ML20080H044
| ML20080H044 | |
| Person / Time | |
|---|---|
| Site: | Brunswick  |
| Issue date: | 08/24/1983 |
| From: | Cook F, Perry R, Toland R CAROLINA POWER & LIGHT CO. |
| To: | |
| Shared Package | |
| ML20080H032 | List: |
| References | |
| NUDOCS 8309200465 | |
| Download: ML20080H044 (159) | |
Text
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RESPONSE
TO NRC (FRANKLIN RESEARCH CENTER) REQUEST FOR ADDITIONAL INFORMATION MARK I CONTAINMENT LONG TERM PROGRAM PLANT UNIOUE ANALYSIS REPORT STRUCTURAL EVALUATION FOR g-- BRUNSWICK STEAM ELECTRIC PLANT l UNITS 1 & 2 - Prepared By: i b Q h-N
. R.H. Toland Date
/ WeMu 2l24h3 R. F. Perry D&te "
8'2b$ F. A. Cook Date E [3 K. L. Bar Date QA Review: Ah 8 8J fJ. Freeman 'Datd Approved By:
/ b d h#- h '
.- L. R.' 5'cott Date N
8309200465 830915 DR ADOCK 05000 2
( INDEX Responses To Franklin Research Center Item Total Sheets Listing of Questions 6 RAI-1 and 2 2 RAI-3a and Attachment (3 drawings) 5 RAI-3b, 4a, 4b, 5, 6, 7a, 7b, 8, 9 and 10 12 RAI-lla and Attachment (6 pages) 8 RAI-12a and 12b 2 RAI-13 and Attachment MPR 751 (45 pages) 47 RAI-14 1 RAI-15 and Attachment (27 pages + 3 sheets) 18 RAI-16, 17, 18, 19a, 19b, 20, 21 and 22a 11 RAI-22b and Attachment (DR 1281A-15 pages) 17 RAI-23, 24, 25a 4 RAI-25b and Attachments (25b-1, 1 sheet and 6 25b-2, 2 sheets) RAI-26 1 RAI-G1 and Attachment (1 page) 4 RAI-G2 1 RAI-C3a and Attachment (2 pages) 5 RAI-G3b, G4, G5 3 I g 9- g- - - - - - - - - . y---,- - --
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( t MARK I CONTAINMENT LONG TERM PROGRAM PLANT UNIOUE ANALYSIS' REPORT - STRUCTURAL EVALUATION REQUEST FOR ADDITIONAL INFORMATION
,FOR BRUNSWICK UNITS 1 & 2 (FRANKLIN RESEARCH CEh'TR)
RAI 1 Indicate whether the vent penetration in the drywell has been analyzed ind whether the calculated stresses are within the allowable specified in the criteria (1). RAI 2 Provide a summary.of :.h: anslysis with regard to the (Torus) vacuum breaker piping systems and the vacuum breaker valves; indicate whether they are considered Class 2 components as (-- required by the criteria (1). RAI 3a Provide a summary of the. analysis for each safety relief valve (SRV) discharge piping which should include the analytical model with piping and supports, from nozzle at main steam line to discharge in suppression pool, discharge device, and its supports. RAI 3b Also, the information should indicate that the time history has been used for discharge thrust loads, and spectrum analysis or dynamic load factors for other loads. Justification should be provided if the above criteria are , not met. RAI 4a Provide a summary of the. analysis with regard to the piping systems and supports which provide a drywell-to-wetwell pressure differential, and also provide classification for these piping systems as essential or. non-essential. - RAI 4b Indicate whether the pumps and valves associated with these piping systems are active or inactive components. b
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,n Ee RAI 5 With reference to Table 1 of Appendix B, indicate whether all ;
loads have been considered in the analysis and/or provide ' justification if any load has been neglected. RAI 6 With reference to Table 1.4.2 in the PDA report (4), provide justification for not considering the following load factors in load cases 1,16, 52, 64 and 65, respectively: 1.7 HL, 1.1. VEQ,1.1 LOCA P,1.7 HL and 1.3 HL as required by the criteria (1). RAI 7a Indicate whether the liner has been analyzed in accordance with the boundary conditions described in Section 5.6 of the criteria (1). RAI 7b Also, indicate that the load combinations defined by Table 5-3 (1) with all load factors taken as unity have been considered. Provide justification if the above criteria are not met.
'~
RAI 8 Justify the reasons for.using a damping ratio of 5% for structural steel instead of 4% for welded structures for
'~' SSE as recommended by NRC Regulatory Guide 1.61 (5).
RAI 9 Provide information showing that the equivalent static and dynamic load f actor analysis for the Torus has yielded conservative results compared to the time history required by the criteria (1). RAI 10 Provide and justify the reasons for not considering a 1800 segment of the Torus in order to determine the effects of seismic and other nonsy'm metric loads. RAI lla Indicate whether each Torus attached piping and its supports have been classified as Class 2 or Class 3 piping, Class 2 or Class 3 component supports, essential or non-essential piping systems. RAI llb Also, indicate whether a pump or valve associated with the piping mentioned above is an active or inactive component, and is considered operable. a s, O m++-ev+* -- ,-w--,e, . , . .. -. ,- , - . . . , - - - - - - ,. ..-y. .--,- . - , - . -,r- -.-
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(a . RAI 12a Provide a summary of the analysis of Torus attached piping consisting of an analytical model which represents piping and supports from Torus to first rigid anchors (or where
- the effect of Torus motion is insignificant), and classi- .
fication of piping systems as essential or non-essential for each load combination. RAI 12b Also, indicate whether a response spectrum or time history analysis for dynamic effectg of Torus motion at the attach-ment points has been considered. RAI 13 Provide details of fatigue analysis for piping systems. > RAI 14 Provide justification for the validity of analyzing SRV piping , in two separate lengths (above and below the vent header) as indicated in Section 2.3.1.1 of PUA report (4). RAI 15 With reference to Section 3.8.4.2 of the PUA report (4), ( provide justification for using the SRSS method for combining responses due to SRV and chugging loads, ( . specifically indicating whether the stress intensity valua corresponds to 84% probability of non exceedance
- as determined from the cumulative distribution function.
RAI 16 Indicate whether the fatigue usage factors for the SRV piping and the Torus attached piping are sufficiently small so that a plant-unique fatigue analysis is not warranted for piping. l The NRC is expected to review the conclusions of a generic j presentation (6) and determine whether it is sufficient for ! each plant-unique analysis to establish that the expected usage factors for piping are'small enough and do not vartant a plant-unique fatigue analysis of the piping. l RAI 17 The ASME Code provides an receptable procedure for computing fatigue usage when a member is subject to cyclic loadings of random occurrence, such as might be generated by excitations from more than one type of event (SSE and SRV discharge, for example). This procedure requires correction be made to the . j stress-range amplitudes considered and to the associated number of cycles, in order to account for the interspersion of stress
- cycles of unlike character. State whether or not the reported usages reflect use of this method. If not, indicate the effect on reported results. .
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r C Y' - RAI 18 During foal swell and seismic events, local concentrated loads are applied to the suppression chamber a't eight vent locations, 450 apart. For example, on page 1-31 (4), it is stated that.the vent header load during foal swell is 342 kips. In the overall analysis of the ruppression chamber, these descrete loads 'are represented as a uniform line load (of equal total magnitude) instead of invoking the Wilson 2 capability to expand the load in Fowner series with approximating load peaks at eight equally spaced stations. Develop the justification of your procedure in greater detail. RAI 19a Provide justification for neglecting the discontinuities in the overall structure of the suppression chamber (a 16-sided polygonal ring). RAI 19b Also, provide justification for the implicit assumption that stress distributions and magnitudes so obtained conservatively represent those of the actual structure. In particular, address the situation at joints between adjacent straight segments. r"
- t. .
SU- RAI 20 Section 1.5 states that the thermal effects were considered, assuming a linear temperature gradient between the liner and, the exterior surface of the concrete. Although a linear temperature distribution is substantially correct for steady state conditions in the wall structures with concentric inner and outer boundaries, it is not theoretically correct for the
! case of a thick section with a circular inner boundary and approximately square outer one. Discuss the adequacy of this l assumption to predict the onset of thermal cracking in the l concrete.
RAI 21 In Section A.l.1, a normal mode and frequency response analysis for CO loading, made to assess effects of fluid /struc-ture interaction, is described. Provide the total number of modes used in the summation, the total percentage of modal
- effective mass thus included, and discuss provisions which may be incorporated in the cocputer program to account for effects of the unsummed portion of the modal effective mass.
RAI 22a Provide a summary of the results for the submerged structures listed 12 Section 2.1.1. (b), page 2-1 and contained in the i calculation books listed in' Appendix A-2 (2 & E), pa'ges 2-21
-_ and 2-11 (4).
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RAI 22b Also include the results of the strainer analysis performed (~ by the supplier, as referenced in Section 2. 3.1.2 (4). RAI 23 Section 3.8.4. 3, page 3-47 and Item 4 of Table 3.4.1.1 (4), second page, indicate that the limits on the local membrane stress intensity and on primary - menbrane plus primary - bending stress intensity were modified by using 1.3 Smc in place of S ac and certain restrictiions are mentioned in a note related to this section. Conform whether these limitations have been applied in the analysis. RAI 24 The stiffness matrix for the downcomer/ vent header intersection given in Table 3.6.1.1-1 (4) and used for the beam model analysis (Section 3.6.2) shows the diagonal terms only, i.e., no-off-diagonal terms indicated. Justify and state the assumptions for not including other significant off-diagonal terms. RAI 25a The downcomer/ vent header finite element model shown on Figure 3.6.1.1-12 (4) indicates that large elements with non-conventional aspect ratio were used at some regions of discontinuities, such as the stiffening plate regions attached to the downcomer and to the vent header. Indicate the calculated stress intensity values at such locations. RAI 25b Also indicate on Figure 3.6.1.1-12 and -13 (4), the locations corresponding to the high stress intensity values given in Table 3.8.4. 3.1 (4). RAI 26 Provide a justification for using a stress concentration factor of 2, stated on page 3-48 (4). RAI G1 With reference to Section 1.5 in the PUA report (4), justify the assumption.that the fluid structure interaction and dynamic load factors for post-chugging are the same as those for condensation oscillation. RAI G2 Provide informa*. ion indicated that the water mass has been included in the seismic analysis of the Torus. Provide justification if it has not been considered. RAI G3a With reference to Section A-1.2.5 in the PUA report (4), provide information on the stress results for condensation oscillation load at modal points listed in Table A-1-10 (4). C
- ---- - _ - -_ w_ _ T.AI G3b Also provide justification for considering those stresses which yielded dynamic load factors greater than 1.45 not reliable or typical. RAI G4 Provide details and schedules for the relocation of the electric penetration box which was found to be overstressed as indicated in Section 2.3.2.9 of the PUA report (4). RAI G5 With reference to Table 2.3.2-3 of the PUA report (4), indicate the schedule for modifications of the platform structure.
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analyzed and whether the calculated stresses are within the , allowable specified in the criteria (1).
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RESPONSE
The vant penetration in the drywell was not analyzed for the .- following reason. The major load transmitted to the drywell k penetration by the vent is pool swell impact and drag acting on ; the bottom of the vent system. It was determined that a modi-fication, consisting of additional upper columns at the vent / ring header intersection, would be required to help resist these loads. To assure adequacy of the modification, the pool swell '
. s loads on the vent and the vent / ring header intersection were ,
t , calculated very conservatively. Using the conservative pool \ f.,
\- swell loads, the ratio of calculated stress to allowable stress in the vent ranged between 8% and 19%. These ratios represent an upper bound and could be significantly reduced by removing conservatism from the load definition. Since the vent transmits only a small portion of its allowable load back to the penetration, there is no need to perform an analysis of the penetration.
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b^ Kr,1 2 Provide.a rummary,of'th( analysis-yith regard to the.(Torus) L ._c , , s vacuum breaker piping systems and the vacuum breaker valves;
% % ,t indicate whether they r.re cons,idered C,la. as 2 coroonents 3
as s' N .D , " '<,'"' required by 1.ha~ criteria (1)._\ ,
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y N > q , Qwl, _) , in accordante with'ths. ASE[ Code Section~ III, Subsection B, s - m 1.--_- .% s - . s, i.e., Cisse 1, without N-stamps. The drywell-wetwell vacuum xy, ,- . _ r - i'
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g ., RAI 3a Provide a sunusary of the analysis for each safety relief valve i_ (SRV) discharge piping which should include the analytical l model with piping and supports, from nozzle at main steam line to discharge in suppression pool, discharge device, and its supports.
RESPONSE
. r The safety relief valve (SRV) discharge piping analysis is divided into two models. The' upper model includes the primary steam line with the attached SRV discharge piping lines in the ,
drywell down to the vent header. The lower model includes the SRV discharge piping from the vent header downstream into the suppression pool and is discussed in detail in the Plant Unique Analysis Report section 2.3.11. The upper model of SRV discharge piping including primary t
- g. steam line was analyzed for the usual thermal, dead weight, and seismic loading conditions as well as SRV discharge transients.
The piping model with supports is shown on the attached isometric drawing, sheets 514, 624 and 626. Seismic analysis used the dynamic response spectrum approach i in accordance with NRC, IE Bulletin 79-07. O f' (.. f r I
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x. ISOMETRIC DRAWINGS
- SHEETS 514, 624 and 626 (3 Sheets)
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CAROLINA POWER & LIGHT COMPANY BRUNSWICK STEAM ELECTRIC PLAITI UNITS 1 & 2
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i RAI 3b Also, the information should indicate that the time history has been used for discLsrge thrust loads, and spectrum analysis or dynamic load factors for other loads. Justification should' be provided if the above criteria are not met.
