61008E+07 3.55079E+07 1.25341E+07 29 4.087E-02 1.56710E+07 9.46462E+06 4.06071E+06 30 3.183E-02 1.18315E+07 2.92956E+06 4.08343E+06 l 31 2.606E-02 7.64521E+06 1.84057E+07 5.40984E+06 32 2.418E-02 7.29383E+06 1.11433E+07 3.93861E+06 33 2.188E-02 2.23029E+07 1.78285E+07 8.46670E+06 34 1.503E-02 4.12070E+07 2.00703E+07 1.19762E+07 3F 7.102E-03 4.04597E+07 2.52165E+07 1.15085E+07 36 3.355E-03 3.90943E+07 1.99575E+07 1.03076E+07 37 1.585E-03 6.33659E+07 3.07840E+07 1.57353E+07 38 4.540E-04 3.75323E+07 1.35174E+07 8.53409E+06 39 2.144E-04 3.76390E+07 1.74130E+07 8.29094E+06 40 1.013E-04 5.01136E+07 2.50163E+07 1.05576E+07 41 3.727E-05 6.23387E+07 3.06902E+07 1.23638E+07
, 42 1.068E-05 3.71304E+07 1.64081E+07 6.92802E+06 43 5.043E-06 4.87669E+07 1.80571E+07 8.48279E+06 44 1.855E-06 3.58256E+07 9.95936E+06 5.74201E+06 45 8.764E-07 3.49526E+07 6.91034E+06 3.16106E+06 46 4.140E-07 1.19921E+08 6.44878E+06 1.10811E+07 47 1.000E-07 1.25993E+09 9.40594E+06 3.51009E+07 48 1.000E-11 (: Total 2.8536E+09 1.1268E+09 3.5669E+08
- q e t 31 i
Table 3.4b. 4-Croup fluxes" for CE Reactor
# "E bound ry OT 1/4T Cavity down omer iange j 1 2.0 MeV - 17 MeV 0.096 0.149 0.158 0.234 0.080.
2 0.111 MeV - 1.0 MeV 0.314 0.099 0.123 0.419 0.313 3 0.4 eV - 0.111 MeV 0.460 0.237 0.235 0.333 0.478 4 10~5 eV - 0.4 eV 0.130 0.515 0.484 0.014 0.129 dpa/s (E > 10~5 eV) 6.909E-13 4.303E-13 6.335E-14
" Normalized to unity. l l
l l 1 1 i sh' . (?- C
e-3 2 ,. Table 3.5 Suannary of fast neutron fluxes (E > 0.1 MeV) for LWR cavities and HFIR surveillance specimens i
/ \ 1 Reactor $(E > 1.0 MeV), n/cm2 .,- dpa/s (E > 0.1 MeV) I l
HFIR 2.4 x 108 - 1.4 x 109 3.7 x 10 2.0 x 10-12 CE (BWR)a 2.8 = 107 6.3 x 10-14 l B&W (PWR)b 2.1 x 10 8 6.8 x 10-13 W (PWR)C 6.1 x 108 4.7 x 10-12 l
, CE (PWR)d 1.8 x 10 9 5.0 x 10-12 8
Ceneral Electric boil ng water reactor. b Babcock and Wilcox pressurized water reactor. Westinghouse pressurized water reactor.
'd Combustion Engineering pressurized water reactor.
I l l k (
33 Table 3.6. Summary of calculated fluxes for " typical" BWR and PWR vessels and cavities s Neutron fluz (n/cm2.s) Reactor Vessel inner surface Cavity Aa .Ba A/B A B A/B j CE (BWR) 6.3 m 10s 3.5 x los 1.8 2.8 x 107 1.1 x los o.3 B&W-(PWR) 5.8 x 109- 7.4 x 108 0.8 2.1 x los 1.6 x 108 0.1 W (PWR) 2.6 = 1010 2.7 x 1010 1.0 6.1 x los 1.3 x.1010 0.5 CE (PWR) 4.6 x 1010 5.6 x 1010 0.8 1.8 x 109 1.0 x 1010 0.2 aAs E > 1.0 MeV. .B: 0.1 5 E 5 1.0 MeV. 1 1 f i i (~~ ! i l
34 l 1 ( 1 I I I I IIIll / I i 1/1 i i lll
/ /
$ HFIR A212-B / /
300 -
/ /
~
G HFIR A350-LF3 V HFIR A105-II ! j j
/ /
/ /
/2.4 X 108 /
C / 3.7 X 1013 /1.2 X 109 L 200 - -
/ / 1.7 X 10-12 C / /
E / /
* / /
/ /
e' ( 100 -
/
/
g!, -
/ /
$ / (E > 1 MeV) =
4
/ VV #,e' Y 1 X 1013 n/cm2.s A DPA/s =
1.4 X 10 8 ) 0 I I I I IIIII I I l I I I III l 10-4 10 3 10 2 DPA (E > 0.1 MeV) Fig. 3.1. Method for extrapolating HFIR vessel surveillance ANDTT vs DPA (E > 0.1 HeV) data, assuming ANDTT = 145 in DPA/d*F
, for ANDTT > 50*F.
.u _
~ _
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l HFIR DATA 10-4 I I I I IIll! I I I I I I ll 10 13 10 12 10-11 ! , .3- , DPA RATE (dpais) i l l Fig. 3.3. DPA (E > 0.1 MeV) vs DPA rate for specific values ! l of ANDIT, based on data in Fig. 3.1. l 1 i l
* = .- 1
37 d. 4 ORNL DWG 864784 ETD i i i i i i a ii... iiisiig i iiiiig ANOTT FF), 10 2 _
~
300 J ~ 2 250
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.. 160 .
140 ) _ 120 j ! 100 i 1 80 i 103 --- -
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. 60 50
- 1 30 20
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/
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f+-HFIR DATA + 10'S ' ' ! ' ' ' 'Ii ' ' ' ' ' ' ' ' 1014 1013 3o 12 3 o 11 DPA RATE I k Fig. 3.4. DPA (E'> 0.1 HeV) vs DPA rate for specific values j of ANDTT, based on data in Fig. 3.2. i i ! l l j
m 38 i . (., 4. SURVEY OF SUPPORT DESIGN FEATURES 4.1 Introduction A survey was made of light water reactor vessel supports using the file of Final Safety Analysis Reports (FSARs) located at the Nuclear Operations Analysis Center of the Oak Ridge National Laboratory (ORNL). Because of limited time and resources, virtually no contacts for data were made with industry, i.e...with nuclear steam system sup-pliers (NSSS), architect-engineer (AE) organizations or with utili-ties. The objective of this survey was two-folds (1)-to determine if supports or any part thereof are susceptible to significant radiation damage, and (2) to determine if the loading conditions and/or fabrica-tion techniques would result in significant primary or secondary tensile (; stresses. Two prior studies t,2 were very useful in elucidating, corroborat-l' ing, and supplementing the FSAR data. For several nuclear plants the - available data from all sources were too sketchy,to be definitive. Con-sequently, the reader should understand. that although most of the FSAR data are clear, there may be specific plants for which the inferences
, are based on meager data and may be incorrect. Nevertheless, for most plants the survey provides data that permit an understanding of - the design rationale followed by NSSS and AE designers.
Of the 125 light-water reactor plants surveyed, 96 plants. were operating, with the. remainder having received construction permits. 1 Since the survey was completed, one of the plants, Lacrosse, has been i
. decommissioned.
m - t # 39 I
/ ! 4.2 Categorization The reactor vessel support designs for the 125 plants were divided into five - broad categories. These categories, which are assigned to each plant in Table 4.1, include'(1) skirt, (2) long column, (3) shield l l
tank, (4) short column, and (5) suspension. All of the General Electric ] Co. BWR plants (41), except two (Big Rock Point and Lacrosse), have skirt supports typically as shown in Fig. 4.1. (As noted previously, the Lacrosse plant has been decommissioned.) Bectuse the skirt supports are located remotely from the core with a large volume of intervening metal and water, radiation damage is not anticipated. In addition, the very large section modulus of the skirt's l l design is anticipated to result in very low stresses resulting from bending loads of all types. The Big Rock Plant' reactor vessel, the only category-5 plant in Table 4.1, has suspension supports as shown in Fig. 4.2. The suspension supports are designed to have significant primary tensile stresses and are also anticipated to reside within a l zone of potential radiation damage. Seven of the PWR plants have skirt supports for the reactor vessel (Fig. 4.3). The comments on the BWR skirt supports are believed to apply also to the PWR vessel supports (i.e., insignificant radiation damage and low tensile stress state for all loading conditions). Babcock and Wilcox served as the NSSS on all seven of thase plants. Supports for 11 plants have been placed in the long column category (Fig. 4.4). Small keyway gap siae leads' to the inference that tensile - stresses in the column from all loadings should be low. On the other hand, the columns are located in a sone of potentially high radiation {.
40
/
\. damage. For nine of the plants in this category, Combustion Engineer-ing, Inc., acted as the NSSS, and for two of the plants, Westinghouse Electric Co. acted as the NSSS. I Eight plants have had the supports categorized as shield-tank type. For all of these plants, Stone & Webster has acted as the AE.
Westinghouse Electric, Inc. was the NSSS for seven of the plants, and Combustion Engineering, Inc., for one of the plants. It is not complet-ely clear from FSAR data if the supports for this latter plant should be classified as shield tank. Figure 4.5 shows design features typical of l shield-tank-supported reactor vessels. It is evident that much of the structure would be highly susceptible to the radiation damage hypothe-sized herein. On the other hand, it would be ~ anticipated that the structure would be quite stiff and consequently have very Icw primary stresses from all loading cases. It is uncertain as to whether second- l ary stresses (e.g., thermal or residual) could have a significant role in a fracture analysis of a support of this type. Category 4 (short column), which includes the largest number of PWR plants (58), is an assemblage of a potpourri of designs having two prin-cipal cotanonalities t (1) loads, vertical and horizontal, are transmit- , ted from. the vessel nozzles (and vessel brackets in a few instances) l l through a radially keyed structure to the biological-shield wall, and l (2) none of the structural components within the reactor vessel biolog-ical-shield annulus extends below the core midplane. In general, the biological-shield wall of most plants is designed for a maximum temperature of 150*F for protection of the concrete. As a consequence, many of the Category-4 supports, because of the short 1 I -
_ x _ _. - __ -__ 41 I
.- heat-ficw path from reactor vessel nozzle to biclogical shield wall, ;
A ' have special designs for heat dissipation. In particular, many of the j short-column designs are weldsents with openings to promote natural-i convection cooling or to assist forced-convection cooling. Others use l cooling coils with water as the cooling medium. Because of the large i variation in designs of these structures few generalizations can be ' made. They appear to have a large variability in susceptibility.to rad- ! l istion damage because of the large variability in vertical location rel-ative to the reactor core.- On the - other hand, designs with relatively broad cavities between the reactor vessel and biological shield are sub-l 1 ject to neutron streaming that elevates the fluence experienced for a ' significant distance above the top of the core. j Because of inferred design characteristics that would promote sig-nificant differences in radiation damage susceptibility or in bending or 1 tensile primary stress fields, Category-4 was subdivided into subgroups 4A through 4G. Figure 4.6 provides several views and details of a PWR vessel with a Category-4A support having some features that are similar to the Trojan and Turkey Point plants, which were selected for detailed evaluations (Sects. 5, 6, and 7). In this support category, vertical loads are transmitted f rom vessel nozzle veld pads through a short-keyed pedestal or girder to steel beams cantilevered radially inward from the biological shield wall; and horizontal loads are transmitted through the same load path or to separate structures. The radial beams of this cat-egory are susceptible to potential radiation damage and are designed to resist vertical loading with primary bending stresses.
m 42 l
/
5 Figure' 4.7 is a sectional elevation view of a PWR vessel with a Category-4B support. In this design, loads are transferred from vessel brackets to a ring girder weldment that is supported from the biological shield. Since the loadings are resisted by the girder by means of a
, torsional moment, primary bending stresses are induced. The girder is located below the top of the core but tends to be shielded by the adja-cent suspended neutron shield tank. Thus, the potential for radiation I
damage is unclear. i } l Figure 4.8 provides perspective and elevation views of a PWR vessel equipped with a Category-4C support. Vertical loads pass from the ves-sei nozzle weld pads through radially keyed shoes to short weldment ped-estals anchored on the bottom surfaces to recesses in the biological shield. Horizontal loadings are resisted by bending and shear stresses ( . in the shoes and pedestal weldments. Insufficient information is avail-able for assessment of potential secondary tensile stresses, i.e., from thermal or residual-stress sources. Because of the short height, heat , transfer provisions (vent openings, water cooling) are incorporated in the design to promote rapid cooling. Radiation damage susceptibility is i among the lowest of Category 4 because of the short height generally . 1 empl oyed. This design concept has the highest frequency of use within Category 4 with its employment in 25 plants. Figure 4.9 provides a section elevation vieu of a PWR vessel equip-ped with Category-4D supports. Features are similar to Category 4E with the employment of short pedestal weldsents to transfer vessel loadings to either reinforced concrete corbels or structural-steel brackets can-tilevered radially inward from the biological shield. Except for the i I
a s 1 43 i ( uncertainty as to whether structural-steel brackets are actually used, the susceptibility to radiation damage and the level'of primary and see-ondary stresses would appear to be similar to Category-4C. Figure 4.10 provides plan and elevation views of a PWR vessel l equipped with a Category-4E support. Provisions are similar to a Category-4C support except for the incorporation of brackets at the level of the radially keyed shoe to resist horizontal loadings. The 1 susceptibility to radiation damage would appear to be similar to Category-4C. Figure 4.11 provides perspective and section elevation views of a PWR vessel equipped with a Category-4F support. With this support scheme, vertical and horizontal loads are transmitted through nozzle weld pads to radial keys mounted on a structural-steel ring girder that (' , is in turn mounted on short weldment pedestals. In some cases the rela- q l i tive position of the ring girder and short veldment pedestals are inver- l l ted. The distinguishing feature of this category, relative to Category-4C, is the employment of a structural-steel ring girder for transference of horizontal loadings. Otherwise, the conclusions with regard to radi-ation damage susceptibility and stress state would appear to be the l ~ same. Figure 4.12 provides a section elevation view of a PWR vessel equipped with a Category-40 support. Features of this support are similar to Category-4C except that horizontal loading is transferred directly to the reinforced concrete structure of the biological shield. The sides, as well as the bottom of the recess in the biological shield wall, provide for load transference consequently,
44 {s no significant primary tensile or bending stress fields should be generated within the support structure. A fundamental objective of the design of all PWR support systems is the incorporation of guides to establish a fixed reactor ves sel center-I line for all operating, upset, and faulted conditions. In general, the l vertical loadings on the supports are applied to ' components designed to accommodate radial thermal expension by sliding. Two exceptions to this i l method of accommodating radial thermal expansion .are the provisions of ) i pivoting columns in the Trojan supports and nested rollers in the Turkey Point supports. As noted by Hopkins,1 the NSSS and AE designers have j assumed that the imposed forces on the supports resulting from differ- j ential friction effects are negligible. The FSAR data on supports verify this observation. ORNL is not aware of any attempt to examine loadings that would result from differential friction effects. References
- 1. W. C. Hopkins, Reactor Pressure Vessel Supports for Pressurized Water Reactors and Boiling Water Reactors, Pesidual Life Assessment of Major Light Water Reactor Components -
Overview Volume 1, NUREC/CR-4731 (ECC-2469, Vol. 1), Idaho National Engineering Labora-tory, June 1987.
- 2. C. A. Knorovski, R. D. Krieg, and C. C. Allen, Jr. , Practure fough-niss of PWR Component support, NUREC/CR-3009 (SAND 78-2347), Sandia National Laboratories, Feb. 1983.
i I 45 I Table 4.1 Basic types of support for LWR vessels i
\. Types of Support Power Type No. station of L ng Shield Short name plant Skirt column 3 , , ,g tank column 1 2 3 4 5 1 Arkansas Nuclear Unit 1 PWRa 3 2 Arkansas Nuclear Unit 2 PWR 2 3 Beaver Valley Unit 1 PWR 3 4 Big Rock Point Unit 1 BWRb $ .
5 Browns Ferry Unit 1 BWR 1 6 Browns Ferry Unit 2 BWR 1 7 Browns Ferry Unit 3 BWR 1 8 Brunswick Unit 1 BWR 1 9 Brunswick Unit 2 BWR 1 10 Byron Unit 1 PWR 4E 11 Callaway Unit 1 PWR 4C 12 Calvert Cliffs Unit 1 PWR 4C 13 Calvert Cliffs Unit 2 PWR 4C 14 Catawba Unit 1 PWR 40 l 15 Cook Unit 1 PWR 4C l 16 Cook Unit 2 PWR 4C 17 Cooper Station BWR 1 18 Crystal River Unit 3 PWR 1 19 Davis-Besse Unit 1 PWR 4A l 20 Diablo Canyon Unit 1 PWR 4F 21 Diablo Canyon Unit 2 PWP 4F 22 Dresden Unit 2 BWR 1 23 Dresden Unit 3 BWR 1 24 Duane Arnold BWR 1 , 1 25 Farley Unit 1 PWR 4C 26 Farley Unit 2 PWR 4C 27 Fitzpatrick BWR 1 28 Fort Calhoun Unit 1 PWR 4C 29 Fort St. Vrain CCRe Not applicable 30 Cinna PWR 4C 31 Crand Gulf Unit 1 BWR 1 32 Haddam Neck PWR 4C
'33 Hatch Unit 1 BWR 1 34 Hatch Unit 2 BWR 1 35 Indian Point Unit 2 PWR 4F 36 Indian Point Unit 3 PWR 4F 37 Kewaunee PWR 4C 38 La Crosse BWR 1 39 La Salle Unit 1 BWR 1 40 La Salle Unit 2 BWR 1 41 Limerick Unit 1 BWR 1 42 Maine Yankee PWR 3
l 46 l l Table 4.1 (Continued) l i ( , Types of Support l Power Type No. station of Skirt b "8 * # plant column Suspension name tank column 1 2 3 4 5 43 HeCuire Unit 1 PWR 4C 44 McGuire Unit 2 PWR 4C 45 Millstone Unit 1 BWR 1 46 Millstone Unit 2 PWR 40 47 Honticello BWR 1 48 Catawba Unit 2 PWR 40 49 Hope Creek Unit 1 BWR 1 i 50 Millstone Unit 3 PWR 3 ! 51 Nine Mile Point Unit 1 BWR 1 52 North Anna Unit 1 PWR 3 53 North Anna Unit 2 PWR 3 l 54 Oconee Unit 1 PWR 1 i 55 Oconee Unit 2 PWR 1 3 56 Oconee Unit 3 PWR 1 57 Oyster Creek Unit 1 BWR 1 58 Palisades PWR 4D l 59 Palo Verde Unit 1 PWR 2 1 60 Palo Verde Unit 2 PWR 2 61 Peach Bottom Unit 2 BWR 1 i ( 62 Peach Bottom Unit 3 BWR 1 63 Pilgrim Unit 1 BWR 1 64 Point Beach Unit 1 PWR 4F 65 Point Beach Unit 2 PWR 4F l 66 Prairie Island Unit 1 PWR 2 67 Prairie Island Unit 2 PWR 2 68 Quad Cities Unit 1 BWR 1 69 Quad Cities Unit 2 BWR 1 70 Rancho Seco Unit 1 PWR 1 f; 71 River Bend Unit 1 BWR 1 ! i 72 Robinson Unit 2 PWR 4C j 73 Salem Unit 1 PWR 4F 3
.74 Salem Unit 2 PWR 4F 75 San Onofre Unit 1 PWR 47 l 76 San Onofre Unit 2 PWR 2 l 77 San Onofre Unit 3 PWR 2 !
l 78 Sequoyah Unit 1 PWR 4C 1 79 Sequoyah Unit 2 PWR 4C ! 80 St. Lucie Unit 1 PWR 2 1 81 St. Lucie Unit 2 PWR 2 l 82 Summer Unit 1 PWR 4C i 83 Surry Unit 1 PWR 3 ! 84 Surry Unit 2 PWR 3
)
i ( l i a 1
. s 47 Table 4.1 (Continued) l 4.
k -Types of Support Power Type No. station of Skirt L ng Shield Short 9 , , ,g, name plant column tank column 1 2 3 4 5 85 Susquehanna Unit 1 BWR 1 l 86 Susquehanna Unit 1 BWR 1 I 87 Three Mile Island Unit 1 PWR 1 88 Trojan PWR 4A 89 Turkey Point Unit 3 PWR 4A i 90 Turkey Point Unit 4 PWR 4A 91 Vermont Yankee Unit 1 BWR 1 92 Washington Nuclear Unit,2 BWR 1 93 Waterford Unit 3 PWR 4D 94 Wolfe Creek Unit 1 PWR 4C t 95 Yankee-Rowe Unit 1 PWR 4B 96 Zion Unit 1 PWR 4E 97 Zion Unit 2 PWR 4E l 98 Vogtle Unit 1 PWR 4C i 99 Vogtle Unit 2 PWR 4C 100 Beaver Valley Unit 2 PWR 3 101 Bellefonte Unit 1 PWR 4D 102 Bellefonte Unit 2 PWR 4D 103 Braidwood Unit 1 PWR 4E (. 404 Braidwood Unit 2 PWR 4E Byron Unit 2 105 PWR 4E 106 Clinton Unit 1 BWR 1 107 Comanche Peak Unit 1 PWR 4C 108 Comanche Peak Unit 2 PWR 4C 109 Fermi Unit 2 BWR 1 110 Crand Gulf Unit 2 BWR 1 111 Limerick Unit 2 BWR 1 112 Nine Mile Point Unit 2 BWR 1 113 Palo Verde Unit 3 PWR 2 114 Perry Unit 1 PWR 1 115 Perry Unit 2 PWR 1 116 Seabrook Unit 1. PWR 4F 117 Seabrook Unit 2 PWR 4F 118 Shearon Harris Unit 1 PWR 4F 119 Shearon Harris Unit 2 PWR 4F 120 Shoreham BWR 1 121 South Texas Unit 1 PWR 4C 122 South Texas Unit 2 PWR 4C 123 Watts Bar Unit 1 PWR 4C 124 Watts Bar Unit 2 PWR 4C
'125 Washington Nuclear Unit 1 PWR 4C 126 Washington Nuclear Unit 3 PWR 2 "PWR - Pressurized water reactor.
b 8WR - Boiling water reactor. ( # CCR - Cas cooled reactor.
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, WAf(nIF.cf y? 'I N N w gggg, gggy w .- conc sveroar etatt cosinot non onivt nozacd, ,9 , > m coat revs momeon no!!Lc Fig. 4.2. BWR (Big Rock Point) vessel with suspension supports.
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- 5. SELECTION OF LWR PLANTS FOR SPECIFIC-PLANT ANALYSIS
~ The criteria established for selection of.two plants for specific-plant analysis are ar. follows: The plants should be among those having a relatively high potential for vessel support failure as a result of radiation-enhanced propagation of flaws. Compliance with this condition was judged on the basis of the following considerat. ions: 1
- a. A potentially critical portion ' of the support should be exposed to. >
L the relatively high-neutron-flux regions of the cavity (from mid-l ( height to the end of the core).
- b. The support material should have a relatively high potential for
')i radiation embrittlement. I { ! l ( c. The portion of .' the support within the high-flux region should be subjected to relatively high tensile stresses, primary-load tensile stresses being of particular concern. Secondary tensile stresses- d (thermal and residual) are also of concern. l d. The potentially critical portion of the support should have a rels- { l tively high potential for flaws of critical size. ,! -s-Information in Table 3.5 indicates that the cavity fluxes for the l BWRs are much less than those for PWRs, and as indicated in Sect. 4, BWR 1 vessels are supported on skirts that are far removed from the bottom end ! I of the core.* Thus, BWRs were excluded from consideration. All but one j of the Babcock and Wilcox (B&W) PWRs are supported on skirts, and they also were excluded from consideration. J .) i ~* Big Rock Point is an exception but was not considered because of h its small size. -l i .l . - _ _. . . ...c . . i 61 - ) t 1 About 10% of the PWR vessels are supported on long columns and ano-ther 10% on shield tanks that extend the length of the core and thus are exposed to the maximum flux. ~ At the outset of this study each of the l PWR WSSS vendors and the Electric Power Research Institute (EPRI) were contacted informally and given the opportunity to contribute - to the j CE and EPRI responded informally with updated analyses of the study. . l long columns (CE) and shield tanks (EPRI, Stone.and Webster), consider-ing the HFIR surveillance data. The preliminary indication was that .) l critical flaw . sizes corresponding to 32 EFPY were ," acceptably" large. Thus, these supports also were excluded from consideration in this ) t study. (ORNL has not reviewed the industry evaluations in detail and thus is not in a position to comment on the industry conclusion.) The remaining PWR vessel supports fall in the "short column" cate-t gory (Sect. 4), which includes, as one extreme,.very short, stubby sup-ports (columns) that rest directly on the concrete biological shield at an elevetion above the upper end of the core, where the. flux is relativ-ely low (Fig. 5.1), to the other extreme of columns extending to about midheight of the core and resting on steel-cantilever beams. As indi-cated in Sect. 4, only Trojan and perhaps Davis Besse are of this latter type,'while Turkey Point Units 3 and 4 (identical supports) are similar to Trojan but with the steel cantilever beam located closer to the. top of the core, where the flux is somewhat less. It appeared that of all those plants in the short-column category,' Trojan and Turkey Point have the greatest potential for fracture-related failure of the vessel sup- ; i ports. Both designs include centilever beams in high-flux regions of the cavity, and both cantilevers are stressed in tension by - primary loads. + .$. 0 9 62 l 1 to 8 1 - e - s - 4 - 3 - 2 - i 1.c _ I :_ , s - i " ~ E r - , l E !
