ML20065B211

Forwards Response Prepared by Nutech Engineers in Response to NRC 820719 Request for Addl Info Re plant-unique Analysis Rept,Including Reasons for Not Analyzing 180-degree Beam Segment

ML20065B211
Person / Time
Site:	Fermi
Issue date:	09/09/1982
From:	Tauber H DETROIT EDISON CO.
To:	Youngblood B Office of Nuclear Reactor Regulation
References
EF2-59-222, NUDOCS 8209140298
Download: ML20065B211 (37)

Text

{{#Wiki_filter:-- H:rry Truber Ng$ YgYM Cmoucte Detroit Edison R?!E"4= Sepbater 9,1982 EF2 - 59,222 Mr. B. J. Youngblood, Chief Licensing Branch No. 1 Division of Licensing U. S. Nuclear Regulatory Commission Washington, D. C. 20555

Reference:

Enrico Fermi Atomic Power Plant, Unit 2 NRC Docket No. 50-341

Subject:

Mark I Containment Request For Additional Information Attached please find our response to your July 19, 1982 request for additional information on the Fermi 2 Plant Unique Analysis Report (PUAR). Due to the time constraint, the response is submitted in the question / response format. After you have reviewed and accepted our response, we will incorporate the attachment into the PUAR, revising PUAR pages if applicable. If you have any questions regarding the above, please contact Mr. Larry E. Schuerman, (313) 649-7562. Sincerely, / f WA (2V L Attachment cc: Mr. L. L. Kintner Mr. J. Ranlet (Brookhaven National Laboratories) Mr. W. Seagraves (Franklin Research Institute) 8209140298 820909 0 PDR ADOCK 05000341 'i h PDR

'7,n Question 10.1 With regard to suppression chamber analysis, provide justification for not analyzing a 180 beam segment in- ,',. '~' cluding the torus, columns, and seismic restraints as required by the criteria for considering the effect of, seismic and other lateral loads. Also, discuss the im-plications of this approach with regard to stresses in the suppression chamber in the region aurrounding the support columns. Response to Question 10.1 The approach used in the Fermi PUAR to evaluate suppression chamber lateral loads results in total lateral leads which envelop those which would be obtalned by a 180 beam model. As discussed in PUAR Section 1-2.4.2, maximum accelerations and dynamic load factors'are used to develop bounding values of lateral loads for seismic loads and for asymmetric torus shell loads due to SRV discharge and pre-chug, irrespective of.the suppression chamber frequency. Specifically, the maximum OBE spectral acceleration of 0.23g is used for-seismic loads, the maximum dynamic load f actor of 2.60 is used for SRV discharge loads, and the maximum dynamic lo3d factor of 13.9 is used for pre-chug loads each of which-ccc.trs at a different frequency. The resulting lateral loads for the seismic, SRV discharge, and pre-chug loadings are added absolutely to obtain a bounding value of the total suppression chamber lateral load, which is conservatively assumed to be transferred by two of the four seismic restraints. a s DET-15-020 1 nut.e_c._h. Revision 0

o ~ The total lateral load which wculd be produced by a 180 beam =cdel would be less since the respcnse to each loading would primarily be determined at the dcminant l'ateral frecuency of the suppression chamber. Lateral loads result in a shear effect and an overturning moment effect en the suppression chamber. The horizontal shear effect is the =cre significant ccaponent and is resisted by the seismic restraints shown in Figure 2-2.1-10. The t overturning mcment effect results in vertical leads which are resisted at each mitered joint by the suppression chamber columns and mitered joint saddles shown in Figure 2-2.1-4. The vertical loads on any cne colt =n/ saddle assembly are small ccmpared with those caused by the ma or torus shell loadings which primarily act in the vertical directicn, the results of which are shown in. Table 2-2.5-4. The corresponding stresses in the suppression chamber shell adjacent to the column / saddle assembl}y due to the overturning =ccent would aisc be small. s .DET-15-020 2 . Revision 0 nut.e. ch