RESPONSE
Time-history' fluid force transient analysis of the SRV discharge lines was performed to determine the forces, moments, and stresses in the SRV discharge piping. The input for this analysis utilized the General Electric - Mark I fluid force time-history resulting from SRV fluid transient discharge analysis. To account for a single SRV discharge effect on a primary steam line, the resulting fluid transient loads from each SRV time-history analysis were statically applied one at a time to the corresponding primary steam line with a dynamic load factor of 2.0. To account for the effects of all SRV discharges simultaneously, the resulting fluid transient ' loads from all SRV time-history analyses were statically applied l s'imultaneously to the corresponding primary steam line with a maximum dynamic load factor of 2.0.
[ (; RAI 4a Provide a summary of the analysis with regard to the piping systems and supports which provide a drywell-to-watwell pressure differential, and also provida classification for these piping systems as essential or non-essential.
RESPONSE
The Brunswick plants do not have piping systems which provide a drywell-to-wetwell pressure differential. The comment is not pertinent to Brunswick since Brunswick does not provide a dryvell to vetwell pressure differential. l e r s. i e c-m- -
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1 ( i RAI 4b Indicate whether the pumps and valves associated with these piping systems are active or inactive components.
- _ RESPONSE See response to RAI 4a. '
e t C. , 6 e 1 I w w -,-s,,--- e, c-,- , , -a v,,.1- , , --g.-m-w -,--e,,- ,, , , - .- , - - g,weg,,, y,,swmeee--e , . -, --~ g-.- p- ,- m n , -
[ RAI 5 With reference to Table 1 of Appendix B, indicate whether all loads have been considered in the analysis and/or provide justification if any load has been neglected.
RESPONSE
All applicable loads, listed below, were considered in the torus shell analysis. See PUAR Section 1.3 for detail. a) Containment Pressure and Included Section 1. 3.2 Temperature and 1. 3. 3 b) Pool Swell
- 1) Torus net vertical loads Neglected Section 1.3.7 ii) Torus shell pressure Included Section 1. 3.7 histories iii) Froth impingement Neglected Section 1. 3.7
([ c) Condensation Oscillation Included Section 1. 3.9
. d) Chugging Included Section 1.3.10 e) T-Quencher Loads Included Section 1. 3.4 f) Ramshead Loads Not Applicable All loads pertinent to the vent system were taken into account in the analysis except the following:
- 1) Containment temperature and pressure loads Containment temperature and pressure loads were not significant for vent system structure. ,
i ii) Jet loads on submerged structures Jet loads on submerged structures were not applicable i to vent-header, vents and downcomers. [s- _1_
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(' ( - RAI 5 (Continued) . RESPONSE (Continued) 111) Froth impingement This load was covered implicitly by the conservatism in the impact and' drag load on the vent. All applicable loads were considered for the S/RV piping and other wet wall interior structures. ( .-- N e 0 4 6 6
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RAI 6 With reference to Table 1.4.2 in the PUA report (4), provide justification for not considering the following load factors in load cases 1, 16, 52, 64 and 65, respectively: 1.7 HL, 1.1 VIQ,1.1 LOCA P,1.7 HL and 1.3 HL as required by the criteria (1).
RESPONSE
~
The hydrostatic prescurt: was considered as dead load not live load per Ref, (1). A load factor of 1.4 for the hydrostatic pressure has been used for load cases 1 and 64 and 1.0 for load case 65. The hydrodynamic effect of the pool water is due to SRV, condensation oscillation, chugging and pool swell has already
'been included in the loading cases.
The loading combination for cases 16 and 52 should have been as follows: i, D.L + H.L + T 2 + 1.lP +\ (1.1 HEQ)2 + (1.1 VEQ)2 1.1 SRV Case 16 (Correct) D.L + H.L + T 2 + 1.lP 1 (1.1 HEQ)2 + (1.1 VEQ)2 + 1.1 Pre-chug Case 52 (Correct) The following combination has been used due to input error to the load combination program. D.L + H.L + T2 + 1.lP 1 (1.1 HEQ)2 + (VEO)2 + 1.1 SRV Case 16 (Incorrect) D.L + H.L. + T (1.1 HEQ)2 + (1.1 VEQ)2 + 1.1 Pre-chug 2 + P i\ Case 52 (Incorrect) - Q:
'( RAI 6 (Continued) RESPONSE (Continued) When load cases 16 and 52 are compared with load cases 18 and 54, shown below, it is obvious that the internal forces due to load
, case 18 are greater than those due to load case 16; the internal for~ces due to load case 54 are greater than load case 52, because post-chugging loads are higher than pre-chugging (see Section 1. 3.10.1 of PUAR). In our opinion, it is not required to correct the input error in load cases 16 and 52 since they are not governing load combinations.
D.L + H.L + T 2 + 1.lP 1 (1.1 HEQ)2 + (1.1 VEQ)2 1.1 SRV i 1.1 Pre-chug (Case 18) s. D.L + H.L + T 2+1.1P+fl.1HEO)2+(1.1VEQ)2+1.1 Post-chugging (Case 54) d .
-2
i .. RAI 7a Indicate whether the liner has been analyzed in accordance with the boundary conditions described in Section 5.6 of the criteria (1).
RESPONSE
The liner has been analyzed in accordance with the boundary conditions described in Section 5.6 of the criteria (1). e 9 e
- c. -
m ,..-..-. .- . _ , - . , - ,, - , , . - ., . . .
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(' RAI 7b Also, indicate that the load combinations. defined by Table 5-3 (1) with all load f actors taken as unity have been considered. Provide justification if the above criteria are not met. RESPONSE - The load combinations defined by Table 5-3 (1) with all load
. factors taken as unity have been considered.
Condensation oscillation, chugging and S/RV exert upward pressure on the liner due to the oscillatory nature of the loads. The maximum suction (negative pressure) on the liner, occurring during IBA, is less than 28 psi including the effect of fluid interaction and dynamic load factor and assuming unit load factors in load combination. The stresses in the liner were obtained for the maximum suction ('\" of 28 psi using finite element analysis and assuming that only the anchors provide support for the liner as per Section S.6 of c 1*eria (1). A 30" by 30" liner plate with appropriate support condition was used in the analysis. The maximum tensile stress in the liner due to local bending of the liner is 12.83 ksi and the maximum compressive stress is 14.85 ksi. The total conservative stress in the liner when seismic and
~
thermal loads are also added are 24.25 kai ccapressive tnd l 24.40 tensile. These stresses are below yield stress of the liner.
s, RAI 8 Justify the reasons for using a damping ratio of 5% for structural steel instead of 4% for welded structures for SSE as recommended by NRC Regulatory Guide 1.61 (5). RESPONSE . The 5% damping value (pg. 3-37) was used for the Reactor Building structural steel in the seismic analysis of the coupled Reactor Building / vent system model. This value was obtained from the Brunswick FSAR, Table C-1. In accordance with the " Plant Unique Analysis Application Guide" (GE document NEDO-24583-1), it is acceptable to use the methods employed in the original plant design for those loads, such as seismic, that are not redefined by the i Mark I Containment Program. Therefore, a damping ratio of 5% for welded steel structures is acceptable for SSE. However, in spite of this, 4% damping was used for the vent system for the SSE. i e t l l t
- ('_-
s
f. RAl 9 Provide information showing that the equivalent static and dynamic load factor analysis for the Torus has yielded conservative results compared to the time history required by the criteria (1).
RESPONSE
. Time history analysis was used for hydrodynamic loads to study the effect of fluid structure interaction and calculate dynamic load factors. It was noted that the hydrodynamic effect on the reinforced concrete torus is minimal as compared to a steel torus.
For seismic analysis, results (maximum accelerations) from the original analysis, based on a response spectra method, was used in accordance with the Section 6.2-a of Reference (1) which justifies the use of such results. ( l l l l l s~ i
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I t
. - ._ = . .- .
(- RAI 10 Provide and justify the reasons for not considering a 1800 segment of the Torus in order to determine the effects of seismic and other nonsymmetric loads. 4
RESPONSE
The Wilson II Program, developed by Ghosh & Wilson in 1969, was modified and verified by UNITED for use in the analysis of the torus under quasi-static seismic and other nonsymmetric loads. The concrete portion of the axisymmetric torus is discretized using axisymmetric ring elements of triangular and quadilateral ! cross section. The liner and the reinforcing bars are modelled using truncated conical shell elements. The nonsymmetric load < is expanded in terms of Fourier series to account for variation of the load in the global hoop direction. The solution algorithm . used in the program solves a three dimensional structure by combining the results from each coefficient in the Fourier series. I
'The stresses for each element at several locations around the global hoop direction are obtained. The complete torus is, therefore, evaluated in the analysis.
e
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RAI lla Indicate whether each Torus attached piping and its supports have been classified as Class 2 or Class 3 piping, Class 2 or Class 3 component supports, essential or non-essential piping systems.
. lib Also, indicate whether a pump or valve associated with the piping mentioned above is an active or inactive component, and is considered operable.
RESPONSE
All torus internal piping has been analyzed in accordance with the requirements of the ASME Boiler & Pressure Vessel Code Section III, Subsection NC (Class 2). All new support components have been designed in accordance with ASME Section III, Subsee-tion NF criteria. There are no active components attached to the torus internal piping. All active components associated with ( these piping systems are located outside of the torus and because 4 of the concrete backing, are unaffected by torus induced loads. The torus section strainer qualification is attached to the response of question 22b. The application of these criteria in this manner and incorporation of the required modifications result in the torus internal piping being designed in complete accordance with l ASME Section III requirements. l e
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RAI 12a Provide a summary of the analysis of Torus attached piping '
consisting of an analytical model which represents piping and supports from Torus to first rigid anchors (or where the effect of Torus motion is insignificant), and classi-fication of piping systems as essential or non-essential for each load combination.
RESPONSE
The Brunswick Torus is a steel lir.ed, reinforcad concrete structure. It is very rigid (unlike all other Mark I toruses which are steel structures); no significant motion is input to torus attached piping at anchor points where each pipe penetrates the torus. Therefore, no analysis for torus attached piping was required. (- I e e _. -,. -- -s,_-, -.,_w , , , ___ _ . _ - ,- , _ _ __ _ _. _ _ - , , _ , _, -7, , _ . _ . . - - .-
f. RAI 12b Also, indicate whether a response spectrum or time history analysis for dynamic effect of Torus motion at the attachment ' points has been considered.
RESPONSE
See response to RAI 12a. i e D 4 e n
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J f;- RAI 13 Provide details of fatigue analysis for piping systems. i
RESPONSE
Details of the generic augmented fatigue evaluation performed
,- for Torus Attached and SRV Piping Systems is presented in Section 2.0 and Appendix A of MPR-751, Attachment RAI-13.
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ii ( s MARK I CONTAINMEhT PROGRAM MPR-751 (45 Pages) E m CAROLINA POWER & LIGHT COMPANY , BRUNSWICK STEAM ELECTRIC PLANT I UNITS 1 & 2 l i I l l Attachment RAI 13 a
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! , PROGRdm y
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{ 2 MARK I CONTAINMENT PROGRAM i AUGMENTED CLASS 2/3 FATIGUE EVALUATION r METHOD AND RESUCTS FOR TYPICAL TORUS ATTACHED AND SRV PIPING SYSTEMS
-1 a
,i November 1982 1
F MPR-751 (; .
-f l
I 4 4 1 1 i - 1 1 ! t i : Prepared by , I
'! MPR ASSOCIATES j Washincton, D.C.
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MARK I CONTAINMENT PROGRAM.
. AUGMENTED CLASS 2/3 FATIGUE EVALUATION METHOD AND RESULTS FOR TYPICAL TORUS ATTACHED '
AND SRV PIPING SYSTEMS November 1982
'- MPR-751 e
(
- y. .
Prepared by MPR Associates
- Washington, D.C.