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g i 8 i ! I = i i g ~ i i $ 6=. l i = e- i i I e Z ' ! g s - ' i I i (- i = - l l l ; 2 - h i CORE HEIGHT =l 1 1 0.o1 l > l 1: ' 1 i e - i 4 - l i ) i = - l l 6 1 2 - i cong a 8 WIDPLANE 8 . 0.001 I I ' ! ' 300 300 100 0 100 200 300 .400 DISTANCE PROM CORE MIDPLANE lern) Fig. 5.1. flux and fluenceRelative axial within the variation pressure of fast vessel wall.neutron (E > 1.0 MeV) i , . I 1 63 f l i 1 l r ., 1, 6. BRITTLE FRACTURE EVALUATION OF TROJAN NUCLEAR .!' PLANT REACTOR PRESSURE VESSEL SUPPORTS 6.1 Introduction and Sumary 1 The Trojan vessel support structure is comprised of upper compo-I nents that resist horizontal loads, lower components (cantilever beams) q l that resist vertical loads, and pinned columns that transfer vertical loads from the vessel to the lower components. Only the cantilever beam is considered in this study because this component of the structure is l subjected to a relatively high fast-neutron fluence, and the applied l I loads result in primary tensile stresses. A number of potentially lim-iting locations on the lower' support beam are within the portion of the assembly embedded in the concrete biological shield. A beam-on-elastic-foundation-analysis model was used to provide a detailed definition of stresses at these locations. 1 Stress analyses of the support beam, and postulated flaws within i it, were planned in the form of a parametric study. This approach per-mitted a broad evaluation of the potential impact of premature embrit- { t tiement on the support beam structural integrity. Parameters included ! in the study were locations in the beam (3), loading conditions (4), beam support configurations (2), flaw. aspect ratios (5), flaw depths (6), and plant operating. times (3). The support beam weldsents are fabricated from A36 material. Frac-ture toughness properties of this material have been shown to be strain-rate sensitive. Materials data were taken from the open literature and converted t? a strain rate appropriate to the dynamic response of the reactor vessel on its supports. Equations were fitted to the data and a { lower-bound design curve was defined. __ _ _________ _ _ _ . . 1 64 l c ,
- 5. Irradiation-induced shifts in NDTT were calculated using correla-tions developed from the HFIR data (Sect. 3). Analysis of the plant operating profile gave radiation-induced NDTI shif ts at the critical !
location within ' he t beam ranging from 30*F at the present time to 75'F at a plant operating time of 32 EFPY. Strain rates produced by the ves-i nel system response to accident loading produce a further 60* F ' NDTT l l shift (relative to strain rates associated with slow bend tests). l 1 Design fracture toughness values at the beam locations selected for ana- l l lysis ranged from a high of 49 ksi /in. to a low of 31 kai /in. at ' 32 EFPY. Dynamic fracture toughness limits were superposed on results from i the stress-intensity-factor analyses to generate curves of critical flaw depth as a function of loading applied to the vessel support. Critical flaw depths obtained for the maximum loading condition [small-break l loss-of-coolant accident (SBLOCA)) are summarized in the following table: r" critical flaw depth (in.) Flaw length at Beam location Loading beam surface 7.48 8 32 (in.) EPPY EFPY i Interface with the SBLOCA 0.89 0.68 Full flange I reactor cavity width (16 in.) Maximum bending SBLOCA 0.88 0.61 Full flange moment location width (16 in.) Flange grout SB14CA >2.0 0.42 Full flange hole depth (2.5 in.) " Current condition (late 1988). l ( a l I 65 l 'l ~ The grout hole, which is 4.0 in. in diameter and located in the i upper flange of the embedded portion of the beam ~11 in. from the inner ' I surface of the biological shield, results in the smallest critical flaw ; size for times approaching 32 EFPY. Factors with the potential for influencing the calculated critical q flaw depths in the, beam flange are -(a) refinement of the stress- , I intensity-factor correlation to include shearlag effects,.(b) variations in the material NDT, (c) residual stress effects, and (d) inclusion of .) 1 support for the beams from concrete inboard of the inner support pedes- . tal. Pilot evaluations of these effects, included in this report, indi-f cate that their combined effect will be to produce critical flaw sizes I that are not greater than those given in the summary table. j i Fatigue-induced flaw growth makes a negligible contribution to the i {. ' above critical flaw dimensions. The flaws would thus have to have been l l present at final inspection of the support structure to constitute a l brittle fracture hazard. The question regarding the existence of flaws of critical size has I not been addressed in this study. It is appropriate to note, however, .that at the two locations other than the grout hole, the surface length ~~ ~ of the critical flaw is the full width of the flange (16 in.). It seems that a flaw of this size would be readily detectable at the time of fab-rication. Detection of the critical flaw in the . edge of the grout hole appears less likely because the flaw is smaller (0.4 x 2.5 in.), and the hole was flame cut and dressed, a process normally resulting in a rather rough surface. -r-l 66 In-service inspection for flaws in the lower support beam may be 1 impractical, particularly for the portion of the beam embedded in. the l concrete biological shield.- Inspection records, however, may shed light i on the likelihood of significant flaws existing (utility records were not available to ORNL st the time of this study). Principal uncertainties in the flaw tolerance analysis of the sup-port beams ' are related to (1) the fracture toughness data for'A36 at strain rates and temperature appropriate to'the support beam loading conditions, (2)'the correlation used for applying the HFIR' embrittlement data to the evaluation of PWR vessel support, (3) the initial NDTT for the material, (4) the operating temperatures of the support, and (5) the 1 magnitude of residual stresses. There is also a . considerable uncer- l tainty regarding the existence of flaws in the beams. 6.2 Scope and Objective The Trojan vessel support structure is comprised of upper and ! lower beam structures joined.by pinned columns (Fig. 6.1). This analysis considers only the lower beam because-(1) it is loaded so as to give some of the highest bending induced tensile stresses in the sup-port,- and (2) it is the component located closest to the core mid plane, where the fast-neutron fluence is nearly a-maximum.- At the time of preparation of this analysis, no data were available l on the geometric characteristics of potential flaws in the support ] beams. .The analysis was therefore performed as a parametric study with a range of . flav depths and aspect . ratios considered. The analysis objective was to define the critical flaw size envelope for the l . . _ _ . l 67 i l high-stress and/or high-fluence locations within the support ~ beam. It l l is intended that this envelope be used as a basis for evaluating the postulated beam flaws, when a definition of the flaw population becomes available. A critical flaw size analysis was performed for three locations on the beam. These locations and the bases for their selection were as follows.
- a. Interface with the reactor cavity liner A transverse fillet weld between the beam flange and the cavity liner exists at this loca- I tion. A potential exists therefore for weld induced cracking over I the full width of the flange. This location also experiences the maximum radiation induced dpa and dpa rate of all three locations l
considered. (
- b. Location of the beam maximum bending mcments l
the beam is supported on pedestals and is cast in concrete from the cavity liner interface back into the cavity wall. This type of support can be character-ized as a beam on elastic foundation support in which the support stiffness varies along the length of the beam. With this type of support, the peak bending moment (and the associated beam flange primary bending stress) is located a short distance back from the start of the elastic foundation support. This location is clearly one of the critical locations on the beam.
- c. Four-inch-diameter beam flange grout holest these holes are located in both the upper and lower-beam flanges, directly over the center-line of the inner support pedestal. While not necessarily located l st the point of maximum bending moment, they are close to it. This
( l u______.___ _ . _ _ _ . _ - __ 68 k fact, coupled with the loss of 25% of the flange area and the intro- ] . y duction of a geometric stress concentration (KT = 2.43) makes the ' flange hole a potentially limiting location. The fact that the flange holes.were flame cut adds to the potential for the introduc-tion of undetected flaws during the fabrication process. A further parameter . considered in the analysis was the effective-ness of that portion of the concrete inboard of-the inner support pedes-n tal inner flange. This portion of the concrete foundation would be expected to experience some of the highest elastic foundation compres-sion stresses and could potentially be loaded beyond its crushing i strength. In addition, some uncertainty exists relative to the degree , of fit achieved between the concrete and the support beam flanges at this critical location. To acconsnodate these potential effects, all analyses were run both with and without the inner concrete included in the foundation model. Four loading cases were considered in the analysis. Load sources contributing to the support loads were dead weight (DW), pipe thermal thrusts (T), the operating basis earthquake (OBE), the safe shutdown ; earthquake (SSE) and both small break and large break loss of coolant l accidants (SBLOCA, LBLOCA). Specific load combinations considered in l the analysis were as follows: Normal + upset = DW + T 1 OBE l Faulted (1) = DW + T i SSE Faulted (2) = DW + T 1 SBLOCA Faulted (3) = DW + T 1 LBLOCA 1 ') -q 69 \ Subsequent to the initiation of the analysis, ORNL was informed that NRC had approved the Trojan leak-before-break analysis, thereby eliminating the double-ended pipe-break (LBLOCA) from the plant faulted condition transient loading list. Analysis was completed for the large break LOCA condition, however, because the addition of this fourth load-ing-condition result improves the definition of load dependent variables and thus improves the accuracy of interpolation between the load cases. considered. In this connection, it should be noted that the critical flaw size is a nonlinear function of the applied stress. l The SBLOCA loading condition provided by the utility constitutes a j generic, bounding case that involves failure of auxilliary coolant lines and thus represents a more severe loading condition than often referred i to as SBLOCA. ' [ Three specific times in the life of the plant were considered in l the analysis. These were (1) start of life, (2) current condition (late 1988, 7.48 EFPY), and (3) 32 EFPY, which is the approximate present license period. Interest in the start-of-life condition stems from the strain rate sensitivity of the fracture toughness data for A36. In summary, the scope of the suoport-beam parametric study was as follows. Parameter "" # ' cases Beam locations 3 Loading conditions 4 Beam support configurations 2 Flaw aspect ratios 5 Flaw depths 6 Reactor operating periods 3 v 70 l l l I k 6.3 Loading and Operating Environment ) l All loading cases have been taken from (a) the minutes of the j s 6-29/30-88 meeting of PCE, Westinghouse, NRC and ORNL in Pittsburgh,2 PA f I
- and (b) the PCE letter of August 2, 1988.3 )
i 6.3.1 Normal plus upset loading Maximum load / support for T + DW + OBE 'I l P = 288 + $30 + 281 = 1099 kips 3 6.3.2 Faulted conditions (1) SSE Maximum load / support for T + DW + SSE = Pi ) P = 288 + 530 + 394 = 1212 kips 1 g ; 1 (2) SBLOCA ; i ( Maximum load / support for T + DW + SBLOCA = P2 1 P2 = 288 + 530 + 740 = 1558 Lips j i (3) LBLOCA j j l Haximum load / support for T + DW + LBLOCA = P3 3 I I P3 = 288 + S30 + 1625 = 2443 kips j 6.3.3 Cyclic loading l ' 1 P ,,, P ein Number of ; Event (k,t ps ) (ktps) cycles ; Refueling 818 395 40 , OBE 1099 537 400 SSE 1212 424 20 ] SBLOCA 1558 0 5 6.3.4 Operating environment Beam temperature = 90'F (~ - ] l, - - - - - - - _ - - - _ l 71 ! 1 \. 6.4 Support Beam Analysis j 6.4.1 Analytical model An overview of the support beam geometry is given in Fig. 6.1, with I l additional details of the structural members given in Figs. 6.2 and 6.3. Vertical loading is applied to lugs on the inboard end of the beam via pins and vertical links. Lateral shear loading is reacted by,a sep-arate structure located just below the elevation of the vessel noz-zies. Loads applied to the support beam thus remain vertical for all r loading cases. I Reaction to the applied loads is provided by (a) support pedestals j built into the cavity wall and (b) by interaction directly with the cav-ity wall concrete. It was considered important to model these interac-
- g. tions as accurately as possible since the modelling assumptions could A
have a significant influence on the magnitude and distribution of bend- l l ing moments in that portion of the beam located within the cavity wall. In this connection it is appropriate to note that the 4-in.-diam. holes in the beam flanges are located in that portion of the beam loca- j ted within the cavity wall. , A beam on elastic foundation model was selected on the basis of its ability to provide a detailed definition of the bending moment distribu-tion within the cavity wall. Rather than nee the conventional approach to beam on elastic foundation analysis, it was decided to model the sup-port as a series of closely spaced struts with prototypic axial stiff-ness and very low bending stiffness. This approach permitted a detailed i I representation of the distribution of support stiffness as the beam l passed over the various elements of the support pedestals. 72 The analytical model for both the support beam and the foundation elements is shown in Fig. 6.4, together with a sketch of the beam. and its support pedestals. Node points and numbers are indicated on both the model and the sketch to aid in interpretation of the analysis results. The model beam numbering system is illustrated in Fig. 6.5(a) together with details of the model in Fig. 6.5(b), showing beam and node j numbers near the location of the flange holes. Note that the flexural rigidity of the vertical foundation members is set at a very low value, thereby assuring that they take'only axial loading. j 6.4.2 Model element properties l l 6.4.2.1 Beam (solid section). See Fig. 6.2 for the section geometry. ( A g = 16 x 2.5 x 2 + 13 x 1.75 x 2 i = 1.25.5 in.2 l 7 , (16 x 183 - 12.5 x 133) 1 12 = 5487.4 in.4 ! 6.4.2.2 Beam (section with 4-in.-diam. holes) a- - \ A2 = 125.5 - 2 x 4x 2.5 = 105.5 in.2 1 2 = 5487.4 - 2 x 4x 2.5 x 7.752 - 2 x 4 x 2.53/12 = 4275.8 in.4 6.4.2.3 Concrete foundation. Assume that compressive loads from the beam flange diffuse into the concrete such that the effective ( l 73 boundary of the concrete column projects 130* from verticals drawn from the edges of the beam flange. The foundation depth is taken as the dimension from the underside of the beam to the lower surface of the support column base plate. Column width at top = flange width = 16.0 in.- l Column width at bottom = 16 + 2 e 29 x tan 30' = 49.5 in. The average concrete support column area for a unit length of sup-l port beam is thus obtained as: A = (16.0 + 49.5)/2 = 32.75 in.2/in. 1 E, = 6 x 108 lb/in.2 based upon aging in excess of 5 years l ( . (Ref. 4 and Appendix 5) Note that all foundation areas are calculated in terms of in.2/in, of beam length. Each vertical support member represents the support provided by a segment of the foundation extending equal distances on \ either side of the beam / foundation member interaction node point. The unit length support areas are multiplied by the appropriate foundation lengths to obtain the area for a given foundation support element. The concrete foundation elements can react loads only in the com-pression mode. Concrete foundation material exists, however, both above and below the support beams, thereby giving the concrete elements the capability of providing reaction loads in both directions. In the I model, this condition is simulated by giving these foundation elements both tension and compression capability. l l l 1 E_________________________*__ - 1- "" ~ 74 f k > 6.4.2.4 Flange of the front support pedestal. See Fig. 4.2(a) for geometry A = 16.0 in.2/in. E = 29.0 x 10s 1b/in.2 6.4.2.5 Front support pedestal web. The web steel and the sur-rounding concrete both provide support to the beam in this segment of 1 the foundation. The foundation area per unit length will thus be ' l expressed in terms of an equivalent concrete area with a stiffness equal to that of the steel-concrete combination I A g = A, = 2 x 1.5 = 3.0 in.2/in. l A concrete = Ac = 32.75 - 3.0 = 29.75 in.2/in. ^ equivalent =AE " ^" # ^** '! " . Ac = 29.75 + 3.0 = 29/6 = 44.25 in.2/in. 6.4.2.6 Flange section of the rear support. Section dimensions l shown in Fig. 6.3(b) for the rear support were taken from Ref. 5, pp. 1-12 A = 15.5 in.2/in. 6.4.2.7 Web section of the rear support. This is another area , where both the web steel and the surrounding concrete provide foundation support. The equivalent concrete formulation of Sect. 6.4.2.5 is used. A = (32.75 - 0.688) + 0.688 = 29/6 = 35.39 in.2/in. 75 , 1 k 6.4.2.8 Foundetion modelling sensitivity assessment. It is i 1 evident from the previous sections that a number of assumptions are built into the modelling of the beam elastic foundation support. A sensitivity analysis was performed to determine the degree to which these assumptions influence the distribution and magnitude of the cal-l culated bending moments in the support beams. Parameters included in i this-sensitivity study were (a) representation of the concrete reinfore- ) ing steel, (b) the concrete elastic modulus, and (c) representation of the concrete inboard of the inner support pedestal. Results from this study showed that inclusion of the inner concrete in the foundation model could decrease the peak bending moment in the l support beams by up to 13%. This segment of the foundation was there-fore included as a variable in the parametric analysis of the support [ beams. The other foundation modelling variables however were found to influence the support beam bending moments by less than 1%. These var-iables were not therefore included in the support beam parametric anal- l ysis. Further details of the beam elastic foundation modelling sensitiv-ity analysis can be found in Appendix 5. ~ ' 6.4.3 Beam analysis input The unit length foundation areas were first converted to foundation member areas by multiplying the unit length areas by the foundation len-gth represented by the foundation support members. This operation is summarized in Table A of Appendix 1. Elastic moduli associated with each of the foundation elements are also listed in Table A. i 1 a 76 k A unit load (200 kip) beam bending analysis was run using the Mic-rosafe -2D finite element program.5 The input data file for this analysis is given in Table B of Appendix 1. Note that each support in-corporates two beams (see Fig. 6.2). The unit load of 200 kips thus corresponds to a load of 100 kips on each beam. Analyses were run for two foundation support configurations. In the first configuration, the effective area of foundation element No. 21 l l (extending between nodes 2 and 22 in Fig. 6.4(a)) was input at a value of 108.4 in.2 (see Table A of Appendix 1). This condition represents the concrete inboard of the inner support pedestal remaining fully effective throughout all loading events. In the second configuration l l the area of foundation element No. 21 was reduced to near zero. This represents the other limiting boundary condition in which the inner con-crete segment is assumed to have failed locally in compression and so became ineffective as a foundation element. The version of the input file given in Table B of Appendix 1 represents the condition where ele-ment 21 is not providing support. 6.4.4 Beam analysis output Output files from the two beam / support configurations analyzed are c provided in Tables C and D of Appendix 1. Note that only load and moment data from these tables are used in the subsequent analysis of i postulated flaws in the beams. The beam stresses calculated by the pro-gram are not used because the beam element in the program has a solid rectangular section and this is not representative of the Trojan support beams. { P , 1 . 771 ) 6.4.4.1 Effectiveness of the inner concrete support. Table E of Appendix 1 gives the foundation loading for the unit load case (loads are given for a single beam). The load per inch in the region inboard of the inner support pedestal is obtained from the ta~ ole as 13317 ) lb/in.. The width of the lower beam flange is 16.0 in. The bearing stress produced by the 200 kip unit load case is thus obtained as: l oe = 13317/16 = 832 lb/in.2 i Bearing (compressive) stresses in the upper portions of the inner con-crete can be obtained by factoring the unit load bearing stress by F, where F = Load case maximum load / unit load ( The ultimate compressive stress for the fully aged concrete is taken as 6.0 kips /in.2 (Ref. 4). The inner concrete reserve factor is defined as RF = AH waMe stress Applied stress The inner concrete is taken to be effective in providing support for the beam for all cases in which RF 2 1.0. Using the load case maximum loads '~ from Sect. 6.3 in conjunction with the unit load concrete compressive stress results, the effectiveness of the inner concrete in the four j loading cases considered is assessed as follows: Maximum Maximum Inner I"""# ad base load F concrete concrete P (P***/200) strass RF effective (kI*p5) (832 x F) 7 DW + T + OBE 1,099 5.50 4,576 1.31 Yes (.. DW + T + SSE 1,212 6.06 5,042 1.19 Yes l DW + T + SB LOCA 1,558 7.79 6,481 0.93 Marginal DW + T + LBILOCA 2,443 12.22 10,167 0.59 No 78 , I 1 l l l {5 l Based upon the above results, the inner concrete could be consid-ered effective in supporting the beam for the seismic events but not for-the LBLOCA event. The SBLOCA event represents a transition condition in which the inner concrete will provide something less than full sup-port. All subsequent analyses, however, considered both inner concrete l l support conditions since this provides a basis for assessing issues rel- l ating to clearances between the concrete and the beam. l The unit load foundation reaction distribution is shown in Fig. 6.6. Reaction loads are seen to peak over the inner flange of the inner support pedestal. Compressive loads in that portion of the con-crete located outboard of the inner support pedestal are substantially lower than those for the inboard concrete. Some slightly higher founda-tion reaction loads occur within the web area of the inner support pedestal for the case of no inner concrete support, but these occur in a region of the foundation reinforced by the pedestal webs, and are not considered limiting. 6.4.4.2 Beam bending moments. Beam bending moment diagrams for the 200 kip unit loading condition are plotted in Fig. 6.7 for both inner concrete support conditions. Note that the presence of the inner concrete ' support has the effect of reducing the peak (Node 4) bending moment by 12.7%. The bending moment diagrams of Fig. 6.7 are of interest in that they provide a measure of the impact of simplifying assumptions that are sometimes incorporated into the analyses of partially embedded beams. The assumption that the maximum bending moment in an embedded cantilever occurs at the interface with the concrete would have produced an error ( 79 i in the estimate of the maximum bending moment of -27%. Conversely, the assumption of zero distributed support from the concrete (e.g., the beam being simply supported at the centers of its support pedestals) would have produced an overestimated of the peak bending moment of 57.5%. It would be fairly standard civil-engineering practice, however, to select the inner face of the front support column as the location of maximum moment, in which case the error in the maximum bending moment would be approximately 11% .(low). These results suggest that modelling of the distributed foundation support may in general be necessary for the eval-uation of any postulated flaws located within the embedded length of the support beam. Bending moments are required from the unit load case for input into the flaw tolerance analysis. Figure 6.7 shows the location on the bend-ing moment diagrams of the three locations selected for flaw tolerance analyses. Bending moments at these locations are as follows. Inner concrete Unit load (200 kip) Location on beam effective bending at ? lb in. Cavity liner interface N.A. 0.969 x 106 Node 4: maximum BH Yes 1.333 x 106 .. Node 41 maximum BH No 1.527 x 106 Flange hole center line Yes 1.238 x los Flange hole center line No 1.493 x los 6.4.4.3 Support stiffness. OBE, SSE, and both SBLOCA and LBLOCA loads are all the product of dynamic analyses in which the vessel sup-port stiffness is a critical element. It is appropriate therefore to check the support stiffness given by the beam on elastic foundation ( model with that used in the dynamic load analysis to verify . o 0 E % -80 ; compatibility of the models. The support stiffness used in the vessel dynamic response analysis was obtained from Ref. 2 as 1.94 x 104 kip /in. Deflections of the beam under the 200 kip unit load are plotted in 1 Fig. 6.8. The tip deflection for the case in which the inner concrete is effective is 0.00437 in. This translates to a stiffness K of: l K = 200/0.00437 = 4.58 x 106 kip /in. 3 The stiffness of Ref. 2 includes elements of the support structure. not included in the above calculation (links, pins, shoes, etc.), which would introduce additional flexibility. The calculated beam model stiffness is therefore considered to be compatible with that used in the original generation of the support dynamic loads. 1 ( 6.5. Material Fracture Toughness Fracture toughnese data for A36 in the unitradiated condition were taken from Ref,, 7. These data are reproduced in Fig. 6.9. A number of operations must be performed on the data of Fig. 6.9 to prepare a dynamic fracture toughness curve suitable for a fracture mechanics evaluation of the vessel support beam. As a prelude to per-I forming these operations, the data of Fig. 6.9 were first conver' .nto digital form and entered into a Lotus 123 spreadsheet. Figure 6.10 gives a plot of the data entered into the spreadsheet. A visual compar-ison of Figs. 6.9. and 6.10 was used to verify the accuracy of the data [ transfer. Note that the two data points from Fig. 6.9, for which precise strain rates were not available, were not used in the analysis. m a s I '81 ! l I- 6.5.1 Strain-rate correction The data of Fig. 6.9 show'A36 to be a strain-rate sensitive mater-ial. The test temperature required to achieve a fracture toughness of 40 ksi/in.' increases by -150*F as the test straining rate increases from 10 5 in./in./s to 10+1 in./in./s.. It follows therefore that the frac-ture toughness data used to evaluate the support beam flaw tolerance must be. adjusted to be compatible with the vessel support straining rate. The temperature shif t relationship required to accomplish this adjustment is given on p. 454 of Ref. 8 as follows. aT = (150 - oy ,) 50 21 , (1) where AT = temperature snift relative to static data, *F, ) [ \ \ o = room temperature yield strength, ksi, c = strain rate, in./in./s. The strain rate is calculated for a point on the elastic-plastic bound-ary for the crack tip, according to the following equation: 20 b " tE ' ( where t is the loading time and E is the elastic modelus for the material. 1 The static yield stress for the vessel support beam material was I obtained from Ref. 2 as 40.0 kei. This value has been used in all of the support beam calculations. The beam material yield stress is close to the. mean yield stress value given for A36_ in Ref. 7 (38.1 ksi) and should therefore be representative of the test material yield stress. 2a-_-,._- - _ _ - _ - - - - ---" 1 82 l ( All of the vessel support loading cases. considered involve a sub-stantial increment of dynamic load, generated by either a seismic or a loss of coolant accident event. Loading rates for this dynamic compo-nent of,the support load are dictated by the dynamic response of the j ] -1 vessel, which is dominated by_ the first rocking mode frequency of the -1 1 vessel on its supports. The vessel first mode frequency was obtained ]) from Ref. 9 as 20 cps.* ! It . is evident from examination of the cyclic loading data of Sect. 6.3.3 that the loading for OBE, SSE, and SBLOCA events - cycles j between two positive values. The loading time for those events is ) therefore one half the vessel first mode vibration period. ? I t= 1/(20 x 2) = 0.025 s { r l s \ In the case of the LBLOCA loading, the magnitude of the cyclic load j exceeds that of the combined dead weight and thermal loading. It is to i be anticipated therefore that the vessel will lift off its support dur- l ing the negative segment of the LBLOCA loading cycle, thereby extending the vessel first mode vibration period. Dynamic response time history i j data for the DW + T+ LBLOCA event are however not currently avail-I able.- The loading time for the LBLOCA event has therefore been taken as i identical with that of the other dynamic loading events. i *Information was received from PGE subsequent to completion of this ! phase of the analysis that reduced the loading frequency to 15 cps. I This change influences the straining rate dependent temperature shift by ( I 3.5'F. This small change was not incorporated into the analysis. I 83 l (, The elastic modulus appropriate to the support beam operating tem-perature (90*F) was obtained from Table I.6 of Ref. 10 as 29.0 x 106 lb/in.2 Substituting the previous values into Eq. (2) gives an effective crack tip straining rate for the support beam. 2x 40 x 103 beam = = 0.1103 in./in./s 0.025 x 29 x 106 = 0.11 in./in./s Substituting for e and o in Eq. (1) gives AT beam = (150 - 40) = 0.110 " = 75.6 'F Strain rates for the fracture toughness data of Ref. 7 vary from 10 5 in./in./s. to 10 in./in./s. Inserting these values into Eq. (1) yields the AT values in Table 6.1. The strain rate temperature adjustment in column three of Table 6.1 is applied to the test temperatures for the data of Fig. 6.10 to convert the data to a straining rate appropriate to the vessel support beam. Results obtained from the application of the strain rate correction temperature shift are plotted in Fig. 6.11. All points plotted in the figure have been adjusted to the support beam straining rate (0.11 in./in./s). 6.5.2 Data correlation and extrapolation A best fit mean curve for the data of Fig. 6.11 was obtained using the Statpad-curvefit template in conjunction with the Lotus 123 spread-sheet computer program. The program fits 25 different curves to the data and provides a correlation coefficient for each curve. The highest ( i 1 l 84 ( correlation coefficient (0.866) was obtained for a Cauchy curve with the following coefficients K Id = 1/[1.5 x 10~7 (T - 382.28)2 + o,041014] (3) 1 1 This curve is shown superposed on the data base in Fig. 6.12. Primary uses of the mean curve fitted to the data are.(a) to pro-vide a basis for extrapolating the data trend to higher temperatures, (b) to provide a base upon which to construct a lower bound curve'which l in effect defines the design plane strain fracture toughness curve for A36 at the vessel support straining rate and (c) to provide a base for construction of a plane stress fracture t ughness (K,) curve, for use where the plane strain constraint requirements are not satisfied. It is *f important therefore that the validity of the mean curve when extra-k,- polated to higher temperatures be verified. In particular, it must be shown that the slope (d KId/dT) of the mean curve approaches infinity at a temperature corresponding closely to the limit of ' applicability of linear elastic fracture mechanics. The highest temperature at which. cleavage fracture will occur is defined in Ref. 11 as follows. .a- - T + T,as o / kid + 0 , (4) where T , = the upper bound temperature limit for LEFM applicability. oy= dynamic yield stress at temperature T. kid = fracture toughness at temperature T. m 85- ., An empirical equation for computing the yield stress of a material over the range of temperatures of interest in fracture toughness evalue-tions is given in Ref. 12 as follows. 174,000 oy (T,t) =a y (0,0) + log (2 m 1010 t)(T + 459) - 27.4 (5) where TA) = dynamic yield stress expressed as a function test temper-ature and loading time, ay (0,0) = static room temperature yield stress (40 ksi), t = loading time (0.025 sec), T = temperature (*F). Yield stress values appropriate to the temperatures for the K Id i date points were calculated using Eq. (5). oy / kid values were then cal-culated and plotted as a functien of the test temperature (adjusted for strain rate) in Fig. 6.13. A mean curve was fitted to the data. Extra-polation of this curve to the point where S y/ kid = 0 gave an intersec-tion with the X axis at T = 77'F. This represents the upper bound tem-perature for cleavage fracture based on the mean of the data base. The upper bound limit is obtained in a similar manner as 100'F. It is important' to note that these temperature limits apply to unirradiated material. Extrapolation of the best fit mean data curve to a. temperature of f 75*F is shown in Fig. 6.14. The slope of the curve can be seen to in-crease rapidly at T = 75'F. The calculated slope at T = 77'F is 1.6 ksi/*F. The best fit mean data curve is seen to be consistent with l the limits defined in Ref. 11 and is thus considered validated as a basis for constructing the fracture toughness design curve for A36. 86 b It is important to recognize that the extrapolated kid curves, while providing a valid base for constructing the final' fracture tough-ness design curve, are not suitable for direct use in a! fracture tough-ness evaluation. This limitation is due to the' fact that the constraint conditions at the crack tip in'the material change from plane strain to plane stress as the ratio of fracture toughness to dynamic yield stress increases. The use of the extrapolated curve in deriving the fracture l tough' ness design curve is discussed further in the following section. . l '6.5.3 Plane strain-plane stress transition Irwin has defined the upper-bound . limit for pure plane-strain behavior at a crack tip in terms of a 8 parameter as follows (Ref. 13) ~ l 'g t I g =1 Ic $ g,4 m (6) Ic B o