Question 10.2 With regard to the assumption that only 20% of the total mass of water in the suppression chamber contributes to lateral seismic loads, provide justification to indicate the applicability of the tests cited in the PUA report to support this assumption. Response to Question 10.2 For a suppression chamber partially filled with water subjec-ted to a horizontal seismic excitation, a portion of the total water mass acts as a rigidly attached mass while the remaining water mass acts in sloshing modes. The effective weight of water which acts as a rigidly attached mass was determined from 1/30 scale generic tests performed as part of the Mark I program effort. The seismic slosh test facility is identical and the test procedures used to determine the effective water weight are similar to the tests described in General Electric Report NEDC-23702-P, " Mark I Containment Program Seismic Slosh Evaluation", March 1978. The 1/30 scale model test facility is based on a prototypical Mark I suppression chamber whose gecmetric characteristics l are very close to'those of Fermi 2. Tests were performed i' with three different support stiffnesses (rigid, medium, l and flexible), which covered the range of stiffnesses and i l frequencies for all Mark I plants, including Fermi 2. The analytical model developed for use in the referenced study predicts that 20% of the total mass of water acts l as a rigidly attached mass with the suppression chamber. This Dd DET-15-020 3 t Revision 0 i i

prediction was verified with the 1/30 scale model by comparing the measured frequencies of the test facility with those obtained frcm the analytical medel. This was done by first adjusting the emptied test facility support stiffness to that necessary to cbtain the frequency cf the empty suppression chamber predicted by the analytical =cdel. The test facility was subsequently filled with water to a height belcw the equator, and a series cf tests were performed to determine the frequency. The resulting frequencies ccmpared favorably with those obtained using the analytical mcdel with the same assumed water height. The test results therefore confirm the analytical results which showed that 20% of the total water mass acts as a rigidly attached mass. These results are considered applicable for use in evaluating the Fermi 2 suppression chamber response to seismic loadings. The evaluation of the Fermi 2 suppression chamber for horizontal seismic leads is discussed in Section 2-2.4.2 of the PUAR. The seismic lateral load is conservatively calculated assuming that 20% of the total water mass acts at the maximum spectral acceleration of 0.23g. The remaining'80% of water is assumed to act at the maximum accelerations in the range of sloshing frequencies. The methodology accounts for 100% of the water and results l t in a bounding value for the suppression chamber lateral load due to seismic loads. nutech DET-15-020 4 l Revision 0

Qu_estion 10.3a Provide detailed calculations to indicate how the modal correction factors given in Sections 1-4.2.3 and 2-2.4.1 of the PUA report are obtained. Response to Question 10.3a Mcdal correcticn factors used in calculating the respcnse due tc SRV :orus shell 1 cads are cbtained by dividing the respcnse of an initial value or free vibration problem by that of a transient forced vibration response problem. The physical representations for the analogs used in developing these correcticn factors are shown in Figure 10.3-1. Two representations of source pressure in a rigid tank are shown. The transient response problem consists of a rigid torus with a pressure source, P which is prescribed as a decaying cosine B, function. The initial value problem consists of a similar torus which contains a spring and disk mechanism for providing an impulse to the surrounding pool water. The analogs, described in terms of masses and springs, are shown in Figure 10.3-2 for the transient response er forced vibration problem and the initial value or free vibration problem. The torus system is described in terms of a generalized stiffness, k and a generalized mass, m s, s' DET-15-020 5 Revision 0 nut.e. ch.