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(., ' k~ ACKNOWLEDGMENT This repore was prepared by MPR Associates with assistance provided through the joint efforts of the Mark I Architect / Engineers. The specific organizations which contributed to the precaration of this report and/or provided fatigue evaluation results are as follows: Sechtel Power Corporation Sechtel Power Corporatis.,7 Gaithersburg, Maryland San Francisco, California , EDS Nuclear United Engineers and Constructors San Francisco, California Philadelphia, Pennsylvania NUTECH Southern Company Services
'? San Jose, California Birmingham, Alabama Teledyne TVA Waltham, Massachusetts Knoxville, Tennessee 1
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DISCLAIMER OF RESPONSIBILITY l t(either the General Electric Company nor any of the contributors to this document makes any warranty or representation (express or , it:: plied) .with respect to the accuracy, completeness, or usefulness - of the information contained in this document or that the use of such information may not infringe privately owned rights; nor do they assume any responsibility for liability or damage of any kind which may result from the use of any of the information contained in this document. s I I 6 e
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i I, . . ( V s.. . . . TABLE OF CONTENTS
1.0 INTRODUCTION
AND
SUMMARY
2.0 DISCUSSION OF EVALUATION METHOD - j 3.0 RESULTS AND CONCLUSION
. 4.0 REFER 5NCES APPENDICES A. . Augmented Class 2/3 Fatigue Evaluation Method, Tables and Nomenclature B. Effect of Thermal Gradient Stresses on the Fatigue Life of SRY Discharge Piping r
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1.0 INTRODUCTION
AND
SUMMARY
The fatigue approach being followed in the Mark I Long-Term Containment Program for torus piping attached and SRV discharge piping is to follow appitcable Class 2 piping rules of. Section III of the ASME Code (Reference 1). .The Code requirements address cyclic thermal and anchor motion stresses. In August 1981 the NRC raised a concern regarding the , , cyclic stress due to the Mark I cyclic mechanical loads. The Mark I
' Owners Group and GE met with the NRC to discuss this concern in September 1981 and proposed that a task group be assigned to evaluate l
whether the Mark I Containment Program loadings and acceptance criteria i , already contained sufficient margin for fatigue effects and to define a F course of action for a generic Mark I response to the NRC concern. The approach was to develop a method for piping fatigue evaluation and to I. apply the method on piping systems representative of the most limiting in Mark I plants. It was agreed that the fatigue approach should be h ( developed along the lines of the Class 2/3 piping design method since these methods were already being applied for Mark I Containment Program' ( plant-unique torus piping analyses, } The Mark I fatigue task group was comprised of several of the Mark I N 7 A/E's and coordinated by the General Electric Company. The objectives. ! of the group were as follows: 'f.t i T O 5 Determine the loading cycles 'and loading cycle combinations appli-cable to the Mark I Containment Program load definitions. 1
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(~, t 0 Develop a methodology for performing an " augmented" Class 2/3 fatigue evaluation on torus piping to account for cyclical mechani-cal loads. O Provide the methodology to the Mark I A.rchitect/ Engineers for their use in evaluating representative limiting piping systems. O Prepare a generic evaluation of the torus attached and SRV dis-charge piping with regard to mechanical fatigue.
.. The differing analytical approaches used for Mark I piping analysis required close coordination between the A/E's in developing the fatigue methods to ensure that the fatigue jnethod could be applied with analysis l results available from each A/E's plant-unique piping analyses.
The final augmented Class 2/3 fatigue evaluation method described in this report reflects the input and coments'from all Mark I A/E's and l t guidance provided by the General Electric Company. j
?
s. Tables 1-1 and 1-2 indicate the scope of the fatigue evaluations
. performed with the final fatigue procedure. As can be seen, essentially all of the Mark I plants were addressed. A total of 36 piping systems - were included in the evaluation. SRV discharge lines comprised 30% of the systems.
i ! The results of the evaluation of the piping for fatigue due to mechani-cal loadings in addition to the'rm.1 and anchor motion show that fatigue usage for Mark I torus piping is generally low, with fatigue usage typi-cally well below 0.3. None of the lines has a fatigue usage greater ' than 0.5. It should be noted that the u H a c. k
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- stress results for the most limiting piping systems and locations were selected for each plant. Thus, these results are representative of the most limiting locations for fatigue usage and the remainder of these torus pip'og systems would have even lower fatigue usage. .
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~ . .
l TABLE 1-1
SUMMARY
OF PLANTS AND TORUS ATTACHED PIPING SYSTEMS If4CLUDED IN MARK 1 CONTAINMENT PROGRAM AUGMENTED CLASS 2/3 FAilGUE EVALUATION l NUMBER OF TORUS :
, PLANT TYPE ATTACHED PIPING SYSTEMS ANALYZED j
- Hatch 1 and 2 BWR-4 2 Peach Bottom 2 and 3 BWR-4 2
- Cooper BWR-4 4 Oyster Creek BWR-2, 3 Pilgrim BWR-3 3
-. Millstone 1 BWR-3 1 (i. Vermont Yankee BWR-4 1 l Brunswick 1 and 2 BWR-4 Quad Cities 1 and 2 BWR-3 1 .
and Dresden 2 and 3
~
l- Duane. Arnold BWR-4 1 Browns Ferry 1, 2 BWR-4 3 and 3 .
. Nine Mile Point 1 BWR-2 2 l
l
. TOTALS 25 I
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f b.s TABLE 1-2
SUMMARY
OF PLANTS AND SRV DISCHARGE PIPING SYSTEMS
'l INCLUDED IN MARK i CONTAINMENT PROGRAM AUGMENTED CLASS 2/3 FATIGUE EVALUATION i
PLANT TYPE NUMBER OF SRV : DISCHARGE PIPING SYSTEMS ANALYZED
,j Hatch 1 and 2 BWR-4 1 Peach Bottom 2 and 3 BWR-4 1 Cooper. BWR-4 1 Fitzpatrick BWR-4 3 Millstone 1 BWR-3 1 Brunswick 1 and 2 BWR-4 1 k.. Fermi 2 ,
BWR-4 1 7
, Monticello BWR-3 1 i
1 Browns c arry 1, 2 BWR-4 1
! and 3 i
TOTALS 11 l
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_ _ ~ . - - . . . - . . . . -_- - . -.. .._ .. r (' b 2.0 DISCUSSION OF EVAi.UATION METHOD i l This section contains a description of the evaluation procedure developed for the Mark I torus piping fatigue evaluation: (i) the basis for the method, (ii) the key assumptions made and (iii) the principal .
, conservatisms in the method. Each of these subjects is covered in a separate section below. The abbreviations and nomenclature used in this section are defined in Section 2.0 of Appendix A. !
2.1 Basis for Evaluation Method In developing the basis for the Mark I piping fatigue evaluation two approaches were considered: (i) use of the ASME design fatigue rules for Class 1 piping; or (ii) use of the ASME design rules for Class 2/3 i piping augmented.to include both mechanical and thermal cyclic stresses k in the applicable evaluation equation. The latter approach was followed since the Mark I analysis guidelines already specify that Class 2/3
- f. piping design should be used for thi plant-unique evaluation and thus I
the results of plant-unique evaluations which were already available could be used. In this way the considerable effort of reanalysis of l piping with the Class I rules could be' avoided. There are two ASME Class 2/3 piping design equations which account for repeated loadings: Equations 10 and 11 of Paragraph NC-3652.3 j (Reference 1). As has been shown in Reference 4, there is reasonably good agreement between the C1'ss a 2/3 method and the Class 1 (previously u B31.7) method, especially below 20,000 cycle:; where all of the Mark I 1 i loadings fall. Reference 4 clso describes the basis for the Class 2/3 i design equations which were developed by Mark 1 (Reference 5). Marki j 4 k~
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developed his design equations from piping fatigue data. The stress 1 I range reduction factors, which are included in the Class 2/3 design equations, are basically a " stepped" approximation of Mark 1's equations. The augmented Class 2/3 method uses the original Marki equations (one for carbon steel and one for stainless steel). A comparison with the
" current ASME Cla.ss 1. design fatigue curve and the Class 2/3 " stepped" approximation is shown in Figure 2-1. As can be seen, there is :
reasonable agreement between the Mark 1 equation used in the augmented Class 2/3 procedure and the Class 1 design fatigue curve and the Mark 1 equation is conservative when compared to the Class 1 curve below 10,000 cycles. ;
\
Tables 2-1 through 2-3 cover the details of the evaluation method \ i including the loadings considered, the loading durations and the load j combinations. These loadings are consistent with the loadings defined
' 1 in the Mark I Load Definition Report (Reference 3) and the load '
I I comoinations specified in the PUAAG (Reference 6). As shown in [ Tables 2-2 and 2-3, two evaluations were perfomed for each piping system: (1) design break accident (DBA) plus normal operating conditions (HOC); and (ii) intermediate or small break accident I (IBA/SBA) plus NOC. Each evaluation includes both safe shutdown and operating basis earthquakes. The IBA/SBA evaluation was performed so as to envelope the loading cycles of both the ISA and SBA events. 1 The stress results for each loading typically correspond to the maximum value which will not, in general, occur for each cycle of the loading. 4 Many time history analyses of these loadings have been performed on Mark I piping systems which, show that most of the response cycles are } less than the maximum value. To account for this variation, factors j g were computed to convert the calculated fatigue cycles into effective 8, [i
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- y. 2.2 Assumotions This section lists the principal ?.ssumptions which were made in developing the augmented piping fatigue analysis met,hodology.
; O I
A typical loading history for a piping system attached to a Mark I containment includes:
' periodic SRV actuations for the life of the plant (NOC) with the total number of actuations determined for the specific plant. One combined thermal and anchor motion load is assumed to act during each actuation.
T Five operating basis earthquakes (Reference 2). One accident condition.- either DBA or IBA/SBA which includes; (i) one combined thermal and anchor motion loading, (ii) operating basis earthquake (OBE) and safe {: shutdown earthquake (SSE) earthquake stresses, and (iii) l periodic SRV actuations during IBA/SBA with the total
- number of actuations determined for the specific plant.
0 Other thermal cycles due to normal operating conditions (for example, system testing) are considered negligible. O Stresses due to thermal gradients in the wall of SRV discharge piping do not significantly contribute to fatigue usage. See Appendix B for a discussion of the basis for this assumption. ; ! I t . O G I 2-4 b b n - y g y_x:7-; y y- _ m: . -mw. c -:--- -
f t i t 0 Leads are grouped according to the combincticns listed in l l' ,
- Reference 6 considering the number of cycles for each }*
loading. Equation 11 from ASME Section III, Class 2, Paragraph NC-3552.3 is the basis for calculating fatigue stress ranges. O
- To evaluate the allowable number of fatigue cycles the Markl equations are used (Reference 4): *.
i N = (245/15)5 -- carbon steel N = (281/15)5 -- stainless steel where:
- i N = number of cycles 15 = intensified stress range in ksi ; ;
, l 0
The fatigue usage due to SRV discharge.actuations taking place !It prior to the Mark I Long-Term Containment Program is accounted 11 y!, for by including an estimated number of discharge actuations corresponding to 'he full 40-year plant lifetime. [] - LI a . 0 $ 2 Each earthquake load contains ten (10) significant response
>}
cycles (Reference 2). 4, 1 O e , Prechug and IBA condensation oscillation (IBACO) loads are !
; i assumed to act at the highest load frequency of 9.5 Her m, thus giving a reasonable estimate of the number of response "
t cycles per second for this loading. e p' s g' : .
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I i., 2.3 Conservatisms in Evaluation Method This section summarizes the principal conservatisms in the fatigue evaluation method.
- 0 Stresses from hydrodynamic loads were computed on the basi; of peak loads which bound full-scale test results.
I
' 0 Absolute summation was used to combine all dynamic responses.
O All effective full stress cycles were assumed to be in phase when two events are combined; further, events were assumed to combine peak on peak for the duration of the combined event. O Deadweight (DW) is not a fatigue load and would not normally be included in the fatigue stress summation; however, it is included in all combinations since it is a required loading for ASME Class 2 piping, Equation 11 (Reference 1) and it is (, . conservative to do so. O Stresses for operating and safe shutdown basis earthquakes are combined with most limiting DBA/ISA/SBA stresses for the appropriate numoer of full stress cycles. :
'i O
Thermal stresses for accident events are ccmbined with the appropriate mechanical stresses although thermal loadings are single cycle loads whose contribution to fatigue would not : normally be cor.sidered. i I 1 i .
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\. .. Stresses for the safe shutdown earthquake (SSE) are considered i in the fatigue analysis. Typically fatigue analyses do not l t
require consideration of SSE stresses. It is noted that j fatigue requirements in the PUAAG (Reference 6) specify that l only operating basis earthquake need be considered in the l fatigue analysis of Mark I torus shells. I G 3
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- l TABLE 2-1 l
! MARK I.LOCA EVENT AND LOADING DURATIONS (Reference 3)
LOCA EVENT DURATION (seconds)
. LOCA TYPE DBACO PRECHG/IBACO CHUG .L DBA 30 30 30 t-I e IBA (with steam-i driven feed j pumps) - 300 200 i
IBA (with motor-driven feed pumps) - 1100 200 t .
'- SBA - 900 900 ~
0 . NOTE:
- l. Since the augmented Class 2/3 fatigue method was applied generically with IBA and SBA combined as one LOCA event, the following bounding ;:
T LOCA. loading times were used in determining the fatigue loading c cycles: i. PRECHG (same as IBACO) -- 1100 seconds . CHUG -- 900 seconds , The rest limiting IBA event duration (i.e., motor-driven feed pumps) was used to determine the bounding IBA/SBA LOCA event durations. . I I. T -. u -= . . . =z.:, , , . : r u --~ ~.n... . . . . . - ~ ~ - - - - ~' - = '" -
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( TABLE 2-2 ; FATIGUE LOADING COMBINATIONS AND CYCLES FOR NOC + DBA CONDITIONS .
. I a..
COMBINATION (Note 4) CYCLES h DBACO + EQ(0) + MCDBA + PRDBA 1 DBACO + EQ(0) 9 , DBACO 890 CHUG + EQ(S) 10 CHUG , (Note 1) PRECHG 102 SRVNOC + EQ(0) + MCNOC + PRNOC (Note 3). 40 is3 SRVNOC + MCNOC + PRNOC (Note 3) (Note 2) SRVNOC (Note 2) !I i ( g NOC DBA lf s' ij s i: SRVNOC MCDBA + PRDBA jj MCNOC + PRNOC lDBACO PRECHG CHUG l EQ(0) EQ(S) 'j NOTES: 1 Ii
- 1. Number of cycles depends on dominant piping response frequency. ,g
>+
- 2. Number of cycles depends on plant evaluation of nud er of normal SRV k actuations and reactuations over the life of the plant (See -:
- Appendix A).
- J l
( 3. MONOC and PRNOC are the thermal expansion and pressure stress ranges 4.f l and are applicable to SRV discharge piping only under normal ,F operating conditions (NOC).