- M B = material thickness v .
K ye a plane-strain fracture toughness oy , = dynamic yield stress The data of Fig. 6.9 are also given in Fig. 4.12 of Ref. 13, with 8 = 0.4 intercepts superposed on the fracture , toughness . curves. The s e~ intercept s indicate that the limiting kid value for plane-st ** conditions at the crack tip is 48 kei/E . Above this value, adjustment. must be made for the onset of plane stress behavior at the crack tip. The . relationship between the plane strain fracture toughness K Ic and the plane stress fracture toughness Ke is given in Ref. 12 as fol-lows. ( K2=gj (1 + 1,4 x glc) (7) 1 s, . 87 ( Equation (7) has been demonstrated to be accurate only for cases in which differences between K ye and Ke are small. In the absenck of plane stress fracture toughness data, however, Eq. (7) was used to predict j plane stress fracture toughness behavior. The ' plane stress fracture toughness curve is shown superposed on the data base, together with the plane' strain fracture toughness curve in Fig. 6.15. Note that the difference between the L best fit plane strain and the estimated plane stress curves at the 8 = 0.4 limit is small (2 kei 5 ). The two mean curves'were blended together, starting ct the point KIc = 48 ksi5. on the plane strain fracture toughness curve. Note that the K Ic value of 48 kei 5 . in corresponds to a strain - rate adjusted temperature of -15'F. The blending curve is shown in Fig. 6.16. [ \ 6.5.4 Fracture toughness design curves A lower bound plane strain frseture' toughness curve was constructed by applying a 50*F temperature shif t to the plans strain portion of the mean value curve.* The plane stress extension of the lower bound curve was then constructed by applying Eq. (7) to the lower bound plane strain curve for temperatures above -15'F. The resulting curve is shown in Fig. 6.17,. superposed on the data set. Note that the lower bound curve passes through the lowest data point. *The 50
- F shift is suggested by examining the shift of the ASME lower-bound K ye or KIR curve from the mean curves of the corresponding data sats (Flaw Evaluation Proceduress ASME Section XI, EPRI NP-719-SR,
( August 1978). l 88 . Requirements for impact testing of materials . intended for Mvice in nuclear component supports are given in subsection NF of BE6tio6 III of the ASME Boiler and Pressure Vessel Code (Ref.14). These require-ments were used in a further validity check on the lower bound curve of Fig. 6.17. I Impact testing requirements- for materials intended for service in component supports are defined in Paragraph NF-2311 of Ref.14. . Impact testing is not required when the support service temperature exceeds NDTT + 30'F. NDTT for A36 plate is given in Table NF-2311(b)-1 as 40*F, giving a lower bound service temperature (for unirradiated material) of 70*F for exemption from the impact testing requirement. In the follow-ing discussion, this temperature is interpreted as a lower bound temper-ature for design without consideration for brittle fracture. The 70*F temperature limit from Ref. 14 is shown superposed on the lower bound fracture toughness curve in Fig. 6.17. The lower bound curve is seen to be increasing rapidly as it approaches the Ref. 14 lim-w. iting value. Intersection of the lower bound curve with the 70*F limit - occurs at a kid value of 120 kai 5 . Analysis results presented in Sec-tion 6.7 (Table 6.17) show critical flaw depths to be increasing rapidly
- I as tiie critical stress intensity factor increasu by 10 kaiS. from 70 kei M . to 80 kai 5 . Critical flaw depths at a K Id value of 120 ksi 5 . would be predicted to be very large and therefore compatible
.5 . with elimination of the requirement for materials impact testing. The .:n ' lower bound curve of Fig. 6.16 is therefore seen to be compatible with 4- /ifi the interpretation of the fracture toughness behavior of unirradiated j . ser i A36 material incorporated into the materials impact testing requirements ( of Ref. 14. l 4 j n j l l 89 ' ( The final step in construction of the fracture toughness design curve was - to replot the curve in terms of T-NDTT. NDTT was obtained f rom Ref. 7 as 28'F. The resulting f racture toughness design curve is given in Fig. 6.18. 6.5.5 A36 NDTT variation NDTT for A36 plate used in this analysis was taken from Ref. 5 as 28'F. Heat to heat variations in NDTT for a range of materials used in PWR component supports were evaluated in an extensive survey conducted by Sandia National Laboratory (Ref.14). Results quoted in Ref.14 for carbon-manganese steels, including A36, give the mean NDTT value as 22'F, and a population standard deviation (c) as 13*F. The (mean + 1.3 c) value for NDTT, corresponding to a 90% confidence limit, is given as 39'F. An increase of 11*F in NDT must therefore be applied to the (' results given in this report, in order to obtain critical flaw depths i I corresponding to a 90% confidence limit on NDT. A carpet plot giving critical flaw depths as a function of (a) T - NDTT and (b) support loading, is given in Fig. 6.49 for the beam flange grout hole location. This plot can be used to rapidly assess the impact of applying the 90*F confidence limit temperature shif t (11'F) on the calculated critical flaw depths. 6.5.6 Supporting analysis A summary of analyses performed to construct the A36 fracture toughness lower bound curves is provided in Tables 6.2 through 6.5. ( 90 6.5.7 Reactor support /bridae structure comparison A36 is widely used in bridge construction where it has performed satisfactorily. It is of interest therefore to compare the service con-ditions appropriate to bridge application with.those encountered in the reactor vessel support application. In particular, the effect.of the differing dynamic response characteristics of the two structural systems i will be considered here. Bridges are typically relatively flexible' structures with low nat-ural frequencies. Peak straining rates in bridge structures during response to impulsive loading are limited by their-low natural frequency.. to relatively low values. Rolfe and Barsom (Ref.13, p. 456) place the upper bound crack tip straining rate for bridge structures under impul-sive loading conditions -at 0.001 in/in/s. This . straining rate is approximately two orders of magnitude slower than that encountered in the vessel support beam application. The temperature-adjustment results of Table 6.1 show that this straining rato differer.ce produces a temper-ature shift of approximately 40'F. Design stress-intensity-factor val-ues at a given temperature for the support beam application are thus lower than the corresponding values for a bridge application. t 6.5.8 Fracture toughness data evolution Recently published analyses of shallow-crack: and deep-crack l A36 CTOD fracture-toughness specimens indicate that . higher fracture, , toughness can be generated in specimens with small crack depths (Ref. 16). This development is noted as a pointer to considerations for possible inclusion in any further evaluation of the support beams. The {- results of Ref. 16 are not incorporated into the current analysis. - ____ _ _ _ _ - _ _ _ _ _ - - _ _ - _ . _ - - _ l q . + 91 ! 6.6 Radiation-Induced NDTT Shift . l 1 NDTT chift (ANDTT) values were calculated for three governing loca-tions on the support beams for both the current condition existing at l i Trojan and for 32 EFPY. The steps in the analysis were as followst. l l
- a. Determine the core midplane dpa rate profile into the reactor cavity ;
wall for the Trojan initial core configuration. Select dpa rates for the three governing locations.
- b. Catalog the Trojan operating history to date, and determine the pro- )
jected operating condition in terms of the ratio (R) of the reactor cavity dpa rate to the cavity dpa rate associated with the first core. Calculate the time averaged dpa rate ratios for both prior i operations and projected lifetime operations.
- c. Using the dpa rate ratios from step (b) in conjunction with the I cycle 1 dpa rates from step (a) calculate effective dpa rates and total dpa's for both the current condition and the projected end of O
life condition. ! l
- d. Using the dpa and dpa-rate data from step (c), enter Fig. 3.4 to !
obtain the ANDTT for each of the three ' governing . locations on the ! i beam (see Fig. 6.19). l tore midplene dpa data were obtained from Westinghouse (Ref. 17) l and are plotted in Fig. 6.20. Note that the data points plotted in j l i Fig. 6.20 are the result of calculations normalized to a measured dpa l l rate for Diablo Canyon Unit 2. The dpa rates for the points of interest , l on the support beams were obtained ast . . .. . j 92 (, Beam location dpa/sec. Cavity inner surface inter- 2.79 x 10 12 l face l Wode 4: Maximum bending at. 2.65 x 10~13 Flange hole center line 1.09 x 10~13 Trojan operating history data and planned future operating condi-tions were obtained from Ref. 18. These data are reproduced in Table 6.6. The data indicate the following dpa ratios Operations to date R = 7.593/7.48 = 1.015 ~a ture operation = 0.65 Results from the analysis of steps (c) and (d) are summarized in Table 6.7. At the present time Trojan has operated for less than 25% of its presently licensed life (32 EFPY). The maximum dpa in that portion of the support beam which projects into the reactor cavity has reached a level where a ANDTT of 75*F is predicted. ANDTT at this location is predicted to increase to 180*F by 32 EFPY. Corresponding values of A NDTT at the beam flange grout hole center line are 30*F at the present time and 75'F at 32 EFPY. T - NDTT values were calculated for both the current condition and 32 EFPY for a range of support beam operating temperatures. The support beam 6perating temperature has been set at 90*F (Ref. 2), but a range of support beam operating temperatures was included in the analyses to pro-vide some insight into the impact of potential refinements in the defi-nition of this parameter. The T - NDTT values were superposed on the fracture toughness design curve of Fig. 6.18 to obtain fracture toughness design values for each of the critical locations on the support beam for initial, current ~ 1 1 93 k ,, and 32-EFPY conditions. The resulting fracture toughness design values are summarized in Table 6.8. In the case of the beam flange grout hole, additional analyses were performed in order to define the design fracture toughness as a function of time between the current time (7.48 EFPY) and 32 EFPY. Results from this additional analysis are suannarized in Fig. 6.21 ' Supporting calcu-lations for Fig. 6.21 are summarized in Table 6.9. Note that while the design fracture toughness is a non linear function of operating time, ANDTT varies with time in a roughly linear manner -between 7.48 EFPY and 32.0 EFPY. An increase in the beam operating temperature can thus be used to offset additions to the reactor operating period 1 using the following relationship ! ( A(EFPY) EFFECTIVE = A(EPPY) ACTUAL ~ ~ ' ' ( where T = operating temperature at the location of concern on the support beam (*F) 6.7 Critical Flaw Depth Analysis . .e - Stress intensity factors were calculated using the methods and data of Article A-3000 of Ref. 19. Lotus 123 spreadsheet templates were pre-pared to accept input from (a) the unit loading beam bending analpis, (b) the load case definitions (c) the flaw shape parameter Q, and (d) the flaw membrane and bending stress correction. factors M ,and M b . Tne spreadsheets then calculated the flaw tip stress intensity factor K 3 for a range of flaw depths (a) and flaw aspect ratios (a/1). f 4 + o _o_. o++ .-+-*-4-~ _...o 1 94 l l l It should be noted that the flaw tip stress intensity factors are a non linear function of the applied loading.' The analysis results can -not therefore be directly factored for application'to other loading con- I ditions. The analyses were however performed for four different peak loading conditions. A plot of results.for these loading conditions pro-vides a means of estimating the flaw'tip stress intensity factor for intermediate loading conditions. Results from the flaw tip stress intensity analyses were plotted'as a function of flaw depth for each of the flaw aspect ratios evaluated. Fracture toughness values determined in Sect. 6.6 (see Table 6.8) were then superposed on the Ky plots. Critical flaw depths were obtained from.the intersection of the Ky and fracture toughness curves. . Finally, the Critical flaw depths calculated for each of the four loading condi-tions evaluated were combined in a single plot, giving Critical flaw depth as a function the maximum vessel support load. 6.7.1 Cavity linear interface A sample of the spreadsheets' used to calculate stress intensity factors at this location is given in Table 6.10. Engineers bending the-- ory was used to cal'culate a linear distribution of bending' stress in the ~n' overail tieam section. The membrane and bending elements of stress in the upper . flange were then calculated. Using these results, flaw tip stress intensity factors were calculated for each of the flaw depths 4 using the equation of Ref. 19, which is reproduced below as K g = o,M ,/I /a/Q + obbN /T /*/S ' (9} - _ _ _ _ _ _ :_ 1 _ 95 .. 1 I where o,, ob = membrane and bending stresses, poi, in accordance with ! A-3200 a = minor half-diameter, in., of embedded flaw; flaw depth for i surface flaw .Q = flaw shape parameter as determined from Fig. A-3300-1 using (a,+ o )/ b ys and the flaw geometry M, = correction factor for membrane stress (see Fig. A-3300 t for subsurface flaws; Fig. A-3300-3 for surface flaws) q Hb = correction. factor for bending stress (see Fig. A-3300-4 for subsurface flaws; Fig. A-3300-5 for surface flaws) It should be noted that considerations of the effectiveness of the i 1 ) inner section of the concrete foundation in providing support to the ' beam do not influence either the distribution of bending stresses or the. calculation of flaw tip stress intensity factors at the cavity liner l interface location. j i A summary of flaw tip stress intensity factors calculated for the l 1 support beam flange at the reactor cavity liner interface is given in { Table 6.11. The analysis sheets from which these data were taken are j provided in Appendix 2. Plots of the flaw tip stress intensity factors for each of the load combinations considered are given in Figs. 6.22 through 6.25. The 32 EFPY kid value derived for this location on the beam is 31 ksi/in.. Critical flaw depths corresponding to 7.48 and 32 EFPY of operation are j shown plotted in Figs. 6.26 and 6.27. The critical depth under SBLOCA l I loading for a uniform flaw extending across the entire 16 in. width of I L v ; 96 i ( '. the beam flange (a/1 = 0) is seen to range from a maximum of 0.89 in. at 1 l l the current time (7.48 EFPY) to a minimum of 0.68 in. at 32 EFPY. The a/1 = 0 flaw is common to all locations evaluated and therefore provides ) 1 a useful measure of the relative severity of the brittle fracture poten- ! I tial at each location. l Note that the flaw tolerance at the, cavity liner interface increa- ) ses rapidly as the flaw aspect ratio increases above zero. Critical flaw depths were not calculated for cases in which a > t/2. A non-heat-treated fillet weld exists at the interface between the support beam flange - and the reactor -cavity liner. Residual stresses from the weld have not been incorporated in the flaw tolerance analy-sis. This approach was taken based upon-.the assumption that the criti-cal flaw depths would be such as to be beyond the zone of any signifi-cant stresses introduced by this relatively small (the cavity liner thickness is 0.25 in.) fillet weld. The critical flaw depths calculated 6 for this location appear to justify this assumption. 6.7.2 Maximum bending moment location (Node 4) Node 4 is located within the concrete embedment at a point 16.5 in. from the point of application of the load. The beam geometry at this location is identical with that at the cavity liner interface and the method of analysis is thus very similar to that described in Sect. 6.7.1 (only the input parameters are different). Examination of Fig. 6.7 and Table 6.8, however, shows the existence of a substantial difference between the two locations. The bending moment diagram of Fig. 6.7 shows a 57.6% increase (for the case where the inner concrete does not provide support for the beam) at Node 4 relative to the cavity liner interface, { -v , 97 i~ whereas the fracture toughness data of Table 6.8 shows only a 32%' increase between the same two points at 32 EFPY. It is to be antici-pated, therefore, that critical flaw depths at the Mode-4 location will be smaller than those at the cavity liner interface. To balance this view, however, it must be recognized that no attachments or processes with known crack producing potential exist at Mode 4, whereas a fillet veld between the cavity liner and the beam flange exists at the cavity liner interface. A sample spreadsheet .used in the analysis of the Node 4 stress intensity f actors is given in Table 6.12, and a summary of the stress-intensity-factor results is given in Table 6.13. Appendix 3 contains ] all of the spreadsheets used in the analysis of the Mode-4 location. l Note that the results for Mode 4 are influenced by the effective-ness of the internal concrete in providing foundation support for the . beam. The results of Sect. 6.4.4.1 suggest that the inner concrete can be assumed effective for seismic loading cases but not for the double l ended break (LBLOCA) case. The small break (SBLOCA) case represents a transition condition. Plots of the flaw-tip stress-intensity factors are given in Figs. -6.28--6.31 for the case in which the inner concrete provides no foundation support and in Figs. 6.32--6.35 for the case where the inner concrete provides support. Crossplots of the critical flaw size as a function of the applied vessel support loading are given in Figs. 6.36 and 6.37 for the cases of no inner concrete support and in Figs. 6.38 and 6.39 for the case of fully effective inner concrete support. O n .___m_ _ . _ . _ _ . _ _ _ _ _ _ _ _ . _ _ _ _ _ _ _ _ m .A -.m _ 98 i ( Critical 32 EFPY flaw depths for SBLOCA loading for the a/1 = 0 case, with no inner concrete support, range from a maximum of 0.88 in. for the current condition (7.48 EFPY) to a minimum of 0.61 in. at 32 EFPY. With the inner concrete providing foundation support, these values increase to 0.97 in. and 0.65 in., respectively. Critical flaw depths at the maximum bending moment location are thus similar to those calculated for the cavity liner interface location. The trend of rap-idly increasing critical flaw depths as the flaw aspect ratio increases above zero is also observed in the Node-4 analysis results. 6.7.3 Beam flange grout-hole location The grout-hole location in the beam flange differs substantially from the other locations considered in that (a) the crack front of the flaw of concern is parallel to the axis of the hole rather than parallel to the free surface of the flange (as was the case in the previous two locations considered) and (b) the presence of the flange hole introduces a substantial stress concentration at the surface of the flange hole. Feature (a) requires an evaluation of the beam flange geometry and the postulated flaw characteristics prior to selection of an expression for calculating the stress intensity factor (K y). Feature (b) requires that the non linear stress distribution in the flange adjacent to the grout hole be factored into the stress intensity factor calculations. 6.7.3.1 Flaw-tip stress-intensity-factor equations. Solutions exist in the literature for tre calculations of stress intensity factors at the tip of cracks emanating from round holes in both finite width and infinite width plates. The solution used in this analysis was taken from Ref. 20 and has the following form: 99 l k.. Ky = c/E = F (10) where o = the uniform far field tensile stress on the plate, kei, i J a = the half width of a symmetrical radial crack, including the hole radius, and' F =a factor reflecting both the concentration of stress adjacent to the hole and the finite width of the plate. The factor F is obtained from the curves of Ref. 20, which are repro-duced in Fig. 6.40. Note that factor F is a function of three vari- l ables, one of which (h/b)-is defined as follows . l h , distance from crack plane to the uniformly loaded plate edge g) b plate half width. The geon.7:ry of the beam flange and the postulated flaws fits the conditions of the Ky solution exactly, but the loading condition does f not. Tensile end loads in the beam flange are generated by shear flows at the flange to shear web interface. This leads to (a) a non-uniform axial distribution of end load along the length of the flange, and (b) shear stresses in the plane of the flange due to the web location. A comparison of the two plate loading conditions is given in Fig. 6.41. The effect of the departure from prototypic loading on the calcu-lated'Ky' values is not known at this time, but a potential exists for the calculated Ky values to be non-conservative.* .This fact should be recognized when evaluating the analysis results, particular1y' for cases E involving F values taken from the knee of the curve of Fig. 6.40. *1. ate in the preparation of this report, both cases described in Fig. 6.41 were calculated using two-dimensional finite-element analysis. These analyses checked the " exact" solution for problem I 100
- k. 6.7.3.2 Flaw tip stress intensity factors. A sample spreadsheet used in the analysis of the beam flange hole radial flaws is given in Table 6.14. The mean stress entry in this table is based on the full section properties at the hole center line (i.e., the reduction in sec-tion modulus due to the loss of flange area was not included). This 1
approach - generated a pseudo far-field stress corresponding to that defined in Fig. 6.40. Note that the beam flange and load at.the plane of the hole center line corresponds exactly with that given by the cor-rect section properties (i.e., the bending so ent calculation includes the effect of the hole). l A complete set of spreadsheets for the flaw tolerance analysis of the beam flange holes is provided in Appendix 4. A summary of flaw-tip stress-intensity factors for the beam flange hole is given in ( Table 6.15. l Plots of the flaw tip stress-intensity factors for each of the load combinations considered are given in Figs. 6.42 through 6.45. Super-position of the design fracture toughness values from Table 6.8 onto these curves defined limiting flaw depths for each load case consid- , ered. These limiting values are susunarized, along with critical crack lengths for the other locations analysed, in Table 6.16. 6.7.3.3 Lifetime critical flaw depth analyses. The beam flange, hole location was selected for a study of the lifetime variation in, critical flaw depths on. the basis of its having _ the smallest critical flaw size at the end of life. NDTT shifts correspon, ding to four inter-mediate times in the remaining plant life were calculated using the method of section 6.6 and translated into design dynamic fracture 4 101 l '. toughness values using the design curve of Fig. 6.18. A plot of the resulting lifetime variation in the design fracture toughness is given in Fig. 6.21. Fracture toughness design values corresponding to intermediate plant operating periods were then superposed on the curves of Figs. 6.42 l 1 through 6.45 to generate critical flaw depths. Plots of the resulting critical flaw depth as a function of reactor operating period are given ; in Fig. 6.46 for the SBLOCA loading condition. Note that the reduction in critical flaw depth between the current time and 20 EFPY into the design life occurs as an approximately linear function of the plant i equivalent full power operating period. ' Plots of critical flaw depth as a function of ' the maximum dynamic g I load on the vessel support are given in Figs. 6.47 and 6.48. The data used to construct these plots are summarized in Table 6.17. l l The flaw considered in the beam flange hole analysis is a constant depth, through flar.ge flaw, and.thus corresponds to the a/1 = 0 case analyzed for the other two critical locations on the support beam. Note that the average slope of the curve for 32 EFPY between the DW + T + OBE and DW + T + SBLOCA loading cases is -2.6 x 10-3 in./ kip (see Fig. 6.47). The corresponding slope for the cavity liner interface location is obtaired from Fig. 6.27 as -3.3 x 10 6 in./ kip. The slopes are seen to differ by a factor of 7.88. Benefits to be derived from a reduction in the beam flange stress are thus greater at the beam flange hole location than at the other locations. Factors contributing to the 7.88 &a/ kip slope differential are identified and discussed in Section 6.8. 102 5, 6.7.3.4 The impact of NDT variations. Variations in reported values of NDTT for A36 were reviewed in Section 6.5.5, and it was shown that a further temperature shift of 11*F must be applied to the results ; calculated in this report in order to encompass the 90' confidence limit on the NDTT population. The carpet plot of Fig. 6.49 has been con-structed to permit a rapid evaluation of the impact of this temperature ! I shift on critical flaw depths at the beam flange grout hole location, j for the governing case where the inner concrete provides no support to the beams. It should be noted that the plot of Fig. 6.49 can also be used to assess the impact of the departure of the beam operating temper-ature from the 90*F value used in the analysis. Analyses used to construct the plot of Fig. 6.49 are summarized in l Table 6.18. As an example of the use of Fig. 6.49, consider the case of the 32 EFPY critical flaw depth for the SBLOCA loading condition. This value (0.42 in.) is shown entered on the SBLOCA load curve in Fig. 6.49 at a T-NDTT value of -13*F as point A. (See Table 6.8 for the origin of the -13*F.) Application of the 11'F temperature shift required to encompass the 90% confidence limit on NDTT moves the calculated T-NDTT value to -2 .* F, corresponding to point B on Fig. 6.49. The critical 1 flaw depth corresponding to point B is seen to be 0.31 in. . Application of the 90% NDTT confidence limit temperature shift (11*F) thus reduced the critical flaw depth by 0.11 in. In a further example, if the beam operating temperature is 1 increased by 10*F to 100*F and the temperature shift associated with the 90% NDTT confidence limit is neglected, then T-NDTT changes from -13*F (_ 103 . to -3
- F. Point C on Fig. 6.49 corresponds to the -3*F T-NDTT value.