I 1 The forced vibration analog is subjecced to an applied loadi..g, described as P_e 'i+* cosa PB(t) t = o B

where, bubble force at time equal to zero P

= B attenuation constant i = frequency of the SRV bubble w = B time t = The free vibration analog incorporates an additional mass and spring, representing the bubble system, which is utilized to establish a frequency of the bubble oscillation. The apparent or effective mass of the bubble is defined as m. g The numerical value of an apparent bubble mass is estimated by averaging the hydrodynamic mass for the case of an occillating sphere in a still fluid and a fixed sphere in e an oscillating fluid. The bubble stiffness is computed by multiplying the bubble mass by the bubble frequency squared. f Four spherical bubbles are assumed for the T-quencher J i discharge device. Single bubble stiffnesses are additive l since the bubbles are assumed to act in phase (i.e., f parallel springs). i i i 1 \\ DET-15-020 6 j Revision 0 .. ~ - - - - - ~

l The initial conditions for the free vibration analcg are a distance, al, prescribed by compressing the spring, kg, such that at time equal to zero the force in the spring Therefore, at time equal to zero, both the is equal to Pg. forced vibration and free vibration analog have the same applied load magnitude. Damping for the torus system can be described in terms of load attenuation and structural damping. Based upon test observations, it is assumed that the structural response will be decayed in the same manner as the prescribed pressure. -D Accordingly, a decaying exponential function of the form e is used to represent load attenuation and structural damping r in the solution of the forced and free vibration system, respectively. The equations of motion for the free vibration analog described 4 in Figure 10.3-2 are obtained from free body diagrams for i the structure mass and bubble mass as i i (k +k ) X* BB" l mX sg s B s f BB BB+ Bs" m

where, structure mass displacement l-X

= s t structure mass acceleration X = s bubble mass displacement X = B bubble mass acceleration j X = B I Md j DET-15-020 7 i Revision 0 a y __m. _,,.c r-

The solution to these equations expressed in terms of the structure response, X is given as the following function, s, EB -At B B B 1 T' T' m-) I f Xs"T* s s s s

where, "B

bubble frequency = structure frequency w = s The other variables have been previously defined. A family of curves which represents dynamic load factors as B "S) is generated by assuming I a function of frequency (w either (k !"s) r (m /m ) as constant. Based upon estimates B 3 s of the significant modal characteristics of the torus and the oscillating bubble, a range of values for (kB "s) and (k !"s) values is (m I*s) is established. The range of B B estimated to be about 160 to 1600. The range of (m /m I g s values is estimated to be 0.03 to 0.3. Figure 10.3-3 contains a comparison of DLF's for the cases-with (k /*s) equal to 160 and 1600, assuming a structural B -frequency of 20 hz. The maximum DLF for the case with (k /*s) equal to (k !*s) equal to 160 is 2.8, whereas for B B 1600, the maximum DLF is 2.0. It was determined that the DLF's for the case with (k !*s) equal to a constant are B about 20 to 60 percent larger than the DLF's for the (m /*s) equal to a constant. case with B OUkQhb DET-15-020 8 Revision 0

s i The equation of motion for the forced vibration analog 2 described in Figure 10.3-2 is obtained from the free body I diagram for the structure mass as i -At mX -kX

• PB*

B j ss ss l The solution to this equation is given in the following 4 function: -At f gg ,, y) Xs" s 1 4 s i Figure 10.3-4 contains DLF's plotted as a function of (w I"s). A family f DLF curves is included for structure B j frequencies of 11, 14, 20 hz. The DLF's for the resonant condition range from 5.4 to 9.,9 for structural frequencies between 11 and 20 hz. The-forced vibration DLF's are approx-imately 3 to 5 times the free vibration DLF's. Correction factors are obtained by dividing the free vibr-ation response to the system by the response of the forced vibration system. The sensitivity of the correction factor (k /*s. '"B *s, Al is evaluated in order to j to the variables b determine a valid set of correction factor curves to be ) used in design. i-i I Based.upon the above evaluation, it was determined that the lower range of (kB *s) should be used for determining conservative design basis modal correction factors. An i l i DET-15-020 9 Revision 0 i