- 4. Pool swell exclu:!ed from fatigue analysis per guidelines contained t in the PUAAG (Reference 6). For nomenclature, see Appendix A. I y ,
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(g-TABLE 2-3 l FATIGUE LOADING COMBINATIONS AND CYCLES FOR NOC + 15A/5SA CONDITION 5 COMBINATION (Note 5) CYCLES CHUG + SRVIBA + EQ(S) + KIBA + PRIBA 1 CHUG + SRVIBA + EQ(5) 9 . CHUG + SRVI3A (Note 1)
; CHUG (Note 2) i IBACO (same as PRECHG) 10,450 SRVNOC + EQ(0) + MCNOC + PRNOC (Note 4) 50 SRVNOC + ENOC + PRNOC (Note 4) (Note 3)
SRVNOC (Note 3) NOC_ IBA/SBA MCIBA + PRIBA , SRVNOC l SRVIBA MCNOC + PRNOC IBACO l CHUG EQ(0) lEQ(S) , NOTES:
. 1.
Nuter of cycles depends on plant evaluation of the number of SRV actuations (See Appendix A). l
..I 2. Number of cycles depends og frequency of piping (See Appendix A).
- 3. Number of cycles depends on plant evaluation of number of normal SRV l
l actuations and reactuations over the life of the plant (See Appendix A).
. 4 MCNOC and PRN' O C are the thermal expansion and pressure stress. ranges
; and are applicable to SRV discharge piping only under normal . .
j operating conditions (NOC).
~
! 5. For nomenclature, see Appendix A.
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O O O O o 8 O O O O O ,' i'
. r, w N in O o .
" o Augmented Class 2/3 Fatigue Curve * .
t' # ' '.
' Hark 1 FATIGUE CURVE
- O d O
c_ , (Reference 4) I
- COMPARISON OF AUGMENTED CLhSS 2/3 FATIGUE :
i , I.IMITS WITil CLASS 1 DESIGN FATIGUE CURVE l-
+
i
.f FIGURE 2 =
y J g .. , l .i f
. f. - 1; 3 j .,
m..
. . . . . .,. . . . .g%% , #. :_e - . m m,. -,. _ .
1 .. 3- .
( r L 3.0 RESULTS AND CONCLUSIONS this sec, tion centains the results of the fatigue evaluations performed on over 30 torus piping systems. These systems were selected by each A/E as representative of the most highly stressed torus piping systems in their respective plants. Thirty percent of
- these were SRV discharge lines and the remainder were lines attached to the torus with sizes ranging from 2-inch to 24-inch.
l All torus piping systems had a fatigue usage less than 0.5. The P fatigue evaluation results, which are tabulated in Table 3-1, are sunrnarized as follows: I
] O SRV Discharge Piping:
Percent less than 0.3 fatigue usage -- 72.7% Percent less than 0.5 fatigue usage -- 100%
'b i 0
Other Torus Attached Piping: t l Percent less than 0.3 fatigue usage -- 92.0% Percent less than 0.5 fatigue usage -- 100.0% - l i A very conservative methodology has been developed for fatigue l analysis of Mark I Class 2 piping. The fact that the calculated i fatigue usage factors are 1ow coupled with the very conservative i. approach used to develop the fatigue analysis methodology shows k that fatigue is not a concern for attached piping. Thus this report answers the concern expressed by the NRC regarding the , effect of cyclic mechanical loads on fatigue. Accordingly, there . { is no need for a ec plete evaluation cf torus piping fatigue on a plant-unique basis. ' *
. A i
e 3-1 (v - 3
~~
.: m.T,' . . ',' ' *
* ' **~~ **~*'
- W ' ~
- - ~ ' " ' ' '
' ~~ ~
. . _ _ _ . . _ . - . . . ..___ _._ _ -. ___. . . _ _ . - , . . = . _ _ . , . ~ .
F~ , s-3
\
\. -
TABLE 3-1 i
' 5UMMARY OF RESULTS 0. AUGMENTED CLASS 2/3 PIPING FATIGUE EVA.UATIONS FOR REPRESENTATIVE P%RK 1 PIPING SYSTEMS i ,
I t ARCHITECT ENGINEER SYSTEM FATIGUE USAGE [ Utility (Size) NOC+DBA NOC+1SA/SBA
- Piant I
l A. Bechtel-Gaithersbure Georgia Power 1. HPCI Pump Suction 0.000 0.002 - Hatch-2 (16-inch)
- 2. Core Spray Pump 0.001 0.002 (10-inch)
- 3. SRV Discharge 0.077 0.096 (10-inch)
B. Bechtel-San Francisco . Philadelphia Electric 1. Core Spray System - 0.001 ,0.022 Peach Bottom-3 P-14-5 (16 and 14-inch) i
- 2. Inerting System 0.013 0.004 P-9-2 (20-inch)
- 3. SRV F-1-5 0.046 0.202 -
(12-inch) 1 . s t 3-2 C[ f^ t i. 3 * , . .. . : s. ., : . , < ' . s -- ' , .
' "'* ,--k'. .. ..'. 4 ..:..~ ~.w="a '~.T.
- ' *~ ~u. ~ ' "~
L , -. - ,, . .__ -,-. -. . , . . - - . , , . - - . - - - - , - . -
4' . - .. _ ._ ._. . i
= .
'g-_ j Q.,
. TABLE 3-1 (Cont'd)
I SUFFARY OF RESULTS OF AUGMENTED CLASS 2/3 PIPING
- FATIGUE EVALUATIONS FOR REPRESENTATIVE
- MARK I PIPING SYSTEMS .
.I . .-
i ARCHITECT ENGINEER SYSTEM FATIGUE USAGE :
! utility (Size) N00+D3 F NOC+15A/53A
! Plant 5
I C. EDS-Nuclear ! I Nebraska Public Power 1. Core Spray Suction - 0.113 0.149 i Cooper X227A (16-inch) I f9 I t l 2. Core Spray Suction- 0.009 0.012 X227B (16-inch) ,
, s
- 3. RCIC Turbine . 0.279 0.280 !
i e Exhaust (8-inch) ! 3 1 3 ! 4. HPCI Turbine 0.020 0.058 Exhaust (24-inch f i .
- 5. a. SRV Discharge 0.116 0.246 -
i (10-inch) : i b'. SRV Dischargt .035 .006 i (10-inch) 6 l 1
~
l l $ I l l I l r i t 3-3 ,
-} n I
j;;; , . - - ,.- - - . . . , . . _. , ..,c .- q, -.- .-
- w -==w-- w -=w- - ~ ~--~ ~ ~ - - - = ' ~ ' - - - * - * * - * ^ ' ' ' ww- *1
r' (V TABLE 3-1 (Cont'd) SUWARY OF RESULTS OF AUGMENTED CLASS 2/3 PIPING FATIGUE EVALUAT10 tis FOR REPRESENTATIVE MARK I PIPING SYSTEMS 1 ARCHITECT ENGINEER SYSTEM FATIGUE USAG: U ti lity (Size) NOC+DSA NOC+1BA/SBA Plant D. ICR Associates FPU Nuclear 1. Vacuum Relief-Type 2 0.434 0.258 Oyster creek (24 a::s 18-inch)
- 2. Demineralizer 0.087 0.067 Relief (20-inch) l 3. Core Spray Suction (12, 16 and 20-inch) 0.131 0.084 e
L e e l l I I I , e r . l ! i g t ! .- 3-4 l s . r i 1 1 L ^
-" = .
~ j. [. rs L : -~..T. }, ,* ,
v.* .M' .-% " I. ' . , ,
. _ . . _ . . . _ . . . . _ _ _ . . _ . . - - . 2 ... __.___._ _ __ : .a -_.___.- .-___ ._ -
(} t. TABLE 3-1 (Cont'd) SU!. MARY OF RESULTS OF AUGMENTED CLASS 2/3 PIPING FATIGUE EVALUATIONS FOR REPRESENTATIVE - f%RK I PIPING SYSTEMS
~
ARCHITECT ENGINEER SYSTEM FATIGUE USAGE utility (Size) NOC+DBA NOC+15A/SBA Plant E. Teledyne PASNY 1. SRV Discharge-C 0.189 0.117 Fitzpaitrick (10-inch)
- 2. SRV Discharge-B 0.022 0.027 (10-inch)
,- 3. SRV Discharge-A 0.252 0.303 ,
(._ .
. (10-inch) .
Boston Edison 1. HPCI Turbine 0.000 0.000 Pilgrim Exhaust (24-inch) _
- 2. RCIC Pump 0.000 0.000 Suction (6-inch)
- 3. Core Spray Suction 0.000 0.000 (18-inch)
Northeast Utilities . 1..H, eat Exchanger 0.053 0.003 Millstone 1 x 210A (8-inch)
~
- 2. SRV MS-8F (10-inch) 0.027 0.034 e
Q . , ..- . ... ; - - .;., .
. :. -..a .a..-:,' . _ :. .~.. .-, '.. .V
..~~ .~::.~r:::: . .:. ..... . .- .. . ----.::.- .- -....- =-:'..
( e. TABLE 3-1(Cont'd)
SUMMARY
OF RESULTS OF AUGMENTED CLASS 2/3 PIPING FATIGUE EVALUATIONS FOR REPRESENTATIVE MARK I PIPING SYSTEM 5 ARCHITECT ENGINEER SYSTEM FATIGUE USAGE Utility (Size) NOC+DBA NOC+IBA/SBA Plant Yankee AtGir.ic 1. Core Spray - Part 6 0.002 0.002 Vermont Yankee (10-inch) Niagara Mohawk 1. Containment Spray .012 .000 Nine Mile Point X8-326(3-inch) -
- 2. Pump .'II .001 .036 X5-337f12-inch)
- i. F. United Eneineers T'
Carolina Power 1. RCIC Turbine 0.086 0.340 Brunswick Exhaust (8-inch) l , 1
- 2. RCIC Barom. 0.001 0.003 Condenser (2-inch) l l 3. SRV Discharge 0.486 0.475 l (12-inch) ,
! i . 0 4 i
- 3-6 p
s Y. . .. , p . . . . . . . , ~ . ..-- ---. - >~ ~ -.. ---=-..--- -- -
?
-~-.--
i i TABLE 3-1 (Cont'd)
SUMMARY
OF RESULTS OF AUGMENTED CLASS 2/3 PIPING FATIGUE EVALUATIONS FOR REFRESENTATIVE MARK I PIPING SYSTEM 5 ARCHITECT ENGINEER SYSTEM FATIGUE USAGE Utility (Size) NOC+DBA NGC+IBA/SBA Plant G. NUTECH
' Comonwealth Edison - HPCI Turbine Exhaust 0.047 0.056 Quad Cities Unit 1 (24-inch)
Iowa Electric Core Spray Suction 0.059 0.039 Duane Arnold (12-inch) Detroit Edison SRV Discharge Piping 0.099 0.056 Fermi II (12-Inch) . . Northern States SRV Discharge Piping 0.307 0.197 Monticello . (10-inch) O H. TVA Browns Ferry RCIC Turbine Exhaust .010 .021 (2-fnch) RCIC Turbine Exhaust 0.003 0.095 (12-inch) ECCS Suction Line .007 .023 (16-inch)
.(a) SRV - Line H Elbow 0.003 .232 '
-(10-inch)
(b) SRV - Line H .047 .318 Ramshead (10-inch) i
- p. .,
3-7 , w ~ ,
= ;- G ::" gi:=C3~,w . ; ?. ~= ~ : ~~ T7*' '
. ~ ~:5- -
~~ '. ~ - ' -
.. . . . . . . . . . . - ..- - - - - -' - - - - - - - - - - ' ' - - ' ~ ~ ~ ~ * ~ ~ ~
p
4.0 REFERENCES
- 1. ASME Boiler and Pressure Vessel Code, Section III,1977 Edition with Addenda through Summer 1977.
- 2. NUREG-75/087, Standard Review Plan for the Review of Safety .
Analysis Reports for Nuclear Power Plants, May 1980.
- 3. NEDO-21888, Revision 2, Containment Prooram Lead Definition Report, General Electric, November 1981.
- 4. ORNL-TM-3520, " Comparisons of Test Data with Code Methods for Fatigue Evaluation"; E. C. Rodabaugh and S. E. Moore, Oak Ridge Nationa1' Laboratory, November 1971.
- 5. Markl, A.R.C., " Fatigue Tests of Piping Components,"
i ( Transactions ASME, Volume 74, pp. 287-303, 1952.
- 6. NEDO-24583-1, , Mark I Containment Procram Structural Acceotance l Criteria Plant-Unioue Aeolication Analysis Guide, General Electric, October 1979.
g . i 4-1 p. w; n y ,.
- . . .- _ _ _ . ._ . . ..~ _ - .... . . . . . .. ... _ ..
. e .
e e C. . APPENDIX A u I l AUGMENTED CLASS 2/3 FATIGUE EVALUATION METHOD, TABLES AND NOMENCLATURE - i e f e 0 e
- g =
4 e d' I O 1 i k i i o E 4 e b 6 e A-1
/~~
u- ;
- == T r ge .. **- . ...p . ., -_ ,
_., -_ _ , p ap=., , _ . . . _ ,
- , , . , .+y- - , - c--,w -
y .,--ip,., y- , y ,- - - - -%- - - ,,y w w, , , - , -
p r*' . I P 1.0 STEPS FOR AUGMENTED ASME CLASS 2/3 PIPING FATIGUE ANALYSIS In performing the augmented ASME Class 2/3 piping fatigue analysis first enter information on each piping system in the blanks shown in Table A-1. Then proceed with Steps A-G as follows: . i STEP A: Calculate stress resultants for each DBA, IBA/SBA and NOC occasional, thermal and anchor motion , loading condition. See Tables A-2 and A-3 for individual loads and Section 2.0 for nomenclature. - STEP B: For SRV piping, determine the discharge pressure (P3gy) and the stresses due to thermal expansion and anchor motions (MONOC) and discharge conditions (SRVIBA). STEP C: Determine number of SRV actuations that would be k U , expected to occur: (1) During normal and transient operation *over the
, life of the plant - nSRVNOC (2) During an IBA or SBA - n3gyggg i .