The corresponding critical flaw depth is obtained as 0.6 in., an increase of 0.18 in. l l 6.8 Evaluation of Results i i Flaws can be present in the support beams due to cracking induced- ! l l by the fabrication processes and/or cracking induced by fatigu'e load- ! ing. In both cases the flaws will grow due to the effect of cyclic loading on the supports. The magnitude of the flaw growth is of concern in the evaluation in that it contributes an increment to the flaw dimen- ' sions which would not have been present during final inspection of the support beams. 6.8.1 Flaw growth due to cyclic loading l ( Cyclic crack growth data were taken from Ref. 7, and are reproduced in Fig. 6.50. The data for veldsents were selected for use in this I l evaluation since (a) the fillet weld at the cavity liner interface and j the flame cut surface of the grout hole will both be in a metallurgical condition which is representative of weld metal /HA2 and (b) crack growth ! l rates for the welded specimens were higher than those for parent mate- ! ~ rial.' A~ lower-bound curve was added to the data of Fig. 6.50 and was ! i used to generate the following equation. I l da/dW = 1.158 x 10 33 m AK5*837 (12)* l I l i *0ther formulations of the equation for crack growth in A36 have j ( been evaluated (Ref. 21). While these alternative formulations do give j slightly different crack growth predictions, the differences are too ! small to influence the conclusions drawn here. # i 104 ( \ A definition of the cyclic loading to be included in the crack growth analysis is given in Sect. 6.3.3. Critical flaw dimensions are given in Table 6.16 for the SBLOCA loading. These flaws all represent conditions for which the flew tip stress intensity factor equals the local value of K In order to Id. rapidly obtain an upper-bound order-of-magnitude assessment of flaw growth due to cyclic loading, the flaw tip stress intensity factors for all cycles in a given loading condition were calculated using the critical flaw geometry. Results were then sumuned to give a cumulative flaw growth for the total of all of the support beam duty cycle events. Results from the upper-bound flaw growth analysis are summarized in Table 6.19. Flaw growth is seen to be insignificant relative to the i dimensions of the critical flaws. This result leads to the conclusion that presumably the critical flaws identified in Table 6.16 will exist in the support beams only if they escaped detection during final inspec-tion. 6.8.2 Interface with the reactor cavity liner Critical flaw depths given for this location in Table 6.16 are seen to increase rapidly as the flaw aspect ratio (a/1) increases, reaching values in excess of t/2 for a flaw aspect ratio of 0.2 (SBLOCA load-ing). The significance of this finding is that a substantial length of flaw would have been visible on the surface of the component at final inspection. The specific lengths of relevance are 16.0 in. for the a/1 = 0 case and 9.7 in, for the a/1 = 0.1 case. Detection of surface { indications of this magnitude presumably is well within the capability of a trained inspector using visual inspection techniques. Dye m i e 105 ( penetrant or magnetic particle inspection presumably would have_ detected such large surface indications. It is~ considered unlikely, therefore, i that flaws with these critical ' depths and aspect ratios could exist in I the s'.p po r t beams-at the liner interface location. It is recommended that the Trojan inspection procedures (and inspection records if they are available) be reviewed to confirm this tentative conclusion. l 6.8.3 Maximum bending moment location (Node 4) Concip 3 ans reached for this location are identical with those given above for the cavity liner interface. Surface lengths'of flaws having the potential to produce brittle fracture under SBLOCA loading were all in excess of 5.0 in. In this instance there is no weld pre-sent, and.therefore a transverse (normal to the rolling direction) sur-face crack of this magnitude presumably would have been readily apparent to the inspe: tor. 6.8.4 Beam top flange grout hole location The stress concentration produced by this large hole (4-in. diar} in a highly stressed portion of the beam upper flange results in a crit-ical flaw depth for the SDLOCA event of only 0.42 in. (if the inboard concrete is not effective). It is considered unlikely that any prac- -a ~ tical ins' pectic" technique applied at the time of fabrication could have reliably detected so small a flaw in and adjacent to the flame cut sur-face. The small site of the critical flaw at the flange grout hole makes it imperative that a reliable characterization be obtained of the flaws present in this region. The fact that this region of the support beam ( is embedded within the cavity wall makes direct inspection impossible. j It is suggest ed therefore that a generic definition of the population of l ., , . . . - _. __ .-x -n. n 106 k flaws adjacent to the surface of flame cut holes in plates of similar thickness be compiled using data from (a) the manufacturers of flame cutting equipment, (b) destructive examination of any available flame cut off-cut material, and (c) duplication of the flange hole flame cut-ting operation and destructive examination of the finished product. 6.8.5 Compatibility check on the stress-intensity factor equations Two separate sources were used for the stress-intensity-factor equations used in this analysis. Postulated flaws in sections of the beam flange ren te from discontinuities were analyzed using equations and factors taken from Section XI, Appendix A of the ASME Boiler and Pressure Vessel Code (Ref. 19), whereas flaws at the beam flange hole ' location were analyzed using equations from the Stress Analysis of ( Cracks Handbook, p. 19.4 (Ref. 20). The purpose of the check summarized in this section is to verify that these equations, as interpreted and utilized in this report, give compatible results. The stress-intensity factor equations taken from Ref. 19 were for a single-sided surface flaw, whereas those taken from Ref. 20 were for symmetric radial flaws emanating from a circular hole in a finite .o pl a t e .- In order to provide a basis for the compatibility check, the beam flange hole case was divided on its axis of symmetry and idealized as a plate with a single-sided flaw. Plate " thicknesses" were defined as the surface-to-surf.2ce distance in the direction of crack propaga-tion. This definition gave a alue of t = 2.5 in. for the locations remote from the hole and t = 8.0 in. for the hole location. The value 8.0 in. (half the flange width) is used rather than the ligament width (6.0 in.) because the stress-intensity factor equations of P.ef. 20 define the crack length to include the hole radius. 107
- k. With the above analytical models defined, the criteria for compat-ibility were defined as follows. When applied in a compatible manner on a finite width plate, the two equations must predict compatible stress-intensity / flaw-depth relationships. This requirement specifically excludes cracking regimes in which non-linear behavior would be anticipated such as (a) the area of concentrated stress immediately adjacent to the hole and (b) secondary bending effects resulting from large-scale penetration of single sided cracks. This latter constraint on the compatibility -comparison was necessary since one of the models (the beam flange hole model) is constrained from generating secondary bending moments while the other model is not.
The compatibility check was made using the a/1 = 0, SBLOCA results fer the cavity liner laterface and the beam flange hole. These results d' are reproduced in Table 6.20, column 7. The stress-intensity factors of Table 6.20 were produced by differ-ing flange stresses, since the flange bending moment varies alon) ehe length of the beam. In order to provide a basis for comparison, there-fore, the stress-intensity factors were normalized by dividing by the local flange far fleid membrane stress (column 8). The resulting param- ~ eter has'the units of kai/in!/ksi or /in. (column 9). A similar normalizing procedure was applied to the flaw depth to beam thickness ratio. In the case of the beam flange hole location, both the plate thickness "t" and the flaw depth "a" are pseudo values, having the geometric configuration defined above. The resulting dimen-sionless "a"/"t" values are given in column 5 of Table 6.20. The data of columns 5 and 9 of Table 6.20 were then plotted to pro-duce the curves of Fig. 6.51. The linear portions of those curves are 1 108 k clearly predicting similar relationships and the equations, as utilized in this report, are therefore considered compatible. Note that the por-tion . of the cavity tirer interface curve extending from a/t = ' d4 to 1 a/t = 0.5 was defined by a single point and would not be espected to be ' linear over its entire length. 6.8.6 Load sensitivity comparison 'It war noted in Section 6.7.3.3 that the ate of decrease of allow-i able flaw size corresponding to a given increase in support load dif-1 l fered by a factor of almost 8 between the beam flange hole and the cavity liner interface locations. Higher rates of change were obtained l st the beam flange hole location. The representation of the flaw depth 1 vs. stress-intensity factor relationship given in Fig. 6.51 provides a j basis for identifying the parameters responsible for the bulk of the l l factor of 8 rate change. Stress-intensity factors for the beam flange hole and cavity liner interface locations are plotted as a function of the dimensionless flaw depth parameter "a"/"t" in Fig. 6.S2. Curves are plotted for both the DW + T + OBE and DW + T + SBLOCA cases. The linear portions of the curves have differing slopes, reflecting (a) the differing beam flange membrane stress levels at the two locations for a given support load due . to variation of bending moments along the beam and (b) increases in sup- I port load between the W + T + OBE and the W + T + SBIDCA cases.
- l End of life design stress-intensity factor lines were superposed on the two sets of curves and the dimensionless flaw depth increments A("a"/"t") between the W + T + SBLOCA and the W + T + OBE curves were determined from the intercepts as follows.
m 1109 .t" (. Location A("a"/"t") Cavity liner interface 0.06 .. Beam flange hole 0.144 In the case of the cavity liner incerface, both the "a" and "t" values correspond directly to the actual flaw depth and the actual beam flange thickness. The flaw depth increment corresponding to a A("a"/ "t") value of 0.06 on the 2.5 in. thick flange can thus be obtained as: l i cLI = 2.5 x 0.06 = 0.15 in. Aa At the beam flange hole location, the pseudo crack length "a" cor- .j responds to R + a, where R is the hole radius and a is the depth of a uniform flaw measured from the surface ' of the hole. The thickness of the pseudo beam measured in the direction of propagation of the crack is ] j in this case the half width of the plate (8.0 . in. ) . The flaw depth increment at the hole location corresponding to the A("a"/"t") value of 0.144 is thus obtained as follows l A(R + a) = 0.144 x 8.0 = 1.152 l t i Since the hole radius R is constant, ths value AaBFH is equal to l l 1.152. The ratio of the critical flaw depth change at the cavity liner . l interface ~ (CLI) and the beam flange hole '(BFH) locations for the same j increment in applied loading is thus obtained as: 1 ' A*BFH 1.152 = .68 As CLI ".0.15 This value is ' very close to the value noted in Section 6.7.3.3. The rate difference is therefore seen to be due to diitarences in the following governing parameters between the two locations. L _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ 110-k., Parameter Cavity liner Beam flange hole Bending moment / beam 4.845 lb. in./lb. 7.465 lb. in./lb. Thickness in direction 2.5 in. 8.0 in. of crack propagation End of life design 31.0 KSI/ E 48.6 KSI/in. fracture toughness I 6.8.7 Residual stress effects i 6.8.7.1 Analysis approach. Residual stresses will be generated I within the support beam elements by both the welding and flame cutting operations. Post weld stress relief will reduce but not eliminate the residual stresses. The residual stresses are important in that they add to the stresses produced by external loading of the support beams and thereby increase the stress intensity factors at the tip of any flaws embedded within the structure. t Residual stress data for the reactor vessel support beams have not been located. Estimates of the potential effect of residual stresses were therefore made using the following ground rules to define the residual stress magnitude: (&) The residual stress shall not encaed the yield stress of A36 at the assumed 1100*F post weld stress relief temperature. (The yield 1 stress at 1100'F was obtained from Ref. 7 as 16.3 ksi.) Creep induced relaxation of the residual stresses will not be considered ' since no data are available on the temperature-time history used in any post weld stress relief operations. (Data reported in Ref. 22 indicate that residual stress levels approaching 10 ksi can be .l achieved using conventional post weld stress relief cycles.) ( i l 4 i { i 111 (b) The maximum stress existing at any point within the beam due to the sum of (i) residust stressas and (ii) stresses produced by external loading, shall not exceed the material yield stress. This ground rule reflects the fact that the residual stressee are strain con-trolled and therefore self limiting. A reduction of the maximum residual stress will occur should local stresses exceed the yield stress. (c) The material yield stress as defined in (b) above shall be based on the initial "as built" yield stress of 40 kai and include yield stress increments due to (i) prototypic straining rates and (ii) irradiation hardening. Estimates of the distribution of residual stress are discussed in See-tion 6.8.7.4. 6.8.7.2 Material yield stress. The material yield stress is influenced by both straining rate and irradiation hardening. Prototypic straining rates produce a 9 ksi increase in yield stress of A36 (see Table 6.4). Irradiation induced increases in yield stress for the HFIR material (A212B) are defined in Ref. 22 as follows! Aoy = AT/C (13) where' ~ aoy = incremental increase in yield stress, MPa AT = NDT temperature shift (*C) C = constant = 0.6 ~ The evaluation of potential residual stress effect was limited to the beam flange hole location, where residual stresses can be introduced by the flame cutting operation. Irradiation induced NDT temperature ( shifts for the location are given in Table 6.7 as 16.7'c (30*F) and 9 ____ ____________--_________m _____ 112 i ( 41.7'c (30'F) for the current . condition and the 32 EFPY ' condition respectively. Substituting these temperature shifts and constant C into Eq. (13), and converting the yield stress increments to kei, gives the following results Current condition: Ao y= 4.0 kai 32 EFPY: Ao = 10.1 kai 7 In summary, the yield stress values to be considered- in the analysis are as follows. Static Dynamic Irradiation yield yield Plant operating period o increment stress stress (EFPY) Y (ksi) (ksi) (ksi) ZERO ZERO 40.0 49.0 7.48 (current condition) 4.0 44.0 53.0 32.0 10.1 50.1 59.1 g .- In the above table the assumption that strain rate and irradiation effects on yield stress are directly additive was made to assure a con-servative result. 6.8.7.3 Upper Bound Residual Stress Estimate. A finite element i analysis using the Microsafe 2D computer program (Ref. 6) was performed l to define the distribution of stresses in the beam flange ligament I adjacent to the grout hole. Results from this analysis are summarized in Table 6.21 and plotted in Fig. 6.53. Peak stresses from these plots were used in conjunction with the yield stress result from sect. 6.8.7.2 - to define the limiting value of residual stress in the heat affected zone adjacent to the hole surface. The first step in defining the limiting residual stress was to define the maximum residual stress which could - exist at the start of L ) 113 j ( k life, when the material yield stress-has its lowest value. This start of life limiting residual stress is the lower of (a) the material yield stress at the strass relief temperature (16.3 kei) or '(b) the static l 1 yield stress minus the stress produced by the vessel dead weight plus l pipe thrust loading. This latter value is 14.3 kai (40.0-25.7) and is therefore the governing value. In effect, the maximum residual stress throughout life can not exceed this start of life value, which is governed by local yielding as the dead-weight and pipe-thrust-induced stresses (not self limiting) are superposed on the residual stresses from the flame cutting operation (strain controlled and self limiting). At the start of life, application of any static support load in excess of the DW+T load (818 kips) would cause further yielding and a further reduction in the residual stress. The yield stress is however i increased by both high strain rates and irradiation hardening. The net result of these hardening mechanisms is to enable the material adjacent to the hole surf ace to sustain higher stresses without yielding. The residual stress element of the total stress however can not exceed the 14.3 kai value derived earlier. The residual stress may be reduced below the 14.3 kai value if the combination of non self limiting and self limiting stresses exceeds the dynamic yield stress of the irradia-tion hardened material. Analysis of the maximum possible residual stress associated with the three principal dynamic loading cases considered is summarized in Table 6.22. Note that the start of life limiting value (14.3 kai) governs through to the end of life for the DW + T + OBE and DW + T + SSE loading cases. In the case of the DW + T + SBLOCA loading however the maximum residual stress is reduced by local yielding to 4.1 ksi at 114 (, 4.78 EFPY and 10.2 kai at 32.0 EFPY. The trend of increasing upper bound residual stress with increasing reactor operating period is due to the effect of irradiation hardening of the material. Analysis results from Table 6.22 are plotted in Fig. 6.54. Two limiting residual stress envelopes are shown in the figure. The envelope defined by points A, B, C and D applies at 7.48 EFPY (the cur-rent condition) while envelope A-B-E-F-C-D applies at 32 EFPY. Crack tip stress intensity factors will be calculated for points E and F on the 32 EFPY envelope in order to define the limiting residual stress impact. 6.8.7.4 Residual stress distribution. Rapid cooling of the flame cut surface is to be anticipated due to the highly localized heat input which is characteristic of the flame cutting process and the " chill-i block" quenching effect of the bulk of the flange plate material. This combination of thermal conditions will produce high localized tensile residual strains in the region immediately adjacent to the flame cut surface and much lower compressive strains in the balance of the beam flange ligament. The objective in this section of the evaluation is to determine a distribution of residual stresses, compatible with the residual. strain distribution, which can be used in the residual-stress / flaw-tolerance impact assessment. No actual data on residual stresses generated by the flame cutting process have been located. The heat-affected zone (HAZ) produce by flame cutting of heavy steel plates has however been investigated. Results reported in Ref. 23 indicate a HAZ extending 3/16 into the material from the flame cut surface. In this analysis the tensile residual stresses define in Section 6.8.7.3 will be assumed to be 115 ( , '. , constant over the depth of the HAZ. For convenience, the HAZ depth is rounded off to 0.2 in. The resulting distribution of residual stresses is given in Fig. 6.55. l 1 It is appropriate to note that data reported in Ref. 23 also show a significant reduction in the NDTT in the region immediately adjacent to the flame cut hole. This is attributed to the fine grain structure pro-duced by the " chill-block" quenching of the ASTM A533, grade B, class 1 material used in the _ evaluation of Ref. 23. No comparable data have l been located for A36, and it is therefore not possible to include an j t \ evaluation of quenching induced microstructure refinements in the cur- l 4 l rent ar.alysis. 6.8.7.5 Residual stress-stress intensity factoes equations. Equations for the calculation of stress intensity factors in clad ( pressure vessel walls, due to weld shrinkage in the chdding, are given 1 in Ref. 24. These equations were adapted to the analysis of residual l stresses adjacent to the flame cut hole by setting the HAZ thickness equal to the thickness of a pseudo cladding layer. The equation taken from Ref. 24 are reproduced below. Ky = o* /sLa (i, qui 3 ) (14) where lo and 1 are the following influence functions 3 i, = 1.122 + 0.9513a - 0.6240a2 + 8.3306a3 (15) 1 = 0.6825 + 0.3704a - 0.0832a2 + 2.8251a3 3 (16) For deeper cracks with the craA tip beyond the HAZ, the solution is j given by f 116 e x3 = o* m a $2 i -o Y (1+R) sin 1 { .+ a q + RP 1- g I 'q -p - (17) + (i,-1) (1+R)-p(q+Rp) 1- g2+ (q+Rp) sin 1 { { Rpa l l l l In the above equations, parameters p and q define the slopes of linear fits to the residual stress distributions. In the . assumed residual stress distribution of Fig. 6.55, these slopes are zero, resulting in a considerable simplification of Eqs. (14) and (17). The remaining terms in these equations, as adapted for the HAZ residual stress analysis, are defined as follows. x a distance from the flame cut surface (in.) L = flange ligament width = 6.0 in. j a = flaw depth (in.) a = flaw depth / ligament width ratio = a/L $ = HAZ thickness / ligament width = 0.2/6 = 0.0333 o* = HAZ residual tensile stress (1b/in2) p,q = stress distribution slope constants = 0 6.8.7.6 Ligament Combined Stress Intensity Factors. Linear super-position was used to combine the residual-stress stress-intensity fac-tors defined in Section 6.8.7.5 with those generated by loads appiled to the supports. Stress intensity factors generated by external loading l were taken from the analysis of Section 6.7.3. Support loads and residual stresses used in the analysis (corresponding to points E and F on Fig. 6.54) were defined as follows: ( 3 l 117 ! -Point Support Residual on load stress Fig. 6.54 Physical significance. (kips) (1b/in.3) ) E -Highest support load for which 1427 14,300 the maximum residual stress can be sustained F Maximum support load currently 1558 10,200 specified for the plant i Ligament stresses corresponding to' points E and F in Fig. 6.54 are. ~ summarized in Table 6.23 and shown plotted in Fig. 6.56 and 6.57. Stress intensity factors corresponding to the applied loading . I stress, the residual stress and the combined stress are calculated in the spread sheets of Tables 6.24 and 6.25 for cases "E" and "F" respectively. The analysis was performed parametrically with the flaw depth "a" as an independent variable. Plots of the stress intensity ) ( factors obtained for the two cases are given in Figs. 6.58 and 6.59. ' Superposed on the plots of Fig. 6.58 and 6.59 is the 32 EFPY K Id value for the beam flange hole location (48.6 kai/in.). The intercept of the K Id line with the Kg curves for (a) support-loading acting alone and (b) support loading combined with residual stresses, provides an indication of the potential effect of the residual strasses. In case -*~ "E" (-Fig. 6.58), residual stresses reduce the critical flaw depth to a combined loading -value of just under 0.2. in. Case "F" results show a smaller critical flaw size reduction (0.21 in.), to 'a combined Icading value of just over 0.2 in. The impact of' incorporating the effect of ~ the upper bound residual stresses is therefore to reduce the critical flaw depth at the beam flange hole location to 0.2 in. 6.8.7.7 Assessment of results. Inclusion of residual stress ) effects in the flaw tolerance analysis of the beam flange hole reduces i 118 i (. the 32 EFPY critical flaw size to about 50% of the value calculated when residual stress effects were ignored. A number of critical assumptions y 'l had to be made in the analysis which produced this result, and these_ l 1 could have a significant impact on the calculated flaw size reduction. t There appears to be a potential for significant reductions in critical- ] i flaw sizes due to residual stress effects, however, and recognition of this potential should be factored into any evaluation of the. overall analysis results. j Refinement of the microstructure adjacent to flame' cut holes in heavy section material has been shown to have the potential for produc-ing beneficial shifts in the nil ductility temperature of . ASTM A533, grade B, class 1 material. It is possible that similar beneficial i microstructure changes may occur adjacent ' to flame cut holes ' in . A36- ; (- material. No data on flame cutting induced microstructure refinement in A36 material have been located however, and it was not therefore possible to include consideration of this effect into the flaw tolerance analysis. The generation of data on the possible beneficial effects of chill-quenching induced microstructure refinement is recommended as an inte- -^' gral part of any further evaluation of the flame cut grout hole. Poten-tial sources for these data are identical with those identified in Section 6.8.4. 6.9 Conclusions and Recommendations Accelerated irradiation induced embrittlement of the Trojan reactor vessel support beams is predicted to occur to the point where brittle fracture becomes a credible end of life (32 equivalent full power years) failure mode. 119 i A. Critical flaws calculated for regions of the A36 support beams, remote from structural discontinuities, have s configuration and dimen-sions that would make them relatively easy to detect using surface in-spection techniques (visual, dye penetrant, magnetic particle). The contribution of fatigue flaw growth to the dimension of the critical flaw is negligible. For the critical flaws to be present in the in-stalled support beams, they would have had to be present at the time of final inspection of the support beam assembly. It is considered un-likely therefore that a brittle fracture potential exists in those por-tions of the support beams remote from structural discontinuities, since flaws of the required size would have been readily detected and elim-insted during the fabrication process. Bases for this conclusion should be verified by a review of the inspection procedures, and any available d inspection records. Flame cut 4-in.-diam holes exist in the beam flanges, immediately above the inner support pedestals. Stress concentrations produced by these features reduce the end of life critical depth of radial flaws eminating from the hole to 0.42 in. for the SBLOCA loading condition. Residual stresses induced by the flame cutting process could further reduce this critical flaw depth. It is considered unlikely that flaws of this magnitude could have been reliably detected in any inspection of the flame cut surface of the hole. Recommendations are given for actions which could provide a generic definition of (a) the distribution of flaws to be anticipated adjacent to the flame cut hole and (b) any l beneficial effects produced as a result of chill quenching induced refinement of the material microstructure adjacent to the flame cut sur-face. Further evaluation of this feature of the support will be required when these data become available. 120 'k Extrapolation of the available fracture toughness data base for A36 was required in order to cover operating conditions for critical por-tions of the beam embedded within the concrete. Generation of fracture toughness data for these conditions is recommended. References
- 1. Bechtel Drawing No. C-368 Rev. 9 for Job No. 6478.