STEP D: Determine maximum characteristic frequency of the piping system for dynamic loadings (f,,x). A-2 (;: W.. .. .. ..<,.~.~.n.- . .. ~ - . . = . . ~
- - - - =
STEP E: Determine the location in the piping system with the most limiting intensified stress conditions. When piping systems have both stainless and carbon steel runs, limiting stresses in both runs should be considered in determining the most limiting 4 iocation for fatigue. STEP F: Perform the NOC - DBA fatigue evaluation by completing the information in Table A-2.
. STEP G: Perform the N00 + IBA/SBA fatigue evaluation by completing the information in Table A-3.
- g O
e e e I t
?
e
'~
A-3 (- . e
,ge* * ,, " .
* **% **.e**~
. *~ .. I " ."*P* 98* *' O* *** *
. . . - . . . - . . -...-.-......_.:=.. .: .-.......- .. . . . . . . . . :.
. . . . _ . . .- ._.. . . . . i V \
l
- k. 2.0 NOMENCLATURE CHUG = Flexural stresses due to chugging loading for j DBA/IBA/SBA as defined in Section 4.5 of the j LDR (Reference 3). Include stresses due to j underwater drag and fluid structure inter-action for submerged piping segments (ksi).
D = Nominal outside diameter of piping in inches
. at location where fatigue evaluatten is per-formed. ,
DW = Stress in ksi due to deadweight of piping system including fluid where appropriate. Corresponds to M Section NC-3652.g/I in Equation 3 of Section 11 of Code III, ASME (Reference 1). DBACO = Flexural stresses due to condensation loading as defined in Section 4.4 of the LDR (Refer-ence 3). Include stresses due to underwater ~ drag and fluid structure interaction for submerged piping segments (ksi). EQ(0) = Flexural stresses due to operating basis
] . earthquake (ksi).
- h. EQ(S) = Flexural stresses due to safe shutdown earth-A quake'(ksi). .
fmax = Highest characteristic or participating frequency of the piping system in Hertz. Use fmax = 30 Hz unless a lower value can be justified from analysis of the piping system. i = Stress intensification factor for ASME Class 2 piping analysis (Section NC-3673.2(b)). IBACO = Flexural stresses due to IBA/SBA condensation (PRECHG) oscillation loading or DBA prechug loading as defined in Section 4.4 and 4.5 of the LDR l (Reference 3) (ksi). i i I l
- . A-4 I
h G
. , - - - . . ~. n - -W -:....--.,.._....,.;.wsr.--
- -.. . - ..-: z...-s .--.--
- ~ , . . _ _ _ , . . . _ _ , _ _ _ . _ , , _ _ _ . . _ _ . _ __ - _ _ _
-s -
__ _ _ _ _ _ _ - . . -. . - o .co .-6. e odw$. . + cmo d 4 . .o eco.e- o. . ===c==..hA-. t K DBA, = Flexural thermal stresses due to thermal expansion of piping during the most limiting K IBA condition of DBA or IBA (whichever is ; applicable) plus the corresponding flexural stresses due to anchor motion (ksi). Tne KIBA stresses should also include the flexural stresses due to anchor point motion resulting from SRV actuation during IBA or SBA. KNOC = Flexural thermal' stresses due to thermal , expansion of SRV discharge piping and anchor motions during SRV actuation (ksi). nCY"*
= Number of significant response cycles for loadings included by SRV actuation. A value
- of 15 is used which is arbitrary since it is used in the calculation of Rggy; however, 15 is a reasonable estimate of the number of significant response cycles for SRY thrust loadings.
nk
= Effective full stress cycles for load combina.
tion k.
~
. N X
= Allowable cyc1es for total combination stress
-STE calculated as follows (Reference 5):
N K
= (245/S for carb6n steel, STE (ksi) TE)6 Ng = (2B1/S for stainless steel, STE (ksi) TE)5 n3gyy34 = Number of SRV actuations and reactuations that would occur during an IBA or SBA accident
- (whichever is greater).
n = Number of SRV,actuations that would occur SRVNOC under normal operating conditions over the remaining life of the plant. Pg = Pressure range inside the piping due to the most limiting pressure ccr.tition of DBA, IBA* or SBA, whichever is applicable (ksi). P 3gy
= Pressure range in SRY discr.arge piping due to SRV actuation. This load is only applicable to SRV piping' evaluation (ksi). -
k Y A-5 1 V.-., . .. ,. ., . - ,- _ . . . a ; p.. . ; ..:-
.~ ...-
. x- --
l l k .
\
Stress range in ksi due to naximum internal s PRDBA = pressure occurring during a DBA event (ksi). PRDBA = PA x D/4 t n PRIBA = Stress range in ksi due to maximum internal pressure occurring during an IBA or SBA event (k si ). PRIBA = PA x D/4 t n
=
Stress range in ksi due to internal pressure PRNOC resulting from SRV actuation (ksi). PRNOC = P3gy x D/4 t n R CHUG
= Equivalent maximum stress cycle factor for the CHUG loading. For recommended value, see Table A-4.
R DBACO
= Equivalent maximum stress cycle factor for DBACO loading. For recommended value, see Table A-4 .
R IBACO
= Equivalent maximum stress cycle factor for the .
IBACO (or PRECHUG) loading. The recommended value is 1.0 since the load 1.s essentially a (.; single frequency harmonic loading. R 3gy = Equivalent maximum stress cycle factor for the
- SRV loadings including thrust bubble drag and
! torus SRV excitation. For recommended values, see Table A-4.
S, i
= Intensified stress ranges in ksi calculated using the resultant moment' and section 5 , etc.
2 modulus as defined in Section III. ASME Code
] :
(Reference 1). SRVIBA = Flexural stresses due to actuation and reactu-
! ation of SRV system during an IBA or SSA -
; accident. Includes stresses due to thrust and SRV bubble drag where appropriate (ksi).
l' SRVNOC = Flexural stresses due to actuation of SRV
'; system during normal operating conditions.
Include stresses due to thrust and SRV bubble drag where appropriate (ksi) unless included in SRVTQF below. ' l i l
,- A-6
- g - _.,a ., .
.- ~- a '- G ~ '- ~
.v =.*. . : . .~ S__ "~
........-...l
. sr
/
t ! l SRVTOF = Flexural stresses due to bubble drag resulting , from actuation of SRys if not included with SRVNOC stresses.
= Total intensified stress range in ksi for STE combination based on Equation 11 of NC-3652.3 of Section III, ASME Code (Reference 1).
Calculated as follows: - STE
- 31*S2 * " ' " + ( 0. 75 i x DW) .
t n
= Nominal thickness of piping in inches at location where fatigue evaluation is per-formed.
4 e f h-L. 9 1 I s l I. I a e (s A-7 h ,.;._
.g.v .-- ;, .. .
<; ~; , , . ..{.; r ' ' - % H u? . . .h .
. -,,v.. - . ., , - . - - . - - -. ,, - w,-
TABLE A-1 - MARK I PROGRAM ATTACHED PIPING AUGMENTED ASME CLA55 2/3 FATIGUE EVALUATION
- 1. GENERAL INFORMATION A. PLANT (UNIT):
B. UTILITY: C. ARCH / ENGR: D. PREPARED BY: DATE: s II. PIPING SYSTEM INFORMATION A. SYSTEM IDENTIFICATION: B. NOMINAL PIPE SIZE: C. MATERIAL: ( - III. PIPING ANALYSIS INFORMATION
~
A. ANALYSIS METHOD: i B. INTENSIFICATION FACTOR (i): C. LOCATION OF MAXIMUM STRESS: _ (Describe briefly or attach figure) D. DEADEIGHT STRESS (DW): l E. NOMINAL THICKNESS'(t n ): F. OUTSIDE DIAMETER (D): i l A-8 I c-- .-.. ._ . . . _ . _. ,. . . - - _ _ ..
. . . .,. s . . .. l l
i
- - - - . , _ _ , - . - . _ - __..__.,__..,_.2_-_, . , ,,. __, ., ., _ . _ _ _ _ . _ , _ _ _ . _ . _ _ . _ _ _ _ . , ..
v
,- a l.... '- .
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ma -
=
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= J W mD
< (.D
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8 1 3 d
=
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. -= 8..
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=.. .. - ~ . . . - . . . := - n -
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'P TABLE A-4 CYCLE REDUCTION FACTORS FOR MARK I DYNAMIC LOADINGS
(
Reference:
A5ME Coce Section III, Suosection NC-3611.2) i Type of Load I - Significant Load Cycle Reduetten (Abbreviation) . Cycles (n) Factor (R) I .
- 1. Acceleration due to SRV 15 0.3 discharge (SRVNOC)
- 2. Thrust due to SRV discharge 15 0.3 (SRVNOC)
- 3. Bubble drag due to SRV 15 0.3 discharge (SRVNOC) 4 Acceleration due to DBA 900(Note 1) 3/fmax or 0.1 condensation oscillation (whicnever
(, (DBACO) is greater, g , Note 2)
- 5. Acceleration due to IBA 102 DBA 1.0 condensation oscillation 10,450 IBA/SBA (IBACD or PRECHG)
-. 6. Acceleration due to 321(Note 1) DBA 2/fmax or 0.1 chugging (CHUG) 9600(Note 1) IBA/SBA (whicnever is greater, Note 2) l NOTES: .
- l. Based on a 30 Hertz maximum participating frequency (actual f max could be used)
- 2. f max = highest participating f'requency of piping system. *
, A-11 p, .
(.
. "J '., . , -Q ,, J. - ~~"'*s.. " s N '.". 7 .." ~.". # ~ *
- eQ (D,' , .,y~
,~7^,'
~
i . i . j . ek HUMBER OF CYCLES 1 (n) i 1 . 10 102 103 - 104 105 10 I 3 , . 10 - ' i. d ; 1 2 T t i VAI.UES '
;j or *% %
S, t ,,,,,, (MSI) '10 *
^ ' -
.f-
"" ~ - '
STAIHLESS STEEL i Y 4 i :- g ~-.,,,,~~,,, sa 231 x n-0.2 ' l l . CARDONSTEELj * ,' , l g
.I s=- 245 x n-0.2 ~ ~ ~ , , , , ,,~ i
,- i
, v
,.1
~~~ ~~ ~- .
t
- .~, l j, 10- I 4
t . , e . ( c '.' f FATIGUE CURVE i , FOR AUGMENTED CLASS 2/3 FATIGUE EVALUATION i i i l i f i d FIGUllE A-I
.+ i E
1 T
= . . .
I APPENDIX B RESPONSES TO NRC QUESTIONS AND COMMENTS STAFF PRESENTATION OF SEPTEM3ER 10, 1982 I i
?
.6
. e h
9 e 8 3 r B-1 - N. g' t t M -
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- s-C'
} .
(.
.( -
} This A;pendix provides additional information and justifictaion for several of the assumptions employed in developing the Augmented Class 2/3 fatigue evaluation methodology. This material is provided in response to comments and questions resulting from a presentation to the NRC staff on September 10, 1982. The NRC staff comment or question is listed first followed by the corre-spending response.
- 1. Comment:
. Provide documentation of the fatigue methodology for the NRC staff.
Response
Section 2.0 of the body of this report and Appendix A contain i the requested documentation.
\-- 2. Comment:-
Identify which piping systems were evaluated for each plant considered and the fatigue usage results for each. . I Response: l l Table 3-1 in the body of this report contains the requested information in tabular form. -
- 3. Comment:
Document how prior fatigue usage has been considered in the l fatigue evaluation methodology.
- p. B-2
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. . _ . . ... .m . . _ .:._ . . _._ . . . . _ . _ . _ _ _ .
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1 . ., i,, Resoonse: ( The number of SRV actuations used in the analyses was based on the expected number of SRV actuations over a 40-year plant lifetime. Thus, the reruits reported in Table 3-1 of the body of this report account for prior fatigue usage. In many cases the support arrangements for the SRV discharge , I piping have been upgraded to withstand all of the Mark I Containment Program loadings. The stress distributions of original piping arrangements are not generally available and in any event would not be comparable to the stress distributions in the upgraded piping arrangenents. Thus, it was concluded that a reasonable approach would be to I extrapolate the fatigue evaluation results for the upgraded piping configuration for the full 40-year life of the plants. 4 Comment: 2
'e
- Provide a rationale to justify not including thermal gradient stresses in the fatigue evalu'ation methodology, t
Resoense: I
-l ASME Class 2/3 piping design equations do not include
, stresses due to thermal gradient stresses since they are ,
; generally not significant to the design of these systems. To justify the assumption for the fatigue evaluation method-ology, calculations were performed of the fatigue usage resulting from the thermal gradient stresses. Typical values .
! for key parameters were taken as follows:
B-3 6'
'k ,
Y*-
-f * ' ... - .: e.svg .a ~~2. w : --,,,;--- ' ~'
. . v -= , - ,n.t.;, --l.
- - ~ . . . . .:-:: - . . - . - . _ _
. .=.z =....;
l '
?%.