- 2. Minutes of a Trojan Reactor Vessel Support meeting held at the Westinghouse offices in Pittsburgh on 6-29/30-88. Attendees repre-sented NRC, PCE, Westinghouse, and ORNL.
- 3. Letter: D. W. Lockfield (PCE) to R. Cheverton (ORNL), Reactor Ves-sel Support Information Request, dated August 2, 1988.
- 4. A. M. Neville, " Properties of Concrete," John Wiley & Sons. Inc.,
1963.
- 5. Manual of Steel Construction, American Institute of Steel Construc-tion, 7th Edition, 1970.
I
- 6. " User's Manual for Microsafe 2-D," Micro-Computer Based Structural Analysis Using Finite Elements, Microstress Corp., Seattle, Washington, 1987.
- 7. " Structural Alloys Handbook," 1987 Edition, Vol. 3, Produced and Published by Batte11e's Columbus Division, 505 King Avenue, Columbus, Ohio 43201-2693.
- 8. J. M. Barsom and S. T. Rolfe, " Fracture Mechanics in Failure Analysis", Fracture Mechanics Eighteenth Symposium, ASTM STP 945, Philadelphia, 1988.
- 9. ' Personal communication from H. Ott (Westinghouse) to W. E. Pennell (ORNL) on June 30, 1988.
- 10. ASME Section III, Appendix I.
- 11. J. C. Merkle, " Fracture Safety Analysis Concepts for Nuclear Pres-sure Vessels, Considering the Effects of Irradiation," Journal of Basic Engineering, Oak Ridge National Laboratory, June 1971.
- 12. R. B. Madison and C. R. Irwin, " Dynamic Kc Testing of Structural Steet," Journal of the Structural Division, ASCE, Vol. 100, No. ST 7, July 1974.
(~ 13. Stanley T. Rolfe and J. M. Barson, " Fracture and Fatigua control in Structures Applications of Fracture Mechanics," Prentice Hall, Inc., Englewood Cliffs, New Jersey, . . i 120 l l l.
- 14. The American Society of Mechanical Engineers Boiler and Pressure Vessel Code, Section III, Subsection NF.
I
- 15. C. A. Knorovski et al., " Fracture Toughness of PWR Component Sup- l ports," NUREC/CR-3009, Sandia National Laboratories, Albuquerque, ;
NH 87185, February 1983.
- 16. W. A. Sonem, R. H. Dodds, Jr., and S. J. Rolte, "An Analytical Com-parison of Short Crack and Deep Crack CTOD Fracture Specimens of an A36 Steel," University of Kansas, presented at the 21st National Symposium on Fracture Mechanics, June 1988. -
- 17. Telephone Conference Report! W. Pennell (ORNL) with S. Anderson (Westinghouse): July 12, 1988,
' 18. Telephone Conference Reports W. Pennell (ORNL) with W. Kershall (P.C.E. ), July 12, 1988.
- 19. Section XI Appendix A of the ASME Boiler and Pressure Vessel Code.
- 20. Stress Analysis of Cracks Handbook (p. 19.4), Hinoshi Tada.
- 21. S. T. Rolfe, " Designing to Prevent Brittle Fractures in Bridges,"
University of Kansas, Lawrence, Kansas.
- 22. R. D. Cheverton et al., " Evaluation of HFIR Pressure-Vessel i Integrity Considering Radiation Embrittlement," ORNL/TM-10444, Oak I Ridge National Laboratory, April 1988. I I
- 23. D. A. Canonico, " Heat-Affected Zone Due to Flame Cutting," Heavy !
Section Steel Technology Program Semiannual Progress Report for Period Ending February 29, 1968, ORNL-4315, 1968.
- 24. A. Simonen et al., " VISA-II - A Computer Code for Predicting the Probability of Reactor Pressure Vessel Failure," Pacific Northwest l Laboratory, Operated by Battelle Memorial Institute, NUREC/CR-4486, l March 1986. I r
i l i 121~ Table 6.1. Temperature adjustment vs strain rate for A36 n' t Strain ~ rate AT Temperature adjustment (in./in./s) (*F) (*F) 10 5 15.5 +60.1 10 3 34.0 +41.6 10+1 162.7 -87.1 0.11 75.6 0.0 1 e , v y 122 i Table'6.2 ' f ( CONSTRUCTION OF THE A36 PLANE STRAIN FRACTURE TmJGNNESS CLAVES FOR REACTOR guPPORT LOADING RATES Test Test Fracture. Straining Temperature Sy/Kic Calculated Calculated T NOT Yield I Temperature ~ Strain Tsushness Rate Temp. Adje ted for Seet Fit Lower seed stress .l Rate Adjustaant Straining Mean Curve Curve Increment Rate .111WIWe .11tvin/a Strein Bete strain Rate Deg. F. In/in/eec KSt+fn*1/2 Dog..F. Dog. F. (Meen*50) Dog. F. 1.0U0E+02 7.702t+01 7.200E+01 9.500E*01 7.33M+01 6.700E+01 9.000E+01 7.010E+01 6.200t+01 8.500E+01 4.713E+01 5.700t+01 8.000E+01 6A44E+01 5.200E+01 7.500E+01 4.199E+01 4.700E+01 7.000t+01 5.974E+01 4.200E+01 6.500t+01 5.767t+01 3.700t+01 6.000E+01 5.577E+01 3.200E+01 5.500E+01 5.401E41 2.700E+01 5.000E+01 7.702t+01 5.237t+01 2.200E+01-4.50M+01 7.337t+01 5.086E+01 '1.700E+01 4.000E+01 7.010E*01 4.944t+01 1.200E+01 j 3.500E+01 6.713t+01 4.812E+01 7.000E+00 3.000t+01 6.444E+01 4.689t+01 2.000E+00 { , 2.500E41 2.000t+01 s.199t+01 5.974E+01 4.573E+01 +3.000E+00 4.465E+01 8.000E*00 1.500E+01 5.767E+01 4.362E+01 1.300E+01 1.000E+01 5.577E+01 4.2ME+01 1.800E+01 5.000E+00 5.401E+01 4.175E+01 a2.300E+01 0.000E+00 5.237E+01 4.090t+01
- 2.800E 41 1.000E+01 4.944E+01 3.932E+01 3.800E+01 1.500E+01 4.812E41 3.859t+01 4.300E+01 !
2.000t+01 4.689E+01 3.790E+01 4.800E*01 2.500t+01 4.573E+01 3.724t+01 5.300E+01 3.000E+01 4.465E+01 3.M2E41 5.400t+01 +7.500t+01 1.000E 03 4.150E+01 4.160E+01 3.340E41 1.434E+00 4.394E41 3.621t+01 6.140E+01 5.960E+01 *1.080E42 1.000t 0,5 4.580841 4.010E+01 4.790E+01 1.33M+00 4.125t+01 3.462E+01 *7.590E41 6.126E+01 3.800E+01 1.000E+01 3.900t+01 +8. 710E41 +4.910E+01 1.574E+00 4.10$E+01 3.450E+01 7.710E+01 6.14M+01 ! 9.300E41 1.000t 03 3.700E*01 4.160t+01 5.140E+01 1.667E40 4.066E+01 3.42M+01 +7.940E+01 6.167E+01 +9.500t+01 1.000E 03 3.900E+01 4.160t+01 *5.340E+01 1.588E40 4.034E+01 3.407t+01 +3.140E+01 4.192t+01 2.60M+01 1.000t+01 4.500E+01 8.710t+01 4.110E+01 1.397E+00 3.915E+01 3.335E+01 4.910E+01 6.287E+01
- 1.24M + 02 1.000t 05 4.050t+01 6.010t+01 +4.390E+01 1.541E+00 3.875E+01 3.310E41 9.190E+01 6.323t41 !
*1.280t+02 1.000E 05 4.450E+01 4.010E+01 6.790E+01 1.432E40 3.818t+01 3.276E+01 +9.590E+01 6.374t+01 ,l 1.40M+01 1.000t+01 3.70M+01 8.710E+01 7.310E41 1.741E40 3.749E+01 3.232E+01 *1.011E+02 4.443t41 +1.330t42 1.000E 03 3.300E+01 4.140t+01 9.140t+01 2.031E+00 3.531E+01 3.095E+01 1.194E+02 ,6.701E+01 *1.000E+01 1.000t+01 3.850E41 8.710E+01 *9.710E+01 1.763E+00 3.470E41 3.054E+01 *1.251E*02- 6.787E+01 *1.400t+02 1.000E 03 3.050E+01 4.160t+01 9.840E+01 2.232E+00 3.457E+01 3.048E41 -1.264E+02 6.807E+01 1 +1.600t+02 1.000E 05 3.430E+01 6.010E41 +9.990E*01 1.991E+00 . 3.M2E41 - 3.038t+01 1.279E+02 6.830t41 1 *1.650t+02 1.000E 05 3.350t+01 4.010E+01 1.049t+02 2.062E+00 3.393E+01 3.007E 01 1.329E+02 6.909E+01 *1.900t+01 1.000E+01 3.450E+01 +8.710t+01 *1.061E+02 2.000E+00 3.381E+01 2.999E*01 1.341E+02 4.928E+01 ( *1.600t+02 ' 1.000E 03 3.050E+01 6.000t+01 1.000t+01 3.300t+01 4.160E+01 1.184E+02 2.339E+00 3.271E41 2.928t*01 1.444E+02 7.133E+01 8.710E*01 *1.471E 42 2.32SE40 3.054E41 2.788E+01 *1.751F42 7.6735*01 l ) *1.900f+02 1.000E 03 2.850E+01 4.140t+01 1.484E+02 2.702E+00 3.048E+01 2.782E41 1.7ME42 7.700E+01 +2.280t+02 1.000E 05 2.960E+01 4.010t+01 +1.679t+02 2.747E40 2.931E*01 2.706E41 1.959E*02 8.131E+01 . *1.100t+02 1.000E+01 2.950E+01 8.710E+01 *1.971t+02 3.014E.00 2.788E+01 2.613E*01 2.251E42 8.897E +01 *1.1408 4 2 1.000t+01 1.050E+01 +8.710t+01 2.051t+02 2.99M40 . 2.754E+01 2.591E+01 2.331E+02 9.138t41 123 Table 6.3 (
- CURVEFIT
- CURVE FITTING No. of data rows 24 CURVE # a b c RR(adj'd)
- 1) bX -0.16314 n/a
- 2) a+bX 41.76978 0.057591 0.6592567
- 3) bX+cX^2 -0.42282 -0.00100 n/a
- 4) a+bX+cX^2 47.51319 0.144399 0.000219 0.7890883
- 5) 1/(a+bX) 0.022956 -0.00005 0.7887316
- 6) X/(aX+b) 0.037985 0.690076 0.6785572
- 7) a+b/X 25.42168 -781.488 0.6837600
- 8) a+bX+c/X 32.91161 0.030140 -464.582 0.7452817
- 9) a+b/X+ /X^2 19.24268 -1933.53 -40096.6 0.8095685 MAX RR CURVES 1 to 9 = 0.8095685 CURVE # a b c RR(adj'd)
- 10) aX^b ERR ERR 0
- 11) ab^(X) 42.42251 1.001770 0.7277982
- 12) ab^(1/X) 25.97992 1.0E-10 0.6887448
- 13) aX^(bX) ERR ERR 0
- 14) aX^(b/X) ERR ERR 0
- 15) aX/b^X ERR ERR n/a
- 16) ab^X*X^c ERR ERR ERR 0
- 17) ab^(1/X)*X^c ERR ERR ERR 0
- 18) 1/(a(X+b)^2 + c] -0.00000 382.2796 0.041014 0.8662148
- 19) a+b*lnX ERR ERR 0
( 20) 1/ (a+b *lnX) ERR ' ERR 0
- 21) ae^(bX) 42.42251 0.001768 0.7277982
- 22) ae^(b/X) 25.97992 -22.9892 0.6887448
- 23) ae^((X-b)^2]/c 24.22951 -350.864 173051.4 0.8327229
- 24) ae^([(b-lnX)^2]/c) ERR ERR ERR 0
- 25) a[(X/b)^c]e^(X/b) ERR ERR ERR 0 MAX RR CURVES 10 to 25 = 0.8662148 DESC STATS X Y TOTAL -3043.4 827.2
- TITLE
- A36 FRACTURE TOUGHNESS DATA MEAN -126.8083 34.466666 TITLE X ADJUSTED TEMPERATURE (Deg.F) n 24 24 TITLE Y FRACTURE TOUGHNESS : Kic SD 90.879939 6.3749282 MAX -33.4 45.8 l MIN -412.1 24.5 FITTED CURVE NUMBER = 18 l DATA BELOW l ROW X Y ESTIMATED Y RESIDUAL Y = 1/[a(X+b)^2 + c) 1 -33.4 41.5 44.06757 2.567574 ,
2 -47.9 45.8 41.35012 -4.44987 1 3 -49.1 39 41.14497 2.144970 4 -51.4 37 40.75941 3.759410 5 -53.4 39 40.43208 1.432083 6 -61.1 45 39.23678 -5.76321 7 -63.9 40.5 38.82614 -1.67385 8 -67.9 44.5 38.26018 -6.23981 ( 10 9 -73.1 -91.4 37 33 37.55880 35.36336 0.558807 2.363361 11 -97.1 38.5 34.75587 -3.74412 , 12 -98.4 30.5 34.62187 4.121879 13 -99.9 34.3 34.46929 0.169298 124 Table 6.3 (cont.) 14 -104.9 133.5 33.97591 0.475911 f \ i 15 -106.1 34.5 33.86087 -0.63912 16 -118.4 30.5 32.75197 2.251972 l 17 -147.1 33 30.59179 -2.40820 $ 18 -148.4 28.5 30.50613 2.006138 l 19 -167.9 29.6 29.32891 -0.27108 20 -197.1 29.5 27.89224 - -1.60775 21 -205.1 27.5 27.55685 0.056854 22 -264.9 25.5 25.68048 0.180482 .23 -283.4 25 25.28937 0.289371 , 24 -412.1 24.5 24.46173 -0.03826 L 6 i l 4 i ? 4 i 1 l l l l _ ._---___-__ _ - 125 ( Table 6.4-1 PLA E STAAIN
- PLANE STRESS TRANS! TION (Ref 22, Pete 114)
Plate Thickness . 2.500E+00 Temperature calculated calculated Dynamic SETAlc Calculated Calcutsted Adjusted for Plane Strafn Plane Strefn Yletd Plane Strees Plane Stress straining Mean Curve 4 ewer Sound Strees Mean Curve Lower Sound Rate .tiln/in/s .11tn/in/a .11tn/In/a .11tn/In/s strain aste strafn aate Strain aate strain Rate Deg. F. (Meen 50) (heen 50) 1.000E+02 1.633E+02 7.702E+01 4.838E+01 4.555t40 - 9.500E+01 1.461t+02 7.337E+01 4.871t+01 3.599t+00 9.000E+01 1.323t+02 7.010E+01 4.903E+01 2.913E+00 8.500t+01 1.210E+02 6.713E+01 4.937E+01 2.405E*00 8.000t+01 1.116E+02 6.U.4t+01 4.971E+01 2.01M+00 7.500t+01 1.036E+02 6.199E+01 5.006E+01 1.715E+00 7.000E+01 9.680t41 5.974t+01 5.041E+01 1.475E+00 1.94M+02 1.202E42 6.500E+01 9.08M41 5.76M+01 5.077E+01 1.281t+00 1.650E*02 1.047E+02 6.000t+01 8.567E+01 5.577E41 5.114E+01 1.123E+00 1.424E+02 9.272E+01 5.500E+01 8.109t+01 5.401E+01 5.152E+01 9.911E 01 1.250E+02 8.323E+01 5.000E+01 7.702E+01 5.237E+01 5.190E+01 8.009E 01 1.112t42 7.565E+01 4.500E+01 7.337E+01 5.086E41 5.229f+01 7.877E 01 1.003E+02 6.952t+01 4.000E+01 7.010E+01 4.944E+01 5.268E+01 7.081t 01 9.145t+01 6.450E+01 g 3.500E41 6.713E+01 4.812t+01 5.309E+01 6.396E 01 8.419t+01 6.035E*01 ( . 3.000E+01 2.500t+01 6.444E+01 6.199t+01 4.6895401 4.573t+01 5.350E+01 5.393E+01 5.8038 01 5.285E 01 7.817t+01 7.311E+01 5.688E+01 5.394t+01 2.000E+01 5.974E+01 4.465t+01 5.436E+01 4.831E*01 6.881t+01 5.142E+01 1.500t+01 5.767t+01 4.362t+01 5.480t41 4.430E 01 6.512t+01 4.92SE+01 1.000E+01 5.577t+01 4.266t 41 5.525t+01 4.075E 01 6.191E+01 4.736E+01 5.000t+00 5.401t+01 4.175E41 5.571t+01 3.759t*01 5.911E+01 4.570E+01 0.000E+00 5.23M41 4.090E+01 5.618t41 3.4TTE
- 01 5.663E+01 4.422E+01
*1.000E+01 4.944E+01 3.932t+01 5.715E41 2.994E*01 5.245t+01 4.171E+01 *1.500E+01 4.812t+01 3.859E*01 5.765E+01 2.787E 01 5.06M+01 4.063E+01 +2.000E+01 4.689t+01 3.790E+01 5.816E*01 2.600E*01 4.906E+01 3.965E+01 2.500t+01 4.573E+01 3.724E+01 5.869E+01 2.429E*01 4.758E+01 3.875E+01 - *3.000E+01 4.465E+01 3.662t+01 5.923E+01 2.273t 01 4.623t+01 3.792841 1 3.340t+01 4.394E+01 3.621t+01 5.960E+01 2.175E 01 4.53M+01 3.739t+01 l *4.790E+01 4.125E+01 3.462t+01 6.126E+01 1.814E 01 4.219E41 3.541E+01 *4.910E 41 4.10$E*01 3.450E+01 4.140E+01 1.780E 01 4.195E+01 3.526E41 5.140t41 l 4.066t+01 3.427E+01 6.16M+01 1.739E.01 4.152E41 3.499E+01 5.340t41 4.034E+01 3.407E+01 6.192t+01 1.698E 01 4.115E41 3.475E+01 *6.110E+01 3.915E+01 3.335t+01 6.28M+01 1.551E*05 3.981E41 3.391E+01 , +4.390t+01 3.875E+01 3.310E+01 6.323E+01 1.502E 01 3.935E+01 3.362E+M 6.790E+01 3.818t+01 3.276E+01 6.374E41 1.435E 01 3.873E+01 3.322E41 *7.310E+01 3.749E+01 3.232E+01 6.443E+01 1.354E 01 3.79M*01 3.274E41 4.140E+01 3.531E41 3.095E+01 6.7018 4 1 1.110E 01 3.541E+01 3.122E+01 9.710E+01 3.470E+01 3.056t+01 6.787E*01 1.044E.01 3.497E+01 3.000E+01 +9.440E+01 3.457E+01 3.048E+01 6.007E*01 1.032E.01 3.483t+01 3.071E*01 9.990t+01 3.442t+01 3.038t+01 6.830E+01 1.016E 01 3.46M+01 3.060E41 f 1.049t+02 3.393t+01 3.00M+01 6.909E+01 9.647E 02 3.415E41 3.026E+01 1.061t 42 3.381t+01 2.999t+01 6.928t+01 9.529E 02 3.403E+01 3.018E41 (k *1.184E+02 3.271E+01 2.928t+01 7.133E*01 8.414E.02 3.287t+01 2.943E+01 al.471t+02 3.056E+01 2.788E+01 7.673E*01 6.347E 02
- 3.065E+01 2.796E41
. 1.484t+02 3.048E+01 2.782t+01 7.700E41 4.264E 02 3.056E+01 2.790E+01 1.679t+02 2.931E+01 2.706t+01 8.1318 4 1 5.197E 02 2.936E41 2.711E+01 126 l Tcble 6.5 i i* COIP081TE Pl.ANE STRAIN / PUWE STRE!! FRACTURE TOUlletill CIRVE \( \ Temperature Calculated Cales!sted Caeposite 'Caeposite T-NOT ' edjusted for Plane Strain Plane !!ress Hean Lauer Sound Strainisq Nean Curve Nean Curve Carve Curve Rate- .111a/le/s .Illalia/s .!!!s/is/s .!!!n/In/s . Strain Aate Strale Rate Strals 8 ate Strals 8 ate leg. F. Set. F. 1.000E+02 7.200E+01 9.500E+0! 6.700E+0! 9.000E41 6.2ME+0! I.500E+0! 5.700E+01 8.000E+0! 5.2ME+0! l 7.500E+01 4.70M+0! l 7.000E41 9.680E+01 4.200E+0! j 6.500E+01 9.087t+0!. L700E+0! 4.000E+0! 8.567E+0! 9.272E+01 3.200E+0! l 5.500E+0! 8.109E41- 8.32M+0! 2.700E+0! 5.000E+0! 7.702E+01 7.565E+01 2.20M+0! 4.500E41 7.377E+0! !.CXE*02 1.000E+02 6.952E+0! 1.7ME+0! 4.000E41 7.0lM+01 9.18!!+0! 9.145E+0! 6.450E41 1.200E+0! 3.500E+0! 6.713E+0! I.4!!!+01 8.419E41- 6.035E+0! 7.M0E+00 3.000E+0! 6.444E+0! 7.l!7E41 7.Il7E+0! 5.688E+0! 2.0ME+00 - 2.500E+01 6.199E+01 7.7:!E+0! 7.282E+0! 5.371E+0! -3.0 ME+00 - l 2.000E+0! 5.974E41 6.E!!!+0! 6.824E+0! 5.097E+0! -0.0ME+00 l ( A'g 1.500E+0! l.000E+01 5.000E+00 5.767E+0! 5.577E+0! 5.40!E*01 6 !!!E41 6.427E+0! 4.857E+01.-1.300E+0! 6.!!!E+0! 6.07BE+0! 4.64E 41 -1.800E+0! '5.911!+0! 5.769E+01 4.456E 4 1 -2.3 ME 4 1 A 0.000E+00 5.237E+0! 5.fi3E+0! 5.49M 41 4.286E+01 -2.800E+0! -1.000E+0! 4.944E+0! .l 5.:4!E41 5.019E41 3.990E+0! -3.800E+0! l . -1.500E41 4.812E+0! 5.0i7E41 4.I!2E+01 3.859E+01 -4.300E+0! I -2.000E*01 4.689E+0! 4.9:5E+0! 4.689E+0! ' 3.79M+0! -4.000E+01 -2.500E+0! 4.573E+0! 4.7!iE+0! 4.57X+01 3.724E+0! -5.30M+0! -3,000E+0! 4.465E+0! 4.123E+01 4.465E41 3.662E+0! -5.800E 41 -3.340E+0! 4.394E+0! 4.537E+01'4.394E+0! L&21E+0! -6.140E41 -4.790E+0! 4.125E+0! 4.21f!+01 4.125E+0! 3. 42E 4 1 -7.590E+01 -4.910E+0! 4.105E+0! 4.19fE+0! 4.105E+0! 3.450E+0! -7.710E+0! -5.140E+0! 4.M6E+0! 4.!!2E+0! 4.M6E+0! 3.427E+0! -7.940E+01 -5.34E41 4.034E+0! 4.!!!E41 4.034E+01 3.407E+01 0.14M 41 4.110E+01 3.91M+0! L981E+0! 3.915E+0! 3.335E 41 -0.91 N 41 4.390E+0! 3.875E+01 3.935E41 3.875E+0! 3.3tM+0! 4.190E41 4.790E+0! 3.IIK41 LI73E41 3.818E41 L 776E+01 -9.590E+01 -7.310E+01 L749E+0! 3.797E+01 3.749E+01 3.232E 41 -1.0llE 42 -9.140E41 3.531E41 3.561E41 L531E+01 3.095E 41 -1.194E+02 -9.71M+0! 3.470E41 3.497E+0! 3.47M+0! 3.056E 41 -1.25tE 42 ' -9.840E+0! 3.457E41 3.41M+01 3.457E41 3.040E 4 1 -1.2HE 42 , l -9.990E+0! L442E+0! 3.467E41 3.442E+0! 3.03M +01 -1.ff9E 42 -1.049E+02 3.39M+0! 3.415E41 3.39M+01 3. M7E 41 -1.129E+02 -1. HIE +02 3.381E41 3.40H41 3.381E+0! 2.999E 4 1 -1.341E+02 . -1.!O4E+02 3.271E+0! 3.2f7t+01 3.2'i!E+0! 2.920E41 -1.M4E*02 -1.471E+02 3.056t+01 L H5E+01 L956t+01 2.708E+0! -1.75tE+02 -1.484E+02 3.04IE+0! 3.055E+01 3.H8E601 2.782E 41 -1.764E 42 -1.679E+02 2.931E41 2.936E+0! 2.931E+0! 2.7HE41 -1.959E42 -1.971E*02 2.788E+01 2.791E+0! 2.788E+01 L 61 H 41 -2.251E 42 -2.051E+02 2.754E*01 2.757E41 2.754E41 2.591E41 -2.331E42 -2.649E+02 2.568E+0! 2.568E+01 2.56M+0! 2.479E 41 -2.929E 42 . 1 l 127 l l Table 6.6. Trojan operating history Reactor i (' Equivalent full 3 Assembly average Weighted j cycle power years power ratio E.F.P.Y. I 1.13 1 3,33 2 0.78 1.26 0.9828 ' 3 0.69 1.23 0.8487 I 4 0.59 1.25 0.7375 j 5 0.37 1.26 0.4662 i 6 0.68 1.12 0.7616 7 0.51 0.75 0.3825 - B 0.69 0.75 0.5175 9 0.74 0.88 0.6512 10 0.5 0.867 0.435 11 0.8 0.85 0.68 12 Onwards 0.65 TOTAI.S 7.48 7,593 1 i 1 ~ .m ' - ( u____.1 & -W-* 128 Table 6.7. NDTT shift (ANDTT) (' Cavity inner Mode 4: Max. Beam flange Beam location surface bending at. hole center , interface location line. j Cycle 1 dpa rate (dpa/s) 2.790000E-12 2.650000E-13 1.090000E-13 Average dpa ratio to date' 1.015 1.015 1.015 EFPY to date 7.48 7.48 7.48 Total dpa to date 0.0006680878 0.0000634564 0.0000261009 l dpa rate to date- 2.831850E-12 2.689750E-13_ 1.106350E-13 l NDTT SHIFT to date (*F) 75 38 30 Remaining lifetime (EFPY) 24.52 24.52 24.52 Planned dpa ratio 0.65 0.65 0.65 ] Additional dpa 0.0014024898 0.0001332114 0.0000547926 32 EFPY dpa 0.0020705776 0.0001966678 0.0000808935 I ' Average lifetime dpa rate 2.051539E-12 1.948595E-13 8.014974E-14 ANDTT at 32 EFPY (*F) 180 96 75 { t i ( c l I ( ! l .3- , t 1 ( s I l -i 1 l,, 129 c. t g 42 . Table tutusk DES!t.N OfuAMIC STRESS INTINS!TY FACTORS l Location Equivalent Irradiation Design Design Deelen Design on vessel Futt Power ineaeed Dynamic Dynamic Dynaefc Dynamic S w oort Years NDY Fracture Fracture Fracture fracture Beam Consund shift Toughness Toughness Toughness Toughness l (a) T
- WOT Support beam temperature (Deg7) 9.000t+01 1.000t+02 1.100t+02 1.200t+02
- Cavity liner irrterf ace 0.000t+03 0.000t+00 6.200t+01 7.200t+01 8.200t 41 9.200E+01 Cavity tirer interf ace 7.480t+00 7.400t+01 1.200t+01 +2.000t+00 8.000t+00 1.800t+01 Cavity liner interface 3.200t+01 1.800t+02 1.180t+02 1.080t+02 +9.800t+01 +8.800t+01 ] Node 4 : Maa. Sending Moment 0.000t+03 0.000t+00 6.200t+01 7.200t+01 8.200t+01 9.200f+01' Node 4 : Max. Sending Moment 7.480t+00 3.800t+01 2.400E+01 3.400E+01 4.400t+01 5.400E+01 Node 4 : Max. Sending Moment l 3.200t+01 9.600t+01 3.400t+01 2.400t+01 *1.400t+01 4.000t+00 s Beam flange hole 0.000t[03 0.000t+00 6.200t+d1 7.200t+01 4.200t+01 9.200E+01 . Ieam ftange hole 7.480!+03 3.000t+01 3.200t+01 .4.200t 41 5.200t+01 6.200E 41 l Beam flenge hole 3.200t+01 7.500t+01 1.300t+01 3.000E+00 7.000E+00 1.700t+01
- 1 l
(b) Design fracture toughness Cavity liner interface 0.000t+03 0.000E+00 2.500t+02 >250 >250 >250 Cavity liner interface 7.480t+03 7.400t+01 4.900t+01 5.434E+01 6.118( 4 1 7.074E+01 Cavity liner interface 3.200t+01 1.800t+02 3.100t+01 3.180t+01 3.250t+01 3.344E+01 I Node 4 : Ken. Sending Moment 0.000t+00 0.000t+00 2.500t+02 >250 >250 #50 Mode 4 rMaar Bending Moment 7.480E40 3.800t+01 7.868t+01 9.750t+01 1.282t+02 1.8108 4 2 Node 4 : Mas. Sending Moment 3.200t+01 . 9.600t+01 4.1088 4 1 4.422t+01 4.815t+01 5.316t+01 Bene flange hole 0.000t+00 0.000E 40 2.516t+02 >250 >250 >250 team flange holo 7.440E*C3 3.000t+01 9.272t+01 1.202E+02 1.447t+02 2.5168 4 2 Ream flange hole 3.200t+01 ' 7.500t41 4.857t 41 5.371t41 6.035t+01 6.952E+01 a x( e . . . _ - . - - _ . . - . - . - _ . - _ _ _ _ - _ - - _ _ i 130 i p- '.(. g , Table M 9 Lifetime Variation in T - NDT.and Material Fracture Toughness Beam Flange Hole Location l i l 1 l l BEAM FLANGE HOLE RADIAL FLAWS , t.F.P.Y. l l7.48t+00 1.0M+01 1.5M+01 2.00t+01 2.5M+01 3.20t+01 - l Total 0.P.A. l l2.61105 3.17t 05 4.29t 05 5.41E 05 4.53t 05 8.09t 05 ! l l1.11813 1.018 13 9.071 14 4.575 14 8.288 14 8.028 14 ) I (* Rf Not shift ! l3.00(+01 3.60t+01 4.60t+01 5.5M+01 6.4M
- 01 7.50t+01 l l
Design fracture toughness l 9.27t*01 8.17t+01 7.08t+01 6.04t+01 5.43t+01 4.86t+01 j l e , + 4 s 9 131 (' % ('l j o fatte [p;49~I~ ikOlAlt : SUPPORI BEM CA!!!C;.L FLA $1*I ANALYS!! ' Stress location : !!nterface eith the cavity liner Flastype :Eliptical surface flas Flas aspect ratio all : 0.:00 Platethicknessin. I 2.500' Stresses in beae flange without flas : Unit loas : Lead case Lead case l Mas, surface l Flange sean Flar.ge bend. :
- (200 tin BM !aansaus load !aanieus BM ! stress ! stress' ! stress :
Norsal + Upset : DW + T + Olst ! 9690N..:00 1099000.000 : 5324635.000 : 8733.079: 7520.151 : 1212.922 : Faulteg a DW + T
- LOCA
- 9690 M.000 1 2643000.000 11836335.000 : 19413.022 : 16716.769 1 2696.253 :
Flas shape paraseter = 0 : Sy a 40000.000 lNoraal+ Upset (Se + $bl/Sy = "' 0.218 : 0= 1.: ! l Faulted ($a + Ebl/Sy = 0.485 : 8s 1.( Flas depth *a' in. 0.050 : 0.100 : 0.200 ! 0.300 : 0.600 1.250 : Flas depth to flange thickness ratio : 0.320 : 0.040 0.000 : 0.120 : 0.240 : 0.500 : Mestrane correction factor Me : 1.100 : 1.100 1.120 t 1.140 : 1.250 ! 1.810 : Bending correction factor Me : 1.040 : 1.020 : 0.980 : 0.940 : 0.870 : 0.890 : ( . Moraal plus upset conditions k' tal SeeMae3505f(Pile 3S;AT(a/0) [ 3125.956 : 4420.769 : 6365.583 : 7935.433 : 12305.260! 25718.091 : lb) a $b' 2 45 RTIPils3S;Rita/DI : 476.685 ! 661.171 : 898.369 : 1055.363 : 1381.365 2039.663 : XI a la) + (b) ! 3602.641 : 5081.940 7263.751 : 8990.796 1 13686.625 l 27757.754 : Faultedconditions : (al e SeeMas350Ri(File;$0t;tv a /0) : 7073.676 : 10010.759 : 14414.758 " 17969.658 : '27865.059: 58238.196 (b) e $ bens 350Ri(Pile 350RT(a/0) : 107'.446 1497.210 : 2034.341 : 2389.853 ! '3128.078 :
- 4618.784 !
K1 = (a) + th) 1 8158.122 : 11507.970 1 16449.099 ! 20359.511 1 30993.1371 6285e.981 : 91 132 i &// Table TROJAN: SUPPORT BEAM CRITICAL FLAW $1ZE ANALY!!! 87RESSINiCNS!IYSimARY INTENACEWITHTHECAV1TYLINER toad cats Inner Flan Flaw Depth (in.) - --- Cor. crete- Aspect Effective i- Ratio 0.00 0.05 0.10 0.20 0.30 0.60 1.25 4 DW+T+0lt No 0.00 0.00 3891.00 5403.00 7964.00 10180.00 :17252.00 !!951.00 DW+T+08E No' O.10 0.00 3402.00 5081.00 7265.00 8990.00 13486.00 '27757.00 W+D0fE No 0.2) 0.00 3322.00 4 80.00 6577.00 7993.00 11582.M !!9575.00 N+T+C4E No 0.30 0.00 2939.00 4140.00 5810.00 7061.00 9965.00 15101.H l DW+T+0lE No 0.50 0.00 2398.00 3379.00 4741.00 5762.00 7966.00 10933.00 DW6T+0!E Yes 0.% 0.00 N.A. N.A. N.A. N.A. N.A. N.A. H+T+0BE Yes 0.10 0.00 N.A. N.A. W.A. N.A. N.A. N.A. CW6T*0ft L: 0.20 0.00 N.A. N.A. N.A. N.A. N.A. N.A. DW+T*0i! les . 0.30 0.00 N.A. N.A. N.A. N.A. N.A. N.A. 0W6D0l! Yes 0.50 0.00 N.A. N.A. N.A. N.A. N.A. N.A. , bef+SSE to 0.00 0.00 4292.00 6047.00 8783.00 11227.00 19026.00 61769.00~ 'H+T+SSE No 0.10 0.00 3973.00 5404.00 8010.00 9915.00 !$093.00. 3)611.00 DW+T+SSE N) 0.20 0.00 3664.00 5162.00 7253.00 8815.00 12773.00 21592.00 H+i+SSE No 0.30 0.00 3231.00 4552.00 6389.00 7764.00 10957.00 16603.00 W+f+SJE 4 ? 0.50 0.00 2667.00 3757.00 5272.00 6407.00 8860.00 12517.00 CW+f*SSE Yes ~ 0.00 0.00 N.A. N.A. N.A. N.A. N.A. W.A. H+i+S!E Yet 0.10 0.00 N.A. N.A. N.A. N.A.- N.A. N.A. DWef*SSE Yes 0.20 0.00 N.A. N.A. N.A. N.A. N.L N.A. 9W6T+$$E Yet 0.30 0.00 N.A. N.A. N.A. N.A. N.A. N.A. Wei+SSE Yes 0.50 0.00 N.A. N.A. N.A. N.A. N.A. N.A. ~ 9W+T+LOCA -No - 0.H 0.00 8784.M 18376.00 17976.M tt978.M 38939.00 1862?5.00 DW+T+LOCA No 0.10 0.00 8158.00 !!507.00 16449.M 20359.00 M993.00 62856.00 DW+1+LOCA No 0.20 0.M 7473.00 10528.00 14794.M 17978.M 26052.H 44029.00 M*T+LOCA No 0.30 0.00 6594.00 9290.00 1M37.00 15043.00 't2359.M 33881.00 SW+1+LOCA No 0.50 0.00 5344.00 7558.00 INN.M 18988.00 17823.H 24454.00 BW+T+LOCA fes 0.00 0.00 N.A. N.A. N.A. N.A. N.A. N.A. . M+T+LOCA fes 0.10 0.00 N.A. N.A. N.A. N.A. N.A. N.A. SW+f+LOCA Yes 0.20 0.M N.A. , N.A. N.A. N.A. N.A. N.A. SW+T+LOCA Yes 0.30 0.M N.A. N.A. N.A. N.A. N.A. N.A. . DU+1*t0CA fes 0.50 0.00 N.A. N.A. N.A. N.A. N.A. N.A. W+i+0SE o 1091 Kips. / BW+T+$!! s 1t12 tips H+T+LCCA e 2443 Kips. (hte il htes : 11) The currestly available LOCA lead is used bere peding reselstion of licensing , actitis a the Trojan 'leat before breal'subeittal. 133 b,I W { istle TROJAN : SUPP0Af IEAM CR!ilCAL FLAW $!22 ANALYSIS Stress location Node 4: e44. IM: Iriner concrete not effective ' Flas type Eliptica! surface flas Flas aspect ratio all : 0. 00 Plate thicknese in. ! 2.!00 Stresses in beas flange without flas Unit loas ! Load case ! Lead case l Mas, surface ! Flange sean l Flange bene. :
- (200 Kip BM !aasiava load laasiava BM ! stress ! stress ! stress Mr,real + Upset s DW + i + DeE : 1527000.000 : 1099000.000 : 1390865.000 : 13762.034 : !!t50.640 1911.3h :
Faulted s OW + i + LOCA ! !$27000.000 2443000.000 :18652305.000 30592.037 : 24343.143 : 4248.894 : Flas shape paraseter = 0 : Sy a 40000.000 lNorsal+ Upset (Se + Sbl/Sy a 0.344 : 0= 1.10- ! l Faulted (Se + $bl/Sy
- 0.765 : 0s 0.97i Flas depth *a' in.
Flas depth to flange thickness ratio 0.(50 : 0.020 0.100 : 0.040 0.200 : 0.000 : , 0.300 : 0.120 0.600 : 0.240 3 1.250 : 0.500 : E Membrane correction factor Ma : 1.100 1.100 1.120 : 1.140 l 1.250 : 1.910 : Dending correction factor Mb : 1.040 l 1.020 : 0.980 0.940 : 0.670 0.990 ' Noraal plus upset conditions (al a Sa MeeiSORf(P!) d50Rita/0) : 4926.042 : 6966.475 : 10031.213 l 12505.063 : 19391.261 : 40527.890 : (b) s $bsMb 450Ri(PllOSDRita/0) ! 751.185 : 1041.907 : 1415.695 1663.096: 2176.825 : 3214.206 : K1 : (al + (b) : 5677.227 8003.3S2 : 11446.908 : 14168.159l 21569.086: 43742.096 : Faulted conditions (al s SatMa 4 SCRf(Pil 0$0Rita/01 : 11660.961 1 16491.089 : 23745.957 : 29602.074 : 45903.130 : 95937.909 : (b)
- Sb8Mn 450Ri(P!) HSCRita/01 : 1775.211 : 2466.409 : 3351.244 : 3936.192 5152.997 : 7608.692 :
Rt = (al + (b) : 13439.172 : 19957.499 l 27097.201 ! 23538.966 1 51056.126 : 103546.5'? g . 134 (o,./9 Table @ TR3JAN: SUPPORT 3EAM CRITICAL FLAW SIZE ANALYSIS STRES3 INTENSITY
SUMMARY
N0DE 4 : LOCAi!0H OF MAI! MUM IENDING MOMENT l Lead case Inner Flan Flam Depth (in.) - -- Concrete Aepect Effective i Ratio U.M 0.05 0.10 0.20 0.30 0.60 1.25 DW+T+0BE No 0.00 0.00 6133.00 8641.00 12551.00 16043.00 27186.00 88178.00 DW+T+CBE No 0.10 0.00 !677.00 8M8.00 !!446.00 14168.00 21548.00 43742.00 DW+T+0!E No 0.20 0.00 !!75.00 72i1.00 10246.00 12452.00 18044.00 30495.00 DW+f+05E No 0.30 0.M 4617.00 6505.M 9129.00 !!004.00 156!?.00 E!?25.00 DW+T+01E No 0.50 0.00 3319.00 5380.00 7550.00 9175.00 12687.00 17408.00 W+T+0!E Yes 0.00 0.00 !!27.00 7506.00 10902.00 13935.00 23615.00 765'3.00 DW+i+05E Yes 0.10 0.M 4933.00 6959.00 9947.00 12312.00 18742.00 38012.00 DW+T+0lE Yes 0.20 0.00 4501.00 6341.00 8910.00 10829.M 15692.00 24520.00 DW+T+0lt Yes 0.30 0.00 'Me 00 5644.00 7921.00 9626.00 13585.00 20586,00 DW+T+0!E Yes 0.!0 0.00 3313.00 4667.00 6:50.00 7959.00 !!007.00 15102.00 W+T+S$1 No 0.00 0.00 6932.00 9626.00 13992.00 17872.00 30286.00 98232.00 (,, f ,, DW+T+$1E ho 0.10 0.00 4348.00 8954.00 12799.00 15842.00 24116.00 48911.00 DW+f+$1E No 0.20 0.00 !!!9.00 819E.00 !!520.00 14000.00 20287.00 34286.00 N+i+S!! No 0.30 0.M !!39.00 7240.00 10160.00 12347.00 17426.00 26406.00 DW+T+35E No 0.50 0.00 4129.00 5958.00 8361.00 10160.00 14050.00 19273.00 W+T+5SE Yes 0.00 0.00 ti34.00 8361.00 12143.00 15523.00 26305.00 85318.00 DW+i+55E Yes 0.10 0.00 5490.00 7745.00 !!070.00 13702.00 20858.00 42303.00 DW+T+SSE Yes 0.20 0.00 5021.00 7073.00 9939.00 12079.00 29583.00 1750.4.00 DW*T+$5E Yes 0.30 0.00 4472.00 6301.00 8842.00 10746.00 15165.00 22?80.00 DW+f*SSE Yes 0.50 0.00 1676.00 5179.00 7268.00 8832.M 12214.00 16759.00 DW+f*LOCA No . 0.00 0.00 14616.00 20594.00 29912.M 38235.00 64792.00 210149.00 DW T+LOCA No 0.10 0.00 13439.00 13957.00 27097.00 33538.00 51054.00 103546.00 DW+1+LOCA No 0.20 0.00 !!067.M 17000.00 23888.00 29031.00 42049.00 71097.00 M*T +LOCA No 0.30 0.00 10694.M 15066.00 21143.00 25695.00 36262.00 549+9.00 DW+T+LOCA No 0.50 0.00 8653.00 12190.00 17107.00 20789.00 28747.00 39443.00 CW+f*LOCA fes 0.M 0.00 !!476.00 17572.M 25531.00 32636 M 55304.00 179373.00 H+f*LOCA Yes 0.10 0.00 11384.00 16059.M 22955.M 43252 M 28412.00 87719.00 H+T+LOCA fes 0.20 0.00 10343.00 14599.00 20514.00 24931.00 34127.00 61055.00 H+i+LOCA fes 0.30 0.00 9157.00 12900.00 18103.M 22000.M 31048.00 47048.00 DW+i+LOCA Yes 0.50 0.M 7457.00 10506.M 14743.00 17916.M 24775.H 33994.00 k +T+0$t = 1099 Kips. .