,l t
(;~ O Temperature of steam inside the SRV discharge piping in the wetwell - 3500F. O Initial temperature of torus water adjacent to SRV discharge piping in the wetwell - 700F. For these conditions and typical sizes of piping used in Mark I plant SRV discharge. piping, the peak thermal gradient stress ranges were calculated as follows for th'e two types of materials used for this piping:
O Carbon Steel - 3,900 psi (compressive)
O Stainless Steel - 15,500 psi (compressive) These stresses are compressive and occur on the inside of the pipe wall. Stresses which occur on the outside of the pipe , are tensile and are somewhat lower in magnitude. The calculated stresses for the stainless steel piping are larger due to the lower thermal conductivity of stainless steel {' - compared to carbon steel. The effect of the maximum. stresses on f atigue usage was estimated based on a bounding number of SRV actuations of 3,100(500-1000 is more typical) occurring l over a plant lifetime. For the two materials, the fatigue usage was calculated using the methods of Appendix A with the result as follows: i . l I O Carbon Steel - less than .001 i 0 j Stainless Steel - 'less than .002 l Since all fatigue usages due to stresses other than thermal I gradient stresses were calculated for all Mark I plants to be a less than 0.5, the effect of.the usage due to thermal I_ gradient stresses on fatigue lifetime would be insignificant. B-4
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. RAI 14 Provide justification for the validity of analyzing SRV piping in two separate lengths (above and below the vent header) as indicated in Section 2. 3.1.1 of the PUA report (4).
RESPONSE
Penetration thru vent header is considered as an anchor point for each piping analytical model (for description of analytical models see response RAI 3a). The penetration is reinforced with plate and structural members which makes it very rigid in comparison with piping downstream and upstream and as such justifies assumption for analytical anchor. 0 O
(. RAI 15 With reference to Section 3.8.4.2 of the PUA report (4), provide justification for using the SRSS method for
- combining responses due to SRV and chugging loads, specifically indicating whether the stress intensity value corresponds to 84% probability of non-exceedance as determined from the cumulative distribution function.
RESPONSE
i Justification for using the SRSS method to combine responses due to SRV and chugging loads is provided in "SRSS Response Combination of Multiple Dynamic Responses" by R. P. Kennedy, Structural Mechanics Association, dated August 1982 l . (September 9, 1982 NRC Presentatioin Document). l l ,~ (. . l l { i 1 k.
FRe
- r. ..
R Ar l5 ATTACHM&v'f* 75 miLc1 ii~ (FAC) SRSS RESPONSE COMBINATION OF MULTIPLE DYNAMIC RESPONSES ~ MARK I by . R. P. Kennedy t . I August, 1982 g f STRUCTURAL mECHAnKS b ' wsm-mmmma R SS OCI.A . ..., TE
.... S 5160 Birch Street, Newport Beach. CaRf. 92660 (714) 833 7552 -
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PARAMETRIC STUDIES TO' JUSTIFY SRSS RESPONSE .
. COMBINATION FOR MARK I SHOULD NOT BE REQUIRED
- 1) LOADINGS AND RESPONSES SIMILAR TO THOSE FOR MARK II AND MARK III
/
/
- 2) GENERIC PARAMETRIC STUDIES HAVE BEEN PERFORMED FOR MARK II AND MARK III TO JUSTIFY SRSS.
RESPONSE COMBINATION FOR THESE LOADINGS C ~
- 3) SUCH STUDIES ARE COSTLY AND UNNECESSARY TO REPEAT FOR MARK I BECAUSE OF THE AVAILABLE MARK II AND III RESULTS
- 4) LIMITED MARK I STUDY SHOWS RESULTS SIMILAR l TO MARK II AND MARK III l
C 1 ( 'N'
- :_:~ T:__ :~ - ~ ~~:~;: T' : . T ~~~ _ _ ~_-_.
( (N. M. NEWMARK AND R. P. KENNEDY, NED0-24010-2, DECEMBER, 1978) BASIC ASSUMPTION BEHIND CRITERIA FOR SRSS COMBINATION ~OF" RESPONSES
- MANY SOURCES OF CONSERVATISM EXIST IN DESIGN AND EVALUATION PROCESS.
- ADDITIONAL CONSERVATISM DOES NOT HAVE TO BE INCORPOR-ATED WITHIN THE RESPONSE COMBINATION PROCESS.
- IT IS NOT NECESSARY FOR THE COMB'INED RESPONSE TO HAVE A LOWER PROBABILITY OF EXCEEDANCE THAN THE INDIVIDUAL RESPONSES.
( e AN 84% NEP FOR COMBINED RESPONSE ASSUMING CONCURRENT APPLICATION OF LOADINGS IS SUFFICIENT l O g:
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MULTIPLE RFEPONSE TIME HISTORIES ,
- 1. TIME HISTORIES HAVE RANDOM RELATIVE START TIMES.
(UNCORRELATED)
- 2. TIME HISTORIES ALSO HAVE RANDOM AMPLITUDES.
- 3. DESIGN AMPLITUDES ARE DEFINED TO BE AT THE 84% NON-EXCEEDANCE PROBABILITY BY CRITERIA.
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- 4. HOW SHOULD PEAK INDIVIDUAL RESPONSE BE COMBINED 7 I
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c . F, t CRITERION 2 R = SRSS8 q SRSS COMBINED RESPONSE WHERE EACH INDIVIDUA1. RESP 0.NSE HAS BEEN DEFINED CONSERVATIVELY AT . 84TH PERCENTILE OR 1.15 TIME MEDIAN, WHICH-EVER GREATER. R = - RAND 0M TIME PHASE COMBINED RESPONSE WHERE TM ' ALL AMPLITUDES DEFINED AT 84TH PERCENTILE. R =
, COMBINED RESPONSE CONSIDERING BOTH RANDOM AMPLITUDE AND TIME PHASING.
GOAL OF SRSS COMBINATION N. RsR P SRSSg a 84% (D CRITERION 2 REQUIREMENT P RTg4 5RSRSS84
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3 I' (DR.1100, 'MECilANICAL BRANCil,' NOVEMBER,1980) i, !- l: Brief Summary of . NRC POSITION ON SRNS . 4
- i -
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- ShaF no+ be apelled universally wHhou4
!' JusMicaHon. - I
- - ResuHs of Parametric 'sludies (NuREG/ce-caso)
L! . _JusuificaHon shall base on case invesH9aHon v L! ,Q of response fun'c4 ion s ~ ( nof Ioad -func+ Ions), : e sess may be used -for combining responses la Loch + SSE w'iihin RCP& and ils Supyords, and V i oil ASME Class I, 2, 3 piping Systems. l
'e SRss may be used +or combining more 4han 2 l li -
responses i
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( BASED ON CONSERVATIVELY BIASED RESULTS
- ALL 167 CASES MEET INTENT OF NEWMARK-KENNEDY CRITERION.2
- SRSS METHOD CONSERVATIVE IN MOST CASES
(
- SRSS METHOD IS A REALISTIC APPROACH FOR CASES WHERE HIGH FREQUENCY RESPONSE IS PREDOMINANT t
!
- SRSS METHOD IS IN NO CASE UNCONSERVATIVE
- SRSS METHOD-IS GENERICALLY APPLICABl.E TO BOTH ASME-AND NON-ASME COMPONENTS FOR LOADINGS CONSIDERED 1
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(;. L. MARK II PIPING AND COMP 0NENTS
~
e GENERIC STUDY OF 291 RESPONSE COMBINATION CASES (NEDE-24010-P, JULY 1977) (NED0-24010-2, DEC. 1978) e ALL 291 CASES MEET INTENT OF NEWMARK-KENNEDY CRITERION 2 s.
- SRSS METHOD IS GENERICALLY APPLICABLE FOR LOADINGS CONSIDERED e
/* ~ bor
- 2,2. -
C. . . . . . , . . , . . . . . - - . . . .
. . ~ . - - - - -
, . . _ _ _ _ ___ _ _ _ _ . _..,_ . - _ _ _ _ , - - _ ~ _ _ - , - _
. = . .. -- .. _ .. -
... ~
( BROOKHAVEN CONTAINMENT AND DRYWELL STRUCTURE GENERIC STUDIES (NUREG/CR-2039, JUNE 1982) (NUREG/CR-1890, JUNE 1982) NO. OF CASES SRSS 1.2* SRSS MARK II 400 87.8 98.8 3 OR MORE LOADS MRK II 400 80.7 97.5 7-2 LOADS MRK III 3904 75.0 97.0 9 IF INDIVIDUAL LOADS ARE CONSERVATIVELY DEFINED (84% NEP OR 1.15 TIMES MEDIAN, WHICHEVER GREATER) THEN SRSS RESPONSE COMBINATION ACCEPTABLE FOR STRUCTURES EXCEPT WHEN ABS /SRSS <1.25 IN WHICH CASE 1.1* SRSS OR ABS (WHICHEVER LESS) IS PREFERRED. l t 1 =l
- C -
__ty-
. - _,.,e,._, -- _ . . _ . _ .. _,.., . .__-. , . . , _ _ . , . , , , . _ . _ , , , , _ _ _ , _ , - . . . . _ _ , _ _ _ _
RECOMMENDATION FOR MARK I' e SRSS RESPONSE COMBINATION BE ACCEPTED FOR FOLLOWING: o e STRUCTURES & COMPONENTS TORUS, TORUS SUPPORTS, ATTACHED PIPING,
& EQUIPMENT
*
- ALL COMBINATIONS OF RESPONSES INVOLVING FOLi.0 WING INDEPENDENT LOADINGS WITH RANDOM RELATIVE START TIMES x
OBE, SSE, SRV, CHUG, CO, POOL SWELL e SRSS RESPONSE COMBINATION HAS NOT BEEN STUDIED FOR MULTIPLE RESPONSES FROM SAME LOADING WITH LIKELY CORRELATION BETWEEN PEAK RESPONSES
*
- EXAMPLE: DIRECT FLUID DRAG AND SUPPORT SHAKING EF7ECTS ON SUBMERGED PIPING FROM SRV I
I C - _15 -
-- - = - - . -. . . . . . _ _ . . . . _ _ , _ _ _ _ , _ ,
- O $ UNITED STATES 3 y f".f. i; ,
NUCt. EAR REGULATORY COMMISSION
' wAp4NGTON. D. C. 20555 s
2, x ~(; s (. h, .....f - MAR 10123 Mr. H. C. Pfefferlen Manager
. 5WR Licensing Programs General Electric Company 175 Curtner Avenue San Jose, California 95125
Dear Mr. Pfefferlen:
SUBJECT:
ACCEPTABILITY OF SRSS METHOD FOR COMBINING DYNAMIC RESPONSES IN MARK I PIPING SYSTEMS We have evaluated General Electric Report NEDE-24632, " Mark I Containment Program Cumulative Distributten Functions for Typical Dynamic Responses of Mark I Torus and Attached Piping Systems" that was transmitted by letter from you dated October 7,1982 and the information provided by R. P. Kennedy of Structural Mechanics Associates during a September 9, 1982 meeting with the staff. Based on our evaluation we have concluded that the use of the SRSS method for combining peak responses of Mark I piping and supports under dynamic loads is acceptable. A copy of our safety evaluation is enclosed. Sincerely, Domenic B. Vassallo, Chief Operating Reactors Branch #2 Division of Licensing
Enclosure:
As stated - G
.-.: .. . u .. . . - . . . _ . - . . - . . - . . . . . . . .
SAFETY EVALUATION BY THE OFFICE OF NUCLEAR REACTOR REGULATION r FOR ACCEPTABILITY OF THE SRSS METHOD FOR COMBINING
\.
DYNAMIC RESPONSES IN MARK I PIPING SYSTEMS Intorduction The information provided in References 1 and 2 by the Mark I Owners Group is intended for justifying the use of the Square-Root-of-the-Sum-of-Squares (SRSS) Method for combining responses of torus structures and attached piping systems under various dynamic loads. The following is care evaluation of
- SRSS acceptability for Mark I piping and supports. The portion of NEDE-24632 concerning torus structures is beyond the scope of this evaluation .'
and hence is not addressed. Information Sunnary The provided information consists of 111 actual cases of response combinations, of which 72 cases were selected from the following three typical Mark I piping systems: (1) RCIC turbine exhaust piping (2) HPCI turbine exhaust piping, and (3) vacuum relief piping systems. Response time functions consist of stress resultants, modal displacements, and support reactions, which were induced by suppression pool swelling and chugging from safety relief valve , discharges. Components in horizontal and vertical directions were considered separately. In each case of responsa combination, the non-excedance probability (NEP) of SRSS was determined by generating a cumulative distribution function (CDF) of the response peak resultant from combining the response time functions with random. time lags. For 72 piping cases investigated, the average NEP is ( 87% with a standard deviation of 6%, which is comparable with the results of previous Mark II piping SRSS investigates. Evaluation The referenced report with respect to the SRSS method for combining dynamic responses in Mark I piping systems is acceptable for the following reasons: i, (1) The response functions for the Mark I design are very similar to those previously reviewed in great detail for Mark II and Mark III' designed plants (Refs. 4 & 5), (2) although somewhat limited in number, for each of the 72 cases of piping and support location response combinations investigated the NEP levels are greater than those prescribed in NUREG-0484 Rev.1 (Ref. 3), (3) the presented NEP levels 'are compatible and consistent with those in the Mark 11 and III SRSS evaluations, and (4) although seismic loads were t included in the response calculation, the combination of sets dc load rv sponse with other hydrodynamic load responses has been adequately represented in the Mark II and III SRSS evaluation and need not be separately conducted. Conclusion Based on our evaluation, we concluded that the use of the SRSS method for combining peak responses of Mark I piping and supports under dynamic loads is acceptable. Dated: MAR 101983 ( Principal Contributor: S. Hou, DE/MEB -
.f*
'i; REFERENCES
- 1. Topical NEDE-24632, " Mark I Containment Program Cumulative Distribution Functions for Typical Dynamic Responses of Mark I Torus and Attached Piping Systems", Prepared by Nutech for GE 12/80,
- 2. Presentation "SRSS Response Combination of Miltiple Dynamic Responses",
by R. P. Kennedy of Structural Mechanics Associates, 8/82.
- 3. NUREG/0484, Rev.1 " Methodology for Combining Dynamic Response", 5/80.
- 4. Memo, J. P. Knight to R. L. Tedesco, Evaluation of GE Topical NEDE-24010 and Supplement 1, 2, 3, " Technical Basis for the use of SRSS Method fer Combining Dynamic Loads for Mark II Plants," 5/28/80.
- 5. Memo, S. Hou to J. Stefano, "MEB Resolutions to LRG II Issues", including Acceptability of SRSS for Mark III Mechnical Components, 4/8/82.
\
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RAI 16 Indicate whether the fatigue usage factors for the SRV piping and the Torus attached piping are sufficiently small so that a plant-unique fatigue analysis is not warranted for piping. The NRC is expected to review the conclusions of a generic presentation (6) and determine whether it is sufficient for each plant-unique analysis to establish that the expected usage factors for piping are small enough and do not warrant a plant-unique fatigue analysis of the piping.
RESPONSE
Plant specific usage factors for the following lines are presented in Table 3-1 (Ref.1) for the Brunswick Plant: (1) RCIC Turbine Exhaust Line (2) RCIC Barometric Condenser Line (3) SRV Discharge Line Usage factors tabulated are less than 0.5. Table 3-1 also presents usage factors for other plants as determined by that plant's Architect / Engineer. A review of the table indicates a reresentative sampling of Torus Attached pipe lines and SRV discharge lines has been included. Typically, the usage factors associated with the SRV lines are higher , than those associated with torus attached pipe lines. Since usage factors are less than 0.5, a plant unique fatigue analysis is not warranted. REFERENCES (1) " Mark I Containment Program Augmented Class 2/3 Fatigue Evaluation Method and Results for Typical Torus Attached and SRV Piping Systems"; MFR-751, November,1982; MPR Associates, Washington, D.C. 1 f
..-. .:..- - .w _ .:=...-- ..... .-._.-..;. _ . . . - . _ . - - . - - . .;.
r ( - RAI 17 The ASME Code provides an acceptable procedure for computing fatigue usage when a member is subject to cyclic loadings of random occurrence, such as might be generated by excitations from more than one type of event (SSE and SRV discharge, for example). This procedure requires correction be made to the stress-range amplitudes considered and to the associated number of cycles, in order to account for the interspersion of stress cycles of unlike character. State whether or not the reported
, usages refleet use of this method. If not, indicate the effect on reported results.
RESPONSE
The ASME procedure for superposition of cycles of various origins was not used to compute usage factors for the vent system.
~
Stresses from seismic loads are very low and were considered to be negligible. SRV discharge and LOCA loads are the, only sig-nificant cycle loads which can occur simultaneously. Of the . 400 postulated SRV actuations, 350 occur during normal operating conditions. Therefore, the simultaneous occurrence of cyclic loads not explicitly addressed in the " Plant Unique Analysis" is limited to 50 SRV actuations (of a specific valve) concurrent with a LOCA. Justification for this is based on the conservative assumptions used in the fatigue analysis. Therefore, the overall ef fect results in calculated usage factors that are conservative.
. The foll wing conservative assumptions were made for the fatigue analysis of the vent system:
- 1. The assumption of'50 actuations of a particular safety relief valve during a LOCA is conservative.
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RAI 17 (Continued) RESPONSE (Continued)
- 2. For the analysis, 5 equivalent peak stress cycles were
, used for each SRV actuation. For a typical SRV time history applicable to the vent system (e.g. Figure 3.8.1-3 of the PUAR), the number of equivalent peak stress cycles per SRV actuation is 3.5.
- 3. CO fatigue was calculated conservatively because each of the three harmonics was considered to individually cause the peak stress. However, the peak stress is actually caused by the summation of the three harmonics.
.The total combined effect of all of the above is that the calculated usage factors are conservative.
e f: v
- a. . =. ::. . -. : - : . . - -.-.-.2.-...--.. . - - . -.- - . . . .- - . . . -...-...--..-.-.. - --.
r b RAI 18 During pool swell and seismic events, local concentrated loads are applied to the suppression chamber at eight vent locations, 45* apart. For example, on page 1-31 (4), it is stated that the vent header load during pool swell is 342 kips. In the overall analysis of the suppression chamber, these discrete loads are represented as a uniform line load (of equal total magnitude) instead of invoking the Wilson 2 capability to expand the load in Fourier series with approximating load peaks at eight equally spaced stations. Develop the justification of your procedure in greater detail.
RESPONSE
There are eight columns (4 existing and 4 proposed new columns) at each vent header intersection that apply a total concentrated load of 684 kips at the suppression chamber. The total number of columns at all the vent headers is 64. The meai mum concentrated load in each column is 102 kips. These 64 point loads applied to the suppression chamber have two effects, local and overall on the structure. The overall effect of the concentrated loads is to induce stress in the critical sections of the torus such as Section 1 and 5 shown in Figure 1.8-1 of PUAR. The local effect of each concen-trated load is to produce punching shear in the torus wall. The allowable shear stresses,however, depend upon the meridional and hoop stresses around the assumed failure surfaces. Therefore, an overall analysis of the torus due to the concentrated loads is required to determine these overall stresses at the critical section. To perform the overall analysis of the torus for eight equally spaced sets of eight concentrated loads along the circumference i
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l RESPONSE (Continued) of the torus, it was assumed that the maximum possible load per unit length along circumference, obtained from maximum concentrated load of each pair of column divided by the minimum distance between two concentrated loads, is applied along the circumference of the torus instead of eight discrete locations. This assumption gives higher stresses in Sections 1 and 5 far from the point of application of the concentrated loads. It also gives higher meridional stresses at the vicinity of the applied loads which results in lower allowable punching shear stresses and therefore, are conse rvative. - The punching shear stresses due to the concentrated loads are then compared with the calculated allowable punching shear stresses. The factor of safety for local punching shear is around 5. i
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f* [ RAI 19a Provide justificatica for neglecting the discontinuities in . the overall structure of the suppression chamber (a 16-sided polygonal ring):
RESPONSE
. The suppression chamber is a 16-sided polygonal ring with a circular inner boundary at its attachment to the circular drywell. All the top bars, seismic bars and global hoop bars are curved forming concentric circles with centers on the drywell axis. The local hoop bars are circles and together with the global hoop bars form a torus.
However, the wetwell lining is composed of 16 right angle cylinders with tapered ends. ~
- f. The analysis for all loading conditions which include internal pressure (60 cases out of 65 load ' eases, with controlling cas'es among them) is performed assuming concrete is totally cracked in the local and global hoop directions. The reinforced concrete suppression chamber is thus, in fact, a toroidal structure with the cylindrical liner. The liner is assumed as a . torus in the analysis.
The joints between adjacent' straight segments of the liner have an enclosed angle of 157.50 Local stress concentratioin may occur at any discontinuity, but according to ASME, Section III, Division 2, since the liner is anchored to the concrete at relatively close intervals (about 2" at joints and 11" at oder places) its integrity - shall be examined by checking the liner strain under service and factored loads. The liner anchor analysis was performed and r.- strains are reported in Section 1.11 of PUAR. V
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RAI 19b Also, provide justification for the implicit assumption that stress distributions and magnitudes so obtained conservatively represent those of the actual structure. In particular, address the situation at joints between adjacent straight segments.
RESPONSE
See response to RAI 19a. e 4
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i (- , I l RAI 20 Section 1.5 states that the thermal effects were considered, i assuming a linear temperature gradient between the liner and the exterior surface of the concrete. Although a linear temperature distribution is substantially correct for steady state conditions in the wall structures with concentric inner and outer boundaries, it is .not theoretically correct for the case of a thick section with a circular inner boundary and
. approximately square outer one. Discuss the adequacy of this assumption to predict the onset of thermal cracking in the concrete.
RESPONSE
The inside temperature of 1400F and 1680F for SBA and IBA, respectively, is occurring with the high internal pressure of LOCA conditions for 61 load cases given in load combinatiion Table 1.4-1 of PUAR. The torus is assumed to be totally (, cracked in local hoop and global hoop (S and T) directions due to high internal pressure. Therefore, only the liner and the reinforcing bars, which are approximately 26" apart (Rliner =200", Reinforcing = 174"), resist the thermal loads. At the iniciation of the LOCA condition the temperature distribution across the torus wall is nonlinear. The liner has a high temperature reducing sharply across the wall. This temperature distribution has the same effect as an internal pressure that produces tension in al-most the entire section of the concrete. This phenomena is not different from the cracking assumption made in the analysis. Using the totally cracked . concrete for each of the 61 load combinations, which includes the internal pressure, is a reasonable assumption. Therefore, the assumption of linear or nonlinear h
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i -. RAI 20 (Continued) kESPONSE (Continued) distribution of the thermal load is immaterial in predicting the
. onset of thermal cracking in the concrete. As the LOCA condition progresses in time, the steady state condition prevails and the assumption that there is a linear distribution of thermal load become more valid in predicting the total forces and moments .
developed in the concrete cross scetions. For the four load cases where no internal pressure exists (cases 2, 3, 64 and 65), the concrete is assumed to be partially cracked. The thermal load for these cases is the normal
\
operating temperature which varies between 760F and 920F with changes occurring over a long period of time. Steady state temperature (1000F inside to normal temperature on outside of the torus) is realistic for these cases. The assumption that the torus is totally cracked in the S and T directions due to high internal pressure eliminates the need for a vigorous cracking pattern investigation which is an iterative process, since the thermal load in the torus is not high. The membrane stresses in the liner from the load combination are all below yield. ASME Sectiori III, Division 2, permits membrane liner strain under factored load to reach between 3 and 5 times the yield strain. p
-- -- ,,-w< ~ - - - -
RAI 21 In Section A.l.1, a normal mode and frequency response analysis for CO loading, made to assess effects of fluid / structure interaction, is described. Provide the total number of modes used in the summation, the total percentage of modal effective mass thus included, and discuss provisions which may be incorporated in the computer program to account for effects of the unsummed portion of the modal effective mass.
RESPONSE
i NASTRAN Rigid Format 26, used to study the fluid-structure interaction per Section A.l.5 of PUAR, is a direct method of frequency response analysis in which the entire differential equations of motion are transformed to the frequency domain. The resulting equations are solved for all frequency ranges of . the input load. The condensation oscillation load is defined
. in terms of frequency and amplitude from 1 to 50 Hz. in w
the Load Definition Report (2). The modal formulation of the frequency response analysis is not used. l ( l (1 v-r
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f+ s.. RAI 22a Provide a summary of the results for the submerged structures listed in Section 2.1.1. (b), page 2-1 and contained in the calculation books listed in Appendix A-2 (B & E), pages 2-21 and 2-11 (4).
- RESPONSE The subject submerged structures were evaluated for various load case combinations and.the resultant piping stresses were found to be acceptable in accordance with applicable stress limits. The moments of selective load cases were com-bined absolutely using the Stardyne " Post" program. The moments were evaluated against applicable piping stress levels in accordance with the following:
(1) ASME Boiler & Pressure Vessel Code Section III - Division 1; Summer 1977 Add ~enda
.(2) Mark I Containment Program Load Definition Report (NEDO-21888, Class I, December 1978, Rev. 0) by General Electric Nuclear Energy Engineering Division (3) Mark I Containment Program Structural Acceptance Criteria, Plant Unique .
Analysis Application ' Guide, Task 3.1.3 (NEDO-24583-1, 79 NED 125, Class I, October 1979) by General Electric Nuclear Energy Engineering Division (-
0 (- , s > RAI 22b Also include the results of the strainer analysis performed l by the supplier, as referenced in Section 2.3.1.2 (4).
RESPONSE
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7,.- ZURN INDUSTRIES INC. DESIGN REPORT 1281A (Total 15 pages) s. s. CAROLINA POWER & LIGHT COMPANY BRUNSWICK STEAM ELECTRIC PLANT UNITS 1 & 2
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l 1 1 Attachment RAI 22b
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DESIGN REPORT 1281A STRESS ANALYSIS CALCULATION .;
- /
. OF SUPPRESSION POOL STRAINERS FOR CUS'ICMER: United Engineers & Constructors, Inc.
UTILIT.': Carolina Power & Light Co. Brunswick Steam Electric Station
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DESIGN SPECIFICATION: Spec. No. 9527.001-236-21, Rev. 2 PURCHASE ORDER NO: 9527.001-236-21, CP542 MANUFACTURED BY: ZURN INDUSTRIES, INC. Fluid Handling Division
. Order No's: 82-N-4965
-4966
-4967
-4968 I
ORIG. 1/27/82 PREPARED BY: Ira Schnall 12-21-81 CHECKED BY: F_FM.bEr4Mred I - /5- 8 2. . APPROVED BY: IN 2 (.