/ DW+7+$$E e 1212 kis .
L H+i+ LOA e 2443 Kips. (Nete 1) 4 tes : Ill the currently avallatte LOCA lsad is used here pending resolution of licensing actisns en
, the trejan ' leal 6afste treat' s6belttal.
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(\{. Table 6rM1r Lifetime Variation of Critical Flav Sizes at the Beam Flange Hole Location E.F.P.Y. ' l7.484*00 1.00E+01 1.50E+01 2.00E+01 2.50E+01 3.20E+01 1 Total D.P.A. l2.61E05 3.17E 05 4.29E 05 5.41E 05 6.53E 05 8.09E 05 l l1.11E13 1.01E 13 9.07E 14 8.57E 14 8.28E 14 8.02E 14 i RT WDT Shift l3.CE+01 3.60E+01 4.60E+01 5.50E+01 6.40E+01 7.50E+01 l Design fracture toughness l 9.27E+01 8.17E+01 7.08E+01 6.04E+01 5.43E+01 4.86t+01 I . l l Limiting Flow Stres : No Inner Concrete I CW + T + ost l t.49E+00 2.01E+00 1.60E+00
, DW + T
- sst l 2.81E+0C 2.07t+00 1.57E+00 1.17E+00 CW + T + SI LOCA l 2.31E+00 1.69E+00 9.70E 01 6.10E 01 4.20E 01 1
DW + T + LI:LOCA l9.3:E01 5.30E 01 3.20E 01 1.85E 01 1.35E 01 1.00E 01 l l Limiting Flaw $ltes : With Inner Concrete i DW
- T + ott l 2.38g 00 2.67.E+00 l
DV + T + SSE l 2.94E+00 2.40E+00 2.00E+00
- 1
- DM + T + SI LocA l 3.20E+00 2.51E+00 1.79E+00 1.32f+00 9.00E 01 1
DW
- T + LS:L W.A l1.73E+00 1.18E+00 4.00E 01 3.50E 01 2.50E 01 1.75E 01
v
. d 139 ,
I 'j ( ! GolY CAR TABLE N CRITICAL FLAW DEPTN$ AS A PtAICTION OF BLPPORT LOAD & (T
- IST)
BEAM FLANGE GROUT NOLE : NO INNER CONCRETE
. . . . . . . . . . . . . . . . LQ40 l NG CONo l T I ON " " " * " " " "
OBE SSE SBILOCA LBtLOCA 1 l 1099.000 1212.000 1558.000 2443.000 (Kips.) (Klps.) (Kfps.) (Kips.) , 1 T
- NOT Desfon "*" " CRITICAL FLAW DEPTM (In.)"""***"
Kid (0eg.F.) (KSIW) 1 40.000 113.986 1.720 ~ 30.000 88.926 2.690 0.750 20.000 73.199 2.990 1.830 0.360 10.000 62.842 2.700 2.240 1.150 0.210 ; 0.000 55.610 2.130 1.680 0.650 0.140 i 10.000 50.010 1.690 1.260 0.470 0.115 ! g' 20.000 45.700 1.300 0.890 0.320 0.085 30.000 42.270 1.100 0.720 0.275 0.078 40.000 39.376 0.720 0.480 0.205 0.065 50.000 37.636 0.550 0.380 0.160 0.060 'k
*60.000 36.400 0.480 0.330 0.135 0.055 l l
...............P + 100*(T NOT).................
40.000 6443.000 30.000 4558.000 5443.000 . 20.000 3212.000 3558.000 4443.000 10.000 2099.000 2212.000 2558.000 3443.000 I 0.000 1099.000 3212.000 1558.000 2443.000 l 10.000 99.000 212.000 558.000 1443.000 20.000 901.000 788.000 442.000 443.000 30.000 1901.000 1788.000 1442.000 557.000 40.000 2901.000 2758.000 2442.000 1557.000 5C.000 3901.000 3788.000 3442.000 2557.000 60.000 4901.000 4788.000 4442.000 +3557.000
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(- Table tuGESF 0408R 07 MAONinct FAfl0UE CRACK OROWTN ANALYSit twent Sefusting 0.8.t. 8.8.t. Smett treek Totals Totals L.O.C.A. e+b+c e+b+d (a) (b) (c) (d) N eber of cycles 4.0000t+01 4.0000t+02 2.0000t+01 5.0000t+00 Maa Lead a pasa 8.1800t+02 1.0990t+03 1.2120t43 1.5540t+03 Min Load a pain 3.9500t42 5.3700t+02 4.2400t42 7.8000t+01 Lead Range a pasa Poln 4.2300t+02 5.6200842 7.8000E+02 1.4400t+03 Load Ratio e e /Pasa = a 5.1711t 01 5.1137t 01 6.5017t 01 9.4994t 01 CAvifY LINER INTERFACE Etc (Est) 3.100W 4 3.1000t+01 3.1000E+01 3.1000t+01
*=R* Ele 1.6031t+01 1.5833t+01 2.0155t+01 2.9448t+01
) d4/dN e 1.158t 13* cst'5.837 1.25021 06 1.1714t 06 4.7578t 06 4.3510t 05
(\ N
- d4/dW 5.0010t 05 4.6855E 04 9.515M 05 2.1755E 04 6.1371t 04 7.3411t 04 N001 4 Kit (K$t) 4.1100t41 4.1100t+01 4.1100E+01 4.1100t+01 set
- R
- Etc 2.1253t+01 2.1017t+01 2.6722t+01 3.9042t+01 dA/m a 1.154E 13* cat'5.837 6.4450t
- 06 6.0759t M 2.4679t 05 2.25698 M N
- dA/*
~D 2.5%M M 2.4304t 03 4.9357t M 1.12848 03 3.18338 03 3.5182t 03 BEAM FLANGE NQLt Ele (K51) 4.8600t+01 4.8600t+01 4.8600E+01 4.8600t+01 er e a
- tic 2.5132t41 2.4853t41 3.1598t+01 4.6167t+01 m/m a 1.158t 13+ cst'5.837 1.7251t 05 1.61638 05 6.5644t 05 6.8035E M M * */m 6.90ME M 6.44508 03 1.3130t 03 3.0014t 03 8.4481t 03 1.0157t 02
_.__r._.___. _ _ _ .n 09. S'L.- *L~
14 1 I. &, Table 6iug5F i j 87Atl$ INTEN$177 FACTOR EOUATION COMPAflI!LITY CNBCE 1 2 3 4 5 6 7 8 9 10 I l Seem Actust 8ffective 8f fective "e*/*t" l Stress Intensity Factor l Flange Necualised] Design j Location Flow Flow Bean j""*"""""*""""l. 14een Elf l Stress l Depth Depth Thickness l DW T+00E DW T+S8 LOCAl Strees lltensityl l l l l factor [ ; e "a" *** K$l*(In.)*1/2 l l KSI Kl/K31 l l 1 I I Flange Note I 0.00 2.00 8.00 0.25l 0.00 Flange Hele 0.00l 16.43 0.00l 31.00l 0.05 2.05 8.00 flange Note 0.26l 15.29 21.68l 16.43 1.32l 31.00 [ i 0.10 2.10 8.00 - Ftange NoIe 0.26l 21.72 30.80l 16.43 1.87l 31.00l 0.15 2.15 8.00 Flange Note 0.27l 25.29 35.86l 16.43 2.18l 31.001 0.20 2.20 8.00 0.28l 27.62 71ange Note 0.30 39.30l 16.43 2.391 31.00l 2.30 8.00 0.29l 31.45 Flange Note 44.60l 16.43 2.71l 31.00i 0.40 2.40 8.00 Flange Note 0.30l 33.88 68.04l 16.43 2.92l 31.00l 0.60 2.60 8.00 7ienge Note 0.33l 38.08 54.00l 16.43 3.291 31.00l 1.00 3.00 8.00 F16nge Note 0.38l 62.68 60.51l 16.43 3.68l 31.00l 2.00 4.00 8.00 Flange note 0.50l 53.39 75.70l 16.43 4.61l 31.00l 3.00 5.00 8.00 Cavity Liner 0.63l 66.59 94.40l 16.43 5.75l 31.00l 0.00 0.00 2.50 Cavity Liner 0.00l 0.00 0.00l 10.M 0.00l 48.60l 0.05 0.05 2.50 Cavity Liner 0.02l 3.89 5.54l 10.66- 0.52l 48.601 0.10 0.10 2.50 Cavity Liner 0.04l 5.48 7.81l 10.64 0.73l 48.60l i 0.20 0.20 2.50 Cavity Liner 0.08l 7.96 11.34l 10.66 1.06l 48.60[ 0.30 0.30 2.50 Cavity Liner 0.12l 10.18 14.51l 10.64 1.36 l 48.60l 0.60 0.60 2.50 Cavity L iner 0.24l 17.25 24.58l 10.66 2.31l 48.60l 1.25 1.25 2.50 0.50l 55.96 79.73l 10.66 7.48l 48.60l
\
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- v 142 io eus 6,2. /
8 TAM FLANGC GROUT Nott LIGAMENT STRES$ts l' '
' Distance from hole surf ace 0.00 0.25 0.63 1.38 2.38 6.00 Load Coeinstion Man. Load ......... **........ Ligament Stresses.*** ..................
(Kips.) OW+fM+$tt0CA DW+1H+$$t 1558.00 48903.00 39906.00 31350.00 23063.00 20049.00 16199.00 DW*1M+00t 1212.00 38042.64 31043.69 24387.80 17941.18 15596.53 12601.53 ; DV+1H 1099.00 34495.76 28149.35 22114.02 16268.44 14142.39 11426.64 818.00 25675.64 20951.93. 16459.76 12108.82 10526.37 85?4.99 [o.E b Table 6serC~ BEAM FLANCE GRCUT NOLE LIGAMENT LIMITING RESIOUAL STRESSES Load Correination DV+1H DW+TN+0st DW+TN*S$t O bfM+$8:LO A Man. Surface Stress = $s (lb/in'2) 25675.64 34495.76 38042.64 48903.00 static Yield Stress o Sys (Ib/in*2) 40000.00 W.A. N.A. N.A. Static Load Limited $r = $r1 ssys
- Ss 14324.36 N.A. N.A. W.A.
Stress tetief Limited $resys(1100 Deg F) 16300.00 N.A. N.A. N.A. (A .g Limitin9 residust stress At tsZero : $rs 14324.36 14324.36 14324.36 14324.36 Dynamic Yield At 7.48tFPY s Sydt 53000.00 $3000.00 53000.00 53000.00 l Oynamic Load Limited $r $r2 e $dy1 $s 27324.36 18504.24 14957.36 4097.00 Dynamic field At 32tFPY = syd2 59100.00 59100.00 59100.00 59100.00 Dynamic Load Limited $r a st3 = $dy2 ss 33424.36 24604.24 21057.34 10197.00 LIMlflNG RE$10UAL $14t$$ At 7.48 (FPY 14324.36 14324.36 14324.36 4097.00 LIMlflNC als10WAL $$ats$ Af 32 LFPY 14324.34 14324.36 14324.36 10197.00 (o .I. 3 table 4;1 3 COM8tNic 8 TAM FLANGC L]GAMENT Sitt$$t$ Distance from hole surface 50.00 0.20 0.20 0.25 0.63 1.38 2.38 6. OvetnesstoCA (1558 Kips) 10200 PSI sesicNal Stress 48903.00 41705.40 41705.40 39906.00 31330.00 23063.00 20049.00 16199 10200.00 10200.00 351.72 351.72 1151.72 351.72 351.72 351. Tot AL CA$t *F"
$9103.00 $1905.40 41353.68 39554.28 30998.28 22711.28 19697.28 15847
( 1427 Eips Applied $umort Load 14300 PSI Residuet stress 101At CA$t "t" 44791.13 38198.72 38198.72 36550.62 28714.02 21123.81 18363.24 14300.00 14300.00 493.10 493.10 493.10 493.10 493.10 14834.' 493.
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3 304 i-'
- 7. BRITTLE FRACTURE EVALUATION OF TURKEY POINT UNIT 3 REACTOR PRESSURE VESSEL SUPPORTS 7.1 Introduction and Summary As indicated in Figs. 7.1-7.6, the reactor vessel for the Turkey Point Unit 3 plant is supported on steel cantilever-beam arrays that are :
positioned under each of the coolant nozzles and that extend from this ) l position into the concrete biological shield. Thesupportload$ teach- l nozzle is transmitted from the nozzle to the cantilever-beam array through'a set of rollers, lateral restraints and a steel girder that is bolted to and joins together the cantilever beams in an array. Three components of this assembly were selected for evaluation with regard to propagation of flaws: the cantilever, beam, the bolts that join the ( girder and beams, and the lateral restraints. The cantilever beam is - considered to be the more critical of the three, and thus the emphasis 1 in this report is on the beam. l I The beam is fabricated from A588, a low alloy structural steel with ! I a minimum yield strength of 50,000 psi. The radiation sensitivity of I A588 was assumed to be essentially the same ' as for the HFIR pressure
. vessel materials, and the radiation-damage trend curves deduced from the
~
HFIR vessel surveillance data and discussed in Sect. 3 were used for this evaluation. Best-estimate present-day and 32-EFPY values of ANDTT corresponding to the location of the minimum critical flaw size are 30*F and 55'F, respectively. Essentially no fracture toughness data, other than CVN data, are available for A588, and thus the ASME Sect. III and XI _ KIc and K IR ( curves were used. Dynamic response data were not available for the l Turkey Point vessel-support system, and thus Egg was used as the best estimate of the appropriate K Id' _ _ _ _ _ _ _ _ _ - _ _ _ : - - " - - ~ ~ - - ~ ~ ~ ' -
205 ( l Loads applied to the supports were supplied by the utility and l l included those for the large-break loss-of-coolant accident (abbreviated j i as LOCA in this sect' ion of the report), operating basis earthquake (OBE), safe shutdown earthquake (SSE), and deadweight or dead load (DL)./ It is ORNL's understanding that Turkey Point still must consider ; I the LOCA credible, and thus it was treated as such in this study.* (The j utility did not provide loads for a small-break LOCA, and as mentioned above, no dynamic loading data were furnished.) I Significant uncertainties exist in many of the parameters included 1 in the calculation of the critical flaw size. Thus, the study was con- l ducted in a parametric manner so that ranges of critical flaw sizes were obtained. Results for the beam, (Figs. 7.7-7.12) indicate that for the most severe credible loading (DL + LOCA) and 32 EFPY the best-estimate minimum critical flaw size (depth) is 0.3 in. The flaw size is insensi- l tive to reactor operating time after -10 EFPY, and at plant startup (O EFPY) the size is 0.6 in. The flaw is a corner crack in the upper flange anywhere between the inner surface of the biological shield and a ' point -4 in. into the shield. Considering the uncertainties in the l t f racture toughness, initial NDTT, and the operating temperature of the l . . i beam, the'ilo (one standard deviation) values of the critical flaw size i at 32 EFPY are -0.2 in. and 0.6 in. i For the DL + SSE condition, the critical flaw size is substantially ! I larger (1.1 in. compared to 0.3 in. for best-estimate values and j 32 EFPY). l C I 1
*Many PWR plants have been exempt from consideration of LBLOCA l loadst but at the time of this study, Turkey Point was not.
)
l
206 l [ 7.2 Description of supports The Turkey Point reactor vessel" is supported on cantilever-beam arrays positioned under each of the coolant nozzles (Figs. 7.1--7.6). A I nested group of three rollers transmits vertical load from the pad under . each nozzle to a wide-flange girder and also accournodates radial thermal expansion'of the vessel. Each girder rests on three wide-flange beams 1 that project inward from the biological shield. Welded to _each girder are lateral restraints that project upward on either side of the nozzle l l pads to transfer horizontal loads to the cantilever beams. Each cantilever beam is welded to two supporting columns that are embedded within the biological shield. The column base plates were anchored to the biological shield at an intermediate stage of concrete pouring.
- i
( . Each column has an array of " Nelson" studs to provide for load transference from the structural steel members to the reinforced concrete biological shield. Two of the beam arrays are synunstrically oriented with respect to the vessel nozzle centerlines; whereas four of the beam arrays are asymmetrically located. The two types of support arrays were designed to have equivalent compliances to vessel loading.
,- The dimensions L= 13 in. and L = 17 in. (Figs. 7.3 and 7.4),
define the locations of flaw assessment for the cantilever beams. They are the distances from the point of load application to the interfacing of the beam with, respectively, the biological shield and the nearest , support column. The location defined by L = 13 in. is the point of maximum radiation damage and significant stress in the portion of the beam within the cavity; whereas the location defined by L = 17 in. is ( the point of maximum stress in the portion of the beam within the biological shleid. r
- 1 i
207 , l l k 7.3 Materials The principal vessel support components of interest for Turkey Point and the corresponding materials aret _ Component Material j i Cantilever beams Bethlehem Steel Co. " proprietary steel Hayari R-50 - conforming to
. ASTM A588 Cirder-to-cantilever ASTN A354 CrBC 1 beam bolts l
Lateral restraints ASTM A302 CrB i l 7.4 Material Fracture-Toughness Properties ' No mechanical or fracture-toughness properties are available for ( the heats used in fabrication of the supports, and no data are available in the open literature for ASTM A354 CrBC. However, a few dynamic data (unknown loading rate) are available for A588 (Ref. 1) and are compared ' with the ASME K ge and XIR curves 2 in Fig. 7.13 (the data as published are not valid in accordance with ASTM E399 (Ref. 3), but those shown in Fig. 7.13 have been converted to " valid" values using a method proposed
. by Herkle4). Except for one data point, the KIR curve, appears to represent the dynamic data adequately. In the absence of more definitive data, ORNL chose to evaluate the cantilever beams premised on the ASME KIR curve. Since ASTM A302 CrB fracture toughness data are an Integral basis for the Kyg curve, evaluation of the lateral restraints logically should use this data base also. The evaluation of girder-to-cantilever beam bolts has also employed the Egg curve.
( Although the KIR curve was selected for evaluation of the Turkey Point vessel supports, it was also of interest to perform some isaim aii
208 t' calculations using the ASME K Ic curve which represents the lower bound of static crack initiation data for a select group of pressure vessel low-alloy carbon steels. As previously noted, no specific data on the initial nil-ductility l transition temperature (NDTTo ) are available for the components j assessed. However, Ref. 1 includes CVN data fu A588 and indicates that }, l for this material NDTTo = T 3 o, the temperature at which CVN = 30 ft lbs. Based on this approach and the included CVN data'(Fig. C-5 f 1 of Ref. 1), NDTTo a 35'F. An inquiry was made to Bethlehem Steel on fracture toughness and j 6 nil-ductility transition temperature of Mayari 150 steel. In response,5 l it was recoassended that data on ASTM A572, provided in an accompanying ! l y paper,6 would represent a " worst case" condition for large rolled sec- ] l ! I tion steels. Assuming NDTT o = T 3 a to be appropriate, Fig. 7 of Ref. 6 l l indicates a range in NDTT, of 10 to 95'F. In earlier studies?,a con-
)\
ducted at Lehigh University, in a few instances, the use of T o 3 would l 1 infer NDTTos exceeding 150*F. i Reference 9 includes a statistical analysis of several heats of A572 steel that indicate mean values for NDTT o of 25'F for the as-hot-rolled ' condition and -50*F for the normalized condition. A review of a l Table B.8 of Ref. 9 indicates that data for a thickness range of 0.625
]
to 2.5 in. were used for the as-hot-rolled condition and that data for a j range of 0.5 to 1.5 in. were used for the normalised condition. Fur- { thermore, when restricted to a range of 1 to 1.5 in., the NDTT, for the as-hot-rolled condition exhibited a range from 20 to 50'F. The pooled 6 ( NDTT, data on A572 from Ref. 9 (Fig. B13), exhibited a range from -60 to j 100*F. The low-temperature values in this range appear to represent
209 l l (' data taken from thin, normalised material, while the higher values cor- j respond to the as-hot-rolled thicker sections. The data of Refs. 10 and 11 indicate that the NDTT, of A588 material and associated HAZ and weld may be raised by as much as 40*F as
~
a result of a postweld heat treatment. It is not known whether the cantilever beam was normalized and/or subjected to a postweld heat treatment. In the absence of this informa-l tion but considering the data that are available (Refs. 5-11) and the > thickness of the beam flanges (2.5 in.), ORNL selected NDTT, = 65 t 65'F. No definitive data for the A354 CrBC bolting, material could be found, and thus NDTT, = 65 1 65'F was selected for the bolts, also. Munsetz estimates that a typical range for NDTT, for 2-1/2 to 4-1/2-in.-thick steel material conforming to ASTM A302 CrB would be 10 to 100*F. Thus, NDTT, = 55 1 45'F was selected for the evaluation of the lateral restraints. As discussed in Sect. 3, the radiation damage trend curve (ANDTT vs dpa, dpa/s) was assumed to be the same for the support materials as for the HFIR vessel materials, and it was necessary to extrapolate the HFIR' data relative to dpa and dpa rate in order to apply it to the supports. Two meth'ods of extrapolation (Method A and Method B, as discussed in Sect. 3), have been used in this evaluation. Figure 7.14' applies to Method A and Fig. 7.15 applies to Method B. As discussed in Sect. 3, Method B is perhaps the preferable procedure for obtaining ANDTT. Nevertheless, because of.the considerable uncertainty as to the best method for extrapolation of HFIR data, it was considered prudent to incorporate both methods in the evaluation of the Turkey Point reactor { vessel supports. 4 ___m.__._._ _ _ _ _ _ ___ _ _ _ a
210 ( As indicated in Fig. 7.13,'the ASME Kg , and KIR values are plotted as a function of T - RTNDT, where RTVDT is the reference nil-ductility . ) temperature. It was assumed for this study that RTNDT = NDTT = NDTT, j
+ ANDTT. Each of the terms in the expression (T - NDTT, - ANDTT) has l l
considerable uncertainty. To account for these uncertainties in a simplistic manner for this initial study, it _ was assumed that the expression has a Caussian distribution, and that the algebraic sums of the uncertainty limits for T and NDTT, correspond to the t3a position of , the Caussian distribution for T.- NDTT (no uncertainty is assigned { 1 to ANDTT). The fracture toughness (Ky , and Egg) were evaluated at the l l following values of the presumed T - NDTT Caussian distribution: ~3o, j
-lo, mean, +10, and +3o. The lo values have been determined as follows:
0. (T - NDTT)yg = i , [(T - NDTT) ,,- (T - NDTT),g)/2, where max and min refer to the uncertainty limits. Thus, the T - NDTT values evaluated as functions of the Caussian distribution are as follows: Caussian Point Evaluation Expression
-30 (T - NDU),g,
-lo ' (T-NDTT),,,-0.6845{T-NDTT}(min-mean) mean 0.5{(T-NDTT),,,+(T-NDU) min}
-J
+1a (T - NDTT)m an + 0.6845 (T - NDTT)( ,, ,,,)
30 (T - NDTT),,, 7.5 Thermal and Radiation Environment { The range of operating temperatures, T, for the Turkey Point reactor vessel support components is not known at this time. The l l
- c. . .. . . .- ...... .. . -- . ---
. . , , l
. , 1 211 1
l i i j maximum operating temperature for the containment is understood to be 1 120*F. Figure V of Ref. 13 provides the maximum anticipated temper-atures that will exist across the biological shield. At the L = 13-in. position (interface with the biological shield), the inferred maximum temperature for the cantilever beam is 130*F. At the L = 17-in. posi-tion, the point of highest vertical and horizontal bending stress, the inferred- maximum temperature is 138'F. Information received from-Westinghouse Electric Corporation on the temperature of the shoe of the Trojan Plant reactor vessel support (see Sect. 6) infers that the lat-eral restraints of the Turkey Point reactor vessel supports will operate at a temperature of about 340*F at the point.of maximum bending. stress. ORNL has assumed a 30*F operating temperature range for all compo-
- nents of the reactor vessel support other than the lateral restraint. A
- k. -
10*F operatind range has been assumed for this latter component. Conse-quently, the operating temperature ranges ares l Component Temperature Range. *F l Cantilever beam L = 13 in. position 100 to 130 L = 17 in. position 108 to 138 Cirder-to-cantilever 90 to 120
,- . b,eam bolt Lateral restraint 330' to 340 Extensive sophisticated calculations of neutron flux and dpa rate were performed by Florida Power and Light Co.,14'15 for - the Turkey Point vessel supports. Figure 7.16 shows the location of ten points on the supports for which neutron fluz and dpa rate were calculated.