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ZURN INDUSTRIES, INC. FLUID HANDLING DIV,901 WEST 12TH ST, ERIE PA, USA 16501 PHONE:814/455-0921
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RAI 23 Section 3.8.4.3, page 3-47 and Item 4 of Table 3.4.1.1 (4), second page, indicate that the limits on the local membrane stress intensity and on primary - membrane plus - primary - bending stress intensity were modified by using 1.3 S ac in place of S ac and certain restrictions are mentioned in a note
, related to this section. Confirm whether these limitations have been applied in the analysis.
RESPONSE
The limitations given in the " Plant Unique Analysis Application j Guide" (GE document NEDO-24583-1), Table 5-1 and repeated in Table 3.4.1-1 of the PUAR on the use of 1.3 S ee in place of S ac were used in the analysis. s._ e 9 6 l l l l
.i
e RAI 24 The stiffness matrix for the downcomer/ vent header intersection given in Table 3.6.1.1-1 (4) and used for the beam model analysis (Section 3.6.2) shows the diagonal terms only, i.e., no-off-diagonal terms indicated. Justify and state tha assumptions for not including other significant off-diagonal terms.
RESPONSE
In the Mark I Containment long term program, it was determined that for all Mark I plants, the downcomers are at least 50 times stiffer than their junctions. These results are summarized in Table 4-1
. page 4-21 of Reference 1. For Brunswick Units 1 and 2, the ratio of downcomer to junction stiffness is 82. Due to symmetry in the header geometry at this location and due to the downcomer being so rigid with respect to the flexible downcomer/ vent header junction, six independent degrees of freedom with no significant coupling can be assigned to the center of intersection. Thus, the down-comer responds as a single degree of freedom structure in each of the six degrees of freedom. This was also shown by an analytical study described on page 4-3 of Reference 1.
In the modified intersectio'n, changes in intersectioin stiffness were accounted for and again, due to the geometry, it was roted that still no significant coupling occurred in each of the six degrees of freedom. Because of the validity of the six independent degrees of freedom assumption, the. contribution of the off-diagonal terms to the junction stiffness was again considered minimal. / Q,
.. ... , . u.--
. 1 i.
RAI 24 (Continued)
Reference:
- 1. Mark I Containment Long Term Program Development of Downcomer Laterial Loads From Full Scale Test Facility Data - Task Number 7.3.2 - NEDE 24537-P, May 1979 O
t . . e e ,y .c n-- -mv. ~ = - < - -- ,, 4,,, . , , ,<-r-o.-- --,,a-.- .--- - , ~=w. e--- -------u., * - - e--, ..,. .
g7 ( RAI 25a The downcomer/ vent header finite element model shown on Figure 3.6.1.1-12 (4) indicates that large elements with non-conventional aspect ratio were used at some regions of discontinuities, such as the stiffening plate regions attached to the downcomer and to the vent header. Indicate
.the calculated stress intensity values at such locations.
RESPONSE
The STARDYNE QUAD 4 element is well-behaved for aspect ratios up to 10. The elements in question in the downcomer model have an aspect ratio of 8.16. The model provides an accurate representation of the primary membrane plus primary and secondary bending stresses in the downcomer shell in the local region immediate to the discontinuity between downcomer shell and stiffener. A stress concentration factor of 2.0 was used to account for the peak
-stress intensity at this local discontinuity for fatigue evaluation. ./
M*' 's. RAI 25b Also indicate on Figure 3.6.1.1-12 and -13 (4), the locations corresponding to the high stress intensity values given in Table 3.8.4. 3.1 (4).
RESPONSE
The calculated maximum stress intensity values at element locations in the regions of discontinuity are shown on Figure Item 25-A, Attachment RAI 25b-1. Locations corresponding to high stress intensity values given in Table 3.8.4.3.1 are marked on Figure 3.6.1.1-12, Attachment RAI 25b-2 and Figure 3.6.1.1-13, Attachment RAI 25-2, (attached). e 9 C
, - - - . . - , ~
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CAROLINA POWER & LIGHT COMPANY BRUNSWICK STEAM ELECTRIC PLANT UNITS 1 & 2 e S Attachment RAI 25b-1
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4-i s. RAI 26 Provide a justificatj = . for using a stress concentration factor of 2, stated on page 3-48 (4).
RESPONSE
Justification for using a stress concentration f actor of 2 for the weld at the downcomer/ ring header intersection is provided in Welding Rescerech Council Bulletin 256, " Review of Data Relevant to the Design of Tubular Joints for Use in Fixed Offshore Platforms", by E. C. Rodabaugh, January 1980. ( l l
,n 's RAI G1 With reference to Section 1.5 in the PUA report (4), justify the assumption that the fluid, structure interaction and dynamic load factora for post-chugging are the same as those for condensation oscillation.
RESPONSE
The condensation oscillation load is an oscillatory pressure on the submerged portion of the torus shell. The load is symmetric, having the same value and distribution at all vertical cross sections of the torus along the circumference. The pressure distribution on the wetted perimeter of the torus changes from zero
. at the water surface to the maximum possible value of 15.37 psi at the bottom dead center of the torus. The load is defined in the frequency domain for the frequency range of 1 to 50 Hz.
This load is applied to the torus for a duration of 30 or 900 seconds depending on the type of break. The post-chugging load is also an oscillatory pressure on the submerged portion of the torus shell. The post-chugging load distribution is similar to'the condensation oscillation load along the torus circumference and the wetted perimeter of the torus cross-section. The maximue possible pressure at the bottom dead center of the torus shell is 3.7 psi. This load is applied to the torus for 30 or 900 seconds depending on the type of break. The
. cycle duration of the post-chugging load is 0.5 seconds every 1.4 seconds for the total duration of the load. This load is s
RAI G1 (Continued) RESPONSE (Continued) defined in the frequency domain for frequencies of 1 to 50 Hz. Table G-1 shows the frequency versus pressure amplitude for condensation oscillation and post-chugging loads. The amplitudes are normalized to the absolute sum of the amplitude of each load for better comparison. The fluid-structure interaction (FSI) factor for the torus depends on the ratio of water mass to torus mass, which is constant for both condensation oscillation loads and post-chugging loads. The variation in these two loads has minimal effect on the FSI factor. Due to very low ratio of water to torus mass the maximum FSI factor obtained for condensation oscillation load is 1.13. The dynamic load factor (DLF) for post chugging load may be different from the dynamic load factor obtained for condensation oscillation load. Since the post-chugging load appears to have relatively higher amplitudes at higher frequencies than the condensation oscillation load and the structural natural frequencies starts from 27 Hz., the dynamic load factor for post-chugging load may be higher than 1.5. A simplified conservative approach using the natural frequencies of the torus and assuming the torus as a single de~ gree of freedom system results in an upper bound DLF for post-chugging load. A DLF of less than 3 was obtained from the analysis. ( k l l 4
r*' e RAI G1 (Continued) - RESPONSE (Continued) In light of this conclusion, the DLF used for hydrodynamic loads was increased to 3 and the stresses calculated. A discussion of the results are given on Page 1-34 of the PUAR. The stresses - are all below allowables. It was concluded that the hydrodynamic loads are not appreciable loads in the concrete torus. For additional information see RAI G3. 3 m a t
- k. .
T:bla C-1 fRC kbZ Gq Comp 2ricen cf Condin3cticn Oscillotien and Post-Chugging Loads Frequency Versus Pressure Amplitude (Normalized to the Absolute Sum of Amplitude of Each Load *)
,r .
Frequency Amplitude Frequency Amplitude Hz. C.O P.C Ratio (CO/PC) plz. C.0 P.C Ratio 1 0.019 0.011 1.73 26 0.016 0.011 1.45 2 0.016 0.011 1.45 27 0.0 37 0.077 0.48 3 0.021 0.014 1.50 28 0.008 0.047 0.17 4 , 0.0 31 0.014 2.21 29 0.012 0.032 0. 38 5 0.077 0.016 4.81 30 0.009 0.024 0.38 6 0.174 0.014 12.42 31 0.005 0.008 0.63 7 0.027 0.0 27 1.0 32 0.002 0.005 0.40 8 0.025 0.027 0.93 33 0.002 0.005 0.40 9 0.0 25 0.027 0.93 34 0.002 0.005 0.40 10 0.025 0.027 0.93 35 0.003 0.005 , 0.60 11 0.051 0.016 3.19 36 0.005 0.008 0.63
<- 12 0.029 0.014 2.07 37 0.007 0.014 0.50 13 0.008 0.008 1.0 38 0.005 0.008 0.63 14 0.005 0.008 0.63 39 0.0 04 0.011 0.35 15 0.005 0.005 1.0 40 0.006 0.011 0.55 16 0.007 0.005 1.40 41 0.021 0.041 0.51 17 0.003 0.003 1.0 42 0.021 0.041 0.51 18 0.003 0.003 1.0 -
43 'O.021 0.041 0.51 19 0.003 0.003 1.0 44 0.021 0.041 0.51 20 0.018 0.011 1.64 45 0.021 0.041 0.51 21 0.013 0.008 1.63 . 46 0.021 0.041 0.51 22 0.020 0.014 1.43 47 0.021 0.041 0.51 23 0.022 0.014 1.57 48 0.021 0.041 0.51 24 0.021 0.014 1.50 49 0.021 0.041 0.51
? 25 0.010 0.011 0.91 50 0.021 0.041 0.51 b
- Absolute sum of C.O. load amplitude = 15.37 psi Absolute spa of P.C. load amplitude = 3.75 psi
l I (~ RAI G2 Provide information indicated that the water mass has been included in the seismic analysis of the Torus. Provide justification if it has not been considered.
RESPONSE
The water mass, about 16 percent of the total weight of the torus, ~
, is located at the bottom portion of the torus. It was included in the original response spectra seismic analysis of the torus and dry well. The actual distribution of the acceleration is linear, with maximum acceleration at the top of the torus decreas-ing to a smaller value at the bottom of the torus. The maximum horizontal and vertical accelerations at the top of the torus were aprlied to the entire torus in the quasi-static seismic analysis to account for the actual distribution of the seismic f' loads.
The calculated shearing stresses due to seismic and other loads are ' far below the allowable stresses and inclusion of the water mass, though not justified, does not impair the integrity of the torus. \ i l i e 1
l s RAI G3a With reference to Section A-1.2.5 of the PUA report (4), provide information on the stress results for condensation oscillation load at modal points listed in Table A-1-10 (4).
RESPONSE
The stress results for the condensation oscillation load applied statically to the torus shall are shown in Attachment RAI G3a. The maximum radial stress is 34 psi, the maximum azimuth (circum-ferential) stress is 9.8 psi, the maximum axial (meridional) stress is 15 psi and the maximum shear stress is 13 psi. The very high dynamic load f actor, listed in Table A.1-10 corresponds to stresses having a value less than 1 psi and were not used. , Dynamic load factors in the range of 1.5 to 3 occur at a few
- nodes. The remaining nodes have a dynamic load factor of 1.5 which was considered typical and used in the analysis for all the stress components with insignificant error.
Stress levels in the concrete torus due to the condensation oscillation and other hydrodynamic loads are low when compared with a steel torus. This is due to the continuous support at the I- bottom of the concrete torus and the very low ratio of water mass to total mass of the torus. To study the effect of hydrodynamic loads on the stresses in the torus for various load combinations given in Table 1.4-1 of the PUAR, hydrodynamic load factors were increased to 3 and stresse's for various load combinations obtained. The stresses were increased between 0 to 40 percent (see page 1-34 f C
RAI G3a (Continued)
RESPONSE
of the PUAR for more information). The maximum stresses remain below allowables. Those limited areas of the torus and loadings which show a dynamic load factor between 1.5 and 3 are enveloped by these results. REFERENCES
- 1. NEDO-24583-1 (79NED125)
" Mark I Containment Program, Structural Acceptance Criteria, Plant Unique Analysis ' Application Guide".
General Electric Company, San Jose, California October 1979 l [ 2. NEDO-21888 l " Mark I Containment Program Load Definition Report" !- General Electric Company, San Jose, California November 1981, Revision 2 i I 7 L. 2-
"- +- .c ,wy--- -r-w. y , - ....-,.m,_ ,m,, , ,_,, ., .. _,
COMPUTER PRINTOUT
. STATIC ANALYSIS
. STRESS RECOVERY PHASE ,
FOR f C.O. LOADS , (Total 2 pages) CAROLINA POWER & LIGHT COMFANY BRUNSWICK STEAM ELECTRIC PLANT UNITS 1 & 2 4 Attachment RAI G3a e ee-
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RAI G3b Also, provide justification for considering those stresses which yielded dynamic load factors greater than 1.45 not re)iable or typical. 1 8
RESPONSE
See Response to RAI G3a. e 4 .i [ e l i
... j
\ RAI G4 Provide details and schedules for the relocation of the electric penetration box which was found to be overetressed as indicated in Section 2.3.2.9 of the PUA report (4).
RESPONSE
Suppression pool electrical penetrations X-232-B and X-232-C, presently equipped with 24" x 24" x 24" termination boxes at the inboard end, have been modified by Plant Modifications PM 81-251 and PM 81-252. Plant Modifications have provided necessary design and quali-fication documentation for:
- a. Removal of the termination boxes from inboard end of
< the above electrical penetrations.
- b. Substitution of the termination points by environmentally and seismically qualified cable splices, manufactured by Raychem Corporation.
D. G J
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' RAI G5 With reference to Table 2.3.2-3 of the PUA report (4),
indicate the schedule for modifications of the platform structure.
RESPONSE
The modifications to the platform structures are to be accomplished for Unit 2 during the maintenance outage scheduled to start in November 1983 and for Unit 1 during the refueling outage scheduled to start in July 1985. e 9 O l* w d at *
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