Points 4, 8, 9, and 10 correspond to respectively, (1) the girder-to-cantilever bolts (2) Interal restraints; (3) cantilever beams at loading are length (L) equal to 13 in. (interface with biological shield
l l 212 l surface)1 and.(4) cantilever beams at loading arm length (L) equal to 17 in. . (intersection with support column and location of maximum bending stress). Table 7.1 sumanarises the dpa-rate data for the ten points as a 1 function of time, i.e., fuel cycles (the dpa rate for point 10 was esti-i mated by multiplying the dpa rates at point 9 by the ratio of fluxes at i point 10 and point 9 given in Fig. 7.17). The data in Table. 7.1 were i adjusted by time weighting to provide an estimate of the mean dpa rate I l and dpa from initial plant startup until the end of the eighth fuel l i cycle (7.04 EFPY, Oct. 1, 1983), ninth fuel cycle (11.79 EFPY, Aug. 3, ! 1988), and 32 EFPY for points 4, 8, 9, and 10. These data are provided in Table 7.2. ( 7.6 Estinstes of ANDTT 1 The radiation damage trend curves (ANDTT vs dpa and dpa rate) used in this study for the support materials are discussed in Sects. 3 and 7.4, and the dpa rate and dpa data for the supports are given in Tables , 7.1 and 7.2. Values of ANDTT derived from this information are provided in Table 7.3 as a function of time and trend cure for each support com-ponent considered. 7.7 Support Loadinas and Stress Analysis Critical-flaw-size assessment of components of the Turkey Point reactor vessel supports has been premised on utilization of the design analysis performed by Bechtel Corporation, the Architect Engineer (AE) for this plant. Only minor corrections or extensions in analysis to the (_ Bechtel stress analyses have been required for the assessments as here-after explained. __ u . . _x . __
$. e 213 i l
l f l As noted in Sect. 4, the AE analysis of PWR supports in general do not incorporate an assessment of the effect of differential friction in support provisions for radial expansion. S 2ch an effect was not exa-mined by the Turkey Point AE, and ORNL has not'made etn assessment of the potential consequences. Furthermore, residual and chermal stresses are important with regard to evaluating critical flaw sizes, but insufficient information was available for their inclusion. -As a result of these omissions, it may be that actual critical flaw sizes are less
, than those calculated in this study.
The Bechtel design document for the Turkey Point containment struc-tures13 and five sets of Bechtel calculations 1s-2o in support of the design thereof were supplied by Florida Power and Light Co. Design cal-culations included in Refs. 16 and 17 appear to involve a conceptual i stage of design, whereas Refs. 18, 19, and 20 appear to represent the final design. In the conceptual stage of design, the AE designer assumed that the cantilever beams were " built-in" at the point of inter-face with the biological shield (indentified as the L = 13-in. position on Figs. 7.3 and 7.4). In the final stage of design, the AE designer l l assumed that the cantilever beams were " built-in" at the inner face of the in' board column (identified in Figs. 7.3 and 7.4 with the dimension L
= 17 in.). The AE designer assumed that the concrete between the biolo-sical shield inner surface and the inboard column provided no effective support to the cantilever beams. These assumptions reflect standard '
design practice and are consistent with the analysis performed in Sect. 6. { The evaluation of the cantilever beams and lateral restraints were performed using the loadings and stresses from Refs. 18-20. 'Ibe loads
. . _ . . m., . . .mm . .
214
'~
and stresses for the girder-to-cantilever-beam bolts are covered only in Refs. 16 and 17, but were corrected "to reflect the subsequent design changes in Refs. 18-20 that are a consequence of a required increase in girder size. In addition to this correction, the stresses that result from the load combination of dead load (DL), and no-loss-of-function earthquake (E') (safe shutdown earthquake (SSE), in the vernacular of l later plant designs] were also evaluated. This latter loading condition was not germane to the - Bechtel ' design anelysis because it did not control the size of the components. Therefore, it was not included in Refs. 16-20. ' It should be noted that in contrast to the Trojan Plant discussed in Sect. 6, pipe loadings due to thermal constraint were considered to be insignificant for Turkey Point. In addition, dynamic methods of ( analysis were evidently unavailable at the time of' the Turkey Point I design so that static analysis of enveloped loads were employed to , i assure that the reactor vessel support design could adequately accommo-
)
date loadings involving LOCAs. In further contrast to Trojan, the i i LBLOCA was included by the utility as a credible loading condition for } Turkey Point. Many PWR plants have been exempt from consideration of j
~ .
\
LBLOCA loads, but at the time of this study Turkey Point was not. { I The LBLOCA " break" was assumed to be a multi-oriented, axially- ! split pipe failure, rather than a guillotine type. A significant , attendant consequence of this assumption is the important role of hori- I sontal loadings in the design analysis and the calculation of critical I flaw size for the cantilever beams. Table 7.4 provides the load combinations used by the Bechtel Corp. ( designer for the final design of the Turkey Point vessel supports, and
{ 215 5. Table _7.5 p evides the corresponding stresses for the cantilever beams, girder-to-cantilever beam b.olts, and lateral restraints. As previously noted, in order to permit evaluation of critical flaw size as a function l of loading condition 3 (Table 7.5) ORNL calculated the stresses shown in parentheses in Table 7.5, using the stress analysis method of the AE given in Refs. 16-20. 7.8 St 1;s Intensity Factor (K y) Evaluation Because of the large' influence that the horizontal load had on the design of the vessel-support cantilever beams, a variety of flaw shapes was assumed,.some more responsive to vertical loading, some to horizon-tal, and one to both directions of loading. Semielliptical surface flaws with varying ratios of major to min.r diameter (from 1.0 (semi-circular) to = (single-edge-notch flav)) were assumed to exist on the top surface of the cantilever-beam flangel these flaws were particularly responsive to vertical loading. Single-edge-notch (SEN) flaws were assumed to exist on the edges of the top flange (these flaws were particularly responsive to horizontal loading), and quarter-circular corner flaws were assumed to exist in the upper corners of the top
~
flange' (these flaws were responsive to both vertical and horizontal loading). The girder-to-cantilever beam bolts were assumed to have complete circumferential flaws oriented perpendicular to the bolt centerline and therefore were responsive to bolt tensile loading. The lateral res-traints were assumed to have semielliptical surface flaws oriented per- {' pendicular to and intersecting the vertical surfaces and therefore were responsive to horizontal loading. Aspect ratios of 1 1 to 6 1 were con-sidered for these latter flaws.
216 ! 7
/
k The stress intensity factors for semielliptical flaws in the canti-lever beam and lateral restraint were' evaluated by means of the method suggested by Merkle.21,22' (Previous comparisons # of this solution with finite element analyses indicate that the stress intensity factors for some elliptical shapes other than semicircular are Overestimated. Nevertheless, its use was considered to be expedient for this study since the analytical format facilitates the parametric it'estigation.) The Herkle solution was developed for a combined tutsion and bend-ing-stress field and permits evaluation of Ky as a funct ion of position on the crack front. In this study the maximum value of Ky on the crack front was assumed to be the critical value. Figures 7.18-7.21 show the stress intensity factor as a function of flaw depth, aspect ratio, and loading condition. The stress intensity factors for the SEN flaws located on the top j I surface and the edge of the top flange of the cantilever beam were eval- , I usted by the expression given for SEN specimens under pure bending given _i in Ref. 24 (p. 2.13). The expression is ; Kg =o /ia F f,
\ /.
where 'g K ' la the stress intensity factor, o is the maximum bending stress, a is the flaw depth and b is the -specimen depth. ! Figures 7.22-7.25 show the stress intensity factor as a function of flaw depth for the SEN flaws located on the top surface of the top flange. Figures 7.26-7.29 show the stress intensity factor as a function of flaw depth for SEN flaws located on the edge of the top flange of the canti-
~
( lever beam.
217 ( Two methods of solution (Liu2s and Newman and Raju2s) were used to l l evaluate the stress intensity factor for the quarter-circular corner l crack subtending the edge and top surface of the top flange of the can-tilever beam. Both solutions are premised on a uniform stress field, and the stress used in the analyses was the sum of the maximum bending stresses resulting from the vertical and horizontal loads. In both cases, the maximum value of Ki on the crack front was used for calculat-ing the critical flaw size. Figures 7.30-7.35 show the maximum stress intensity factor for the 1 corner crack of the cantilever beam evaluated by the Liu method, while Figs. 7.36-7.41 show the maximum stress intensity factor for the corner crack of the cantilever beam evaluated by the Newman-Raju method. Two methods were used for estimating the stress intensity factor for the assumed circumferential flaw of the girder-to-cantilever beam bolts. One method is that given in Ref. 24 (p. 27.1) and the other is one proposed by Brown and Srawley.27 The calculated Ky values are given in Figs. 7.42 and 7.43; as indicated, the agreement between the two methods is good. A similar calculation of - stress intensity factor for an assumed pratightening load of the bolts is shown in Fig. 7.44. '*~ A's previously noted, the method developed by . Merkle was used to evaluate the stress intensity factor for assumed semielliptical flaws in the lateral restraint; results are shown in Fiss. 7.45 and 7.46. 7.9 Discussion of Results and Conclusions Using the definition of the fracture toughness-temperature param-eter (T - NDTT) developed in Sect. 7.4 as a Caussian distribution with ( the values for ANDTT taken from Table 7.3, Table 7.6 provides a summary of T - NDTT as a function of method of extrapolation of ANDTT, Caussian
l 218 l c i Point, and service life at 0, 7.04, 11.79, and 32 EFPY for each of the locations being evaluated for each of"the structural components. Table 7.7 was developed from Table 7.6 by reading the value of toughness, Kyc l or KIR, from Fig. 7.13 for each value of T - NDTT. 1 I Equating K ye or K IR, given in Table 7.7, to K y in Figs. 7.18 ! through 7.46 permits an estimate of the critical flaw size.for the var- 1 1 i Following this procedure, ious parameters considered in Table 7.7. Table 7.8 provides a susanary of the estimated critical flaw sizes for t the cantilever beam for semicircular, 6:1 semielliptical surface, SEN top surface, SEN edge, and corner quarter-circular flaws, for both methods of extrapolation of ANDTT, for evaluation per ASME KIc and K IR' i for each loading condition of interest, as a function of the Caussian point stress intensity factor temperature parameter and as a function of j s l service life. l l The significant points to be noted about Table 7.8 are the follow-ing:
- 1. The calculated mean value of flaw size represents the best estimate.
- 2. In every loading case involving both horizontal and vertical load-
, ing, the critical corner crack is the smallest of the various types an'd consequently the type of gecatast concern.
- 3. The use of Method A versus the use of Method B for extrapolation of ANDTT data does not lead to significantly different results for this evaluation.
- 4. Very little difference in critical flaw size exists between evalua-tions at the L = 17 in. versus L = 13 in. locations.
{ Table 7.9, a sumanary of the estimated critical flaw sizes for gir-der-to-cantilever beam bolts, was developed in the same manner as
.<-_____________________________.- __ m_. _ . _ _ _ _ - _ _ . __
219 5 Table 7.8. Variables considered in Table 7.8 include both methods of extrapolation of ANDTT, evaluation by ASME K Ic and KIR ""* 8 ' C"O 1 loading cases, the assumed temperature parameter Crussian distribution, I and service life. In general, the points made about Table 7.8 apply also to Table 7.9. The small best-estimate critical flaw sizes for loading Case 5 and XIR evaluation are small enough to be of concern. However, tightening of the bolting at the time of erection tends to alleviate the problem. Table 7.10 provides a susunary of the best-estimate critical flaw size as a function of NDTT,, operating tempera-ture (T), and evaluation by both AgME K IC and K IR """*8 I'# P"~ tightening to 70% of the minimum yield load, i.e., 160 kips. .It is evident that the small critical flaw sizes of Table 7.9 would be elim-insted by the application of a pretightening load as noted. The Turkey Point supports were specified to be installed according to Ref. 28 in accordance with the standards of the American Institute for Steel
- Construction. AISC practice requires that bolting conforming to ASTM A325 or A490 be preloaded to 70% of the minimum tensile strength.2s An examination of Table 7.7 shows that the estimated fracture 1
toughness for the lateral resu sints for both asthods of ANDTT shif t, l for both 4y, and KIR evaluations, and as a function of service life is consistently high. The best-estimate value at 32 EFPY is >200 ksi/in., and the smallest value (-30, Method-A ANDTT) is 80 ksi/in., which corresponds to a critical flaw size of -1.2 in. (Fig. 7.46). Thus, the lateral restraints do not appear to be the critical component in the support assembly.
. _ _ _ _ - __ i
220 y f
-References
- 1. R. Roberts, et a1, Tracture Toughness of Bridge Steels - Phase II Report, Lehigh University, Department of Transportation, Federal Highway Administration Report, FHWA-RD-74-59 (National Tech. Infor-t l mation Service PB-239-188), September, 1974.
- 2. ASME Boiler and Pressure Vessel Code, Section XI. Rules for Inser-vice Inspection of Nuclear Power Plant Components Article A-4000 3 Material Properties,1986 Edition, The American Society of Mechan- !
ical Lagineers, N.Y. l
- 3. Standard Test Method for Plane-Strain Fracture Toughness of' Metal-lic Materials, ASTM E 399, American Society for Testing and Materi-als, 1983.
- 4. J. C. Merkle, "An Examination of the Size Effects and Data Scatter Observed in Small-Specimen Cleavage Fracture Toughness Testing,"
FUREC/CR-3672 (ORNLTM-9088), Union Carbide Corp., Nuclear Division, i Gak Ridge National Lab., April 1984. l
, 5. Dean C. Krouse, Bethlehem Steel Corp., Bethlehem, PA, Letter to C.
C. Robinson, Oak Ridge National Laboratory, Oak Ridge, TW, July 20, 1988.
- 6. Gregory J. Buragino, Steven S. Hansen, and Dean C. Krouse, " Metal-lurgical Characterization of Jumbo Sizes," AISC National Engineer-ing Conference, June 10, 1988. l
- 7. Fisher, J. W. and Pense, A. W., " Experience with Use of Heavy W Shapes in Tension", Proceedings - National Engineering Conference and Conference of Operating Personnel, May 1987, p. 18-1 47.
- 8. Fisher, J. W., Pense, A. W. and Kaufmann, E. " Cracking and Tough-ness Problems in Jumbo Rolled Sectionst Occurrence and Avoidance",
NSF Engineering Research Center - Advanced Technology for Large Structural Systems publication, Lehigh University, December 1987.
- 9. C. A. Knorovski, R. D. Krieg, and C. C. Allen, Jr. , Fracture Fough- ;
ness ct PWR Component support, NUREC/CR-3009 (SAND 78-2347), Sandia National Laboratories, Feb. 1983. '
- 10. C. V. Robino et al., The tracture Behavior of A $88 Crade A and A ,
572 crade 50 Weldsents, Welding Research Council Sc11etin 330, ! January 1988. I
- 11. P. J. Konkot, stfect of Long-Time Postweld Heat Treatment on Prop-l erties of Constructional-Sceel Weldsents, Welding Research Council Bulletin 330, January 1988.
( 12. W. H. Munse, " Fatigue and Brittle Fractures," p. 4-1 to 4-19 in Structural Engineering Handbook, edited by E. H. Caylord and C. N. Caylord, McGraw-Hill Book Co., N.Y., 1968.
- 9 -- ' * *
- ^
221 ( t
- 13. Isa Sihwell, " Florida Power & Light Company Turkey Point Units 3 &
4 (Nuclear) Containment Building Interior Civil Structural Design," Bechtel Corporation, 1967-1968.
- 14. J. B. Sun, Florida Power and Light Co., letter to R. D. Cheverton, Oak Ridge National Laboratory, dated June 16, 1988, (FRN-88-447).
- 15. J. B. Sun, Florida Power and Light Co., letter to R. D. Cheverton, Oak Ridge National Laboratory, dated July 13, 1988, (FRN-88-537).
- 16. Isa Sihweil, " Turkey Point Nos. 3 & 4 Calculations, Reactor Sup-ports, Calculation Sheets 1-17," Bechtel Corporation Job. No.
5610, Oct. 31, 1966.
- 17. I. S., " Turkey Point 3& 4 Reactor Support Cantilever Beam and R. S. Etaction, Calculation Sheets 1-11, 14-20, Bechtel Corporation Job No. 5610, Feb. 8.
- 18. Isa Sihveil, " Turkey Point Unit 3 & 4 Reactor Support (Calculation Sheets A-B) and Reactor Vessel Supports -Design Loads (Calculation Sheets 1,2,4,5,7,7a,7b,8-21,21a,22-26)," Job No. 5610, Feb. 8, 1968.
- 19. Isa Sihweil, " Turkey Point 3 & 4 Reactor Supports - Revision & R.
( S. Design, Calculation Sheets 1-4," Bechtel Corporation Job. No. 5610, Feb. 8.
- 20. I. S. Sihweil, " Turkey Point 3 & 4 R. S. Design, Calculation Sheets 1-4, 4a-4f, (3 unnumbered), 5-18," Bechtel Corporation Job No.
5610, Feb. 1968.
- 21. J. C. Merkle, " Stress-Intensity Factors Estimates for Part-Through Surface Cracks in Plates Under Combined Tension and Deading,"
- p. 3-22 in Quarterly Progress Report on Reactor Safety Programs Sponsored by the Division of Reactor Safety Research for July - September 1974, ORNL/TM-4729, Vol. II, Union Carbide Corp.
Nuclear Div., Oak Ridge Natl. Lab.
- 22. R. D. Cheverton et a1., Applicability of LEFM to the Analysis of PWR Vessels Under LOCA-ECC Thermal Shock Conditions, NUREC/CR-0107 (ORNL/NUREG-40), Union Carbide Corp. Nuclear Div., Oak Ridge Natl.
Lab., October 1978.
- 23. R. D. Cheverton and S. E. Bolt, " Pressure Vessel Fracture Studies Pertaining to a PWR, I4CA-ECC Thermal Shocks Experiments TSE-3 and TSE-4 and Updata of TSE-1 and TSE-2 Analysis," ORNL/NUREC-22 Uniori carbide Nuclear Division, Oak Ridge National Lab., December 1977.
- 24. Hiroshi Tada, Paul C. Paris, and George R. Irwin, The Stress Anal-( ysis of Cracks Nandbook, Del Research Corporation, Millertown, PA, 1973.
222 t
- 25. A. F. Liu, "St eess Intensity Factor for a Corner Flaw," Engineering Fracture #cchanics. Vol. 4, p. 175-179 (1972).
- 26. J. C. Newman, Jr. and J. S. Raju, " Stress-Intensity Factor Equa-tions for Cracks in Three-Dimensional Finite Bodies," National Aeronautics and Space Administration, NASA Tech. Memo, 83200, August 1981.
- 27. W. F. Brown, Jr. and John E. Srawley, " Plain Strain Crack Toughness Testing of High Strength Metallic Materials", ASTM STP 410, Ameri-can Society for Testing and Materials, Philadelphia, 1966, Fig. 9,
- p. 15.
- 28. Specification for Structural Joints Using ASTM A325 or A490 Bolta, American Institute of Steel Construction, Section 5 Installation, April 26, 1978, p. 5-214.
l l I
. 1 i
l
)
l l ( l
223 l-
\
Table 7.1. Calculated dpa rates around the Turkey Point reactor vessel supports 1
. dpa/s (prior to dpa/s (beginning with point location October 1, 1983)- January 7, 1984)
- IM' #M E 3 1.0 MeV E 3 0.11 MeV E 3 1.0 MeV E 3 0.11 MeV 1_ 66,35 3.844-13 5.909-13 2.103-13 3.213-13 2 61,60 4.930-13 9.123-13 2.750-13 5.010-13 3 41,62 6.523-13 1.163-12 3.802-13 6.683-13 4 51,67 3.492-13 6.996-13 2.073-13 4.066-13 l 5 58,64 2.088-13 4.713-13 1.204-13 2.787-13 6 41,78 2.297-13 4.666-13 1.440-13 2.830-13 ;
7 51,65 8.888-14 2.009-13 5.816-14 1.249-13 i 8 45,100 4.319-14 1.272-13 2.705-14 7.640-14 9 62,78 3.829-14 1.141-13 2.296-14 6.621-14 i 1 10 3.582-14 2.052-14 I
' Steel displacement cross-sections and calculation procedures ;
j ,. for dpa/s are based on ASTM E 693-79. .I b Read 3.844-13 as 3.844 x 10-13, c ruel cycles 1 through 8 ended on October 1, 1983. d Fuel cycles 9 through 32 EFPY (projected) began with January 7, 1984.
' Point 10 values were determined by multiplying the values at Point 9 by the ratio of fluxes (>0.11 MeV) of Point'10 to Point 9 given on-Fig. 7.13.
1 e g e T
- se - j 224 i
','~ Table 7.2. Time-veighted dpa rates (E > 0.1 MeV) 4 and dpa as a function of fuel cycle !
Data through to. Data through to October 1, 1983 August 3, 1988 Data through to 8 fuel cycles 7 fuel cycles 32 ETPY Data 7.04 EFPY 11.79 EFPY Total time lapses point T;tal time lapset Total time lapse: 10.091 + 8 s No. 2,020 + 8 s 3.719 + 8 s
"- ~---- Mean dpa
- Mean Mean dpa/s dpa dpa dpa/s dpa/s ,
4 6.996-13 1.553-4 5.815-13 2.163-4 4.719-13 2.302-4 8 1.272-13 2.824-5 1.067-13 3.969-5 8.773-14 8.853-5 ' 9 1.141-13 2.533-5 9.02-14 3.525-5 7.67-14 7.744-5 i 10 3.582-14 7.95-6 2.965-14 1.103-5 2.393-14 2.415-5 8 Read 6.996-13 as 6.996 x 10~13 bMean dpa/s were determined for 11.79 EFPY and 32 EFPY by time weighting the values given in Table 7.1. t i I,
)
l
-n- ,
(
325 l l 1 f
\..
Table 7.3. Susanary of estimated radiation damage. shif t, ANDTT, as a function of extrapolation method and service life for each component ANDTT, 'F Component 7.04 EFPY 11.79 EFPY 32 EFPY ; I Aa Ba A B A B I Cantilever beam b 30 a 40 130 75 1 Point 9, L = 13 in. l l Cantilever beam L 20 a 30 30b 55 Point 10, L = 17 in. c cirder-to-cantilever beam b 45 40b 65 70 75 bolt - Point 4 ' Lateral restraint b 30 30h 40 130 80 Point 8 8HFIR data extrapolation methods A and B. b Extrapolation method A yields questionable values below a ANDTT shift of 50*F. ( 4 4 l (
226' Table 7.4. Summary of Turkey Point vessel and vessel-support loads Vessel loadsa Number of Load per Loading condition Force, kips support arrays a pport ar m Moment, assumed reacting kips kips.ft
- 1. Operational dead load (DL) V = 1684 3 V = 562 only
- 2. Dead load (DL) and design V = 1802 6 V = 300 earthquake (E) (operating H = 200 3 H = 66 basis earthquake, 0.3 E')
- 3. Dead load (DL) and no loss V = 2105 6 V = 350 of function earthquake (E*)a H = 623 3 H = 203 (Safe shutdown earthquake, 55 E) 4 Dead load (DL) and vertically V = 3074 6 V = 1418 (max) acting accident load (LOCA-V) M = 21000
- 5. Dead load (DL) and horizon- V = 1684 6 V = 280 tally acting accident load H = 900 1 H = 900 (max)
( LOCA-H ) ..
- 6. Dead load (DL) and 45* hori- V = 2666 5 V = 1084 zontally acting accident load H = 635 (LOCA-45' H) M = 14900 6 H = 635 aV is vertical acting (includes dead load) ,
H is horizontal acting M is moment b A support array includes three cantilever beams. See Figs. 7.3 and 7.4. l
j , U" _
~.
m e w . c ) i r i 3 6 re a ks t v ( 6 6 eem 1 1 g ml a n mie ytb i
) d sn n na oc o/
I 3 0 3 e - k b
. N ( 5 2 0 m -
1 1 t
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