attenuation factor of 6 is selected, since use of this factor results in correction factors which bound the response at resonance conditions for all structural fre-quencies. A typical correction factor curve is shown in Figure 10.3-5 for a structural frequency of 20 hz. For the plant unique analysis, a set of enveloping correction factors is gene-rated for dif ferent modal frequencies of interest as shown in PUAR Figure 2-2.4-5. Table 10.3-1 contains a comparison of analytical results obtained using modal correction factors and measured results for Monticello. The results shown are'obtained by dividing analytical results by test results for key response parameters. The comparisons show that modal correction factors provide a conservative basis for cali-brating the analytical model used to evaluate the response of the Fermi-2 suppression chamber for SRV torus shell loads. The modal correction factors are developed at test conditions and applied at design conditions in accordance with NUREG-0661. DET-15-020 10 Revision 0 4 1 11Utg,Q),) l I

T Test Condition Location Component Cold Pop Hot Pop a) Stress Intensity 78 From Inside She'll Membrane 1.2 1.3 Equator Midbay 78* From Outside Shell Membrane 1.4 2.3 Equator Midbay b) Column Reaction Inside Support Upload 3.7 3.0 Column Inside Support Download 2.8 2.0 Column Outside' Support Upload 4.4 3.5 Column Outside Support Download 2.7 2.8 Column 4 Table 10.3 Corrected Transient Response Analysis Normalized by Test Response DET-15-020 11 Revision 0 nutg.gh

4 gdI k ~ 'C * ** B T TIME ZEHO: PB A PB (g) = PB* us o e8 -"$~b"$$$hk"/.92 s. ,,gfj:$:Tb"-$- bh$ *.[.:x"'s i b-sy.v - me " ' "sy, +s. / " ' "0:<. .v ..v. .-x y, // ""v

p. m "9

. O g ,e y y

s i :s o o"

e N g A"t ~@) 39 "+"s X" ~X SS. M,s f' i (c::' s. i b;.- Q- ,.:Q O' Q .q ~ =5=- 0:# $9 ' : = T.: v,. C.. A .y r. "y 0 e Pg(t) ,jf/ -{ ,'j[ ~~ x ?. yt i 4. ~ k t./. 4. u ,.y .:$$7 _)_ _ J' ,g,gg3 % g g<ss # P. P. a w N a) THAt3SIENT DEFINITION li) INITIAL VAll)E DEFINiilON -i ^ Figure 10. 3 SRV Load De finition 1 i

1 l BUBBLE mB Pg(t) ,I X g k g 9 FLUID FLUID ms m s TCRUS TORUS I Xs Xs m. m (a) (b) FCRCED V!BRATION ANALOG FREE VfERATION ANALOG l i Figure 10.3 Structural Response Analogs l l i l i l l DET-15-020 13 Revision 0 nutQQj] i

1 i* ijijiil lIl_ll IIIIIIIilIt iIIt 5 ! I~i T I T r-h [ ll I llllf f ]!!!!!! 1.0 ,.6 .i.......iiii .,t 6 4. .i.! o I iiieai4a4664i4 66i 6 l ltlj j i ll t1i iitiI y liLEGEND: i l ll lll f lll l y' me j ] k8 /ms 160 0 .... --- - k g / ms=1600 ,l' j, lllllllllll!lllllllI!!!!llll!!!!IIII! 0 .s 3.0 s.s 2.0

• s'*s Figure 10,3 Free Vibration Analog Dynamic Load Factors (k !*s) Constant For B

DET-15-020 14 Revision 0 nutggh

n l i l l i e i 0.0 - J - EN 1 1 4 i 6, ! ii! LfM Ni4 I + l l i i 1[ i g" iII I ~N'iilil tt8 i i I I i! ~ O 56i M ! ll l ! lNb! !!!I 2o 2 5 LEGEN0;

• i! l I b

7 1 t 44 iitI t i ii1 4 i 11 11 lil Illillil T............,. ".. " > i e = g H, l ~ i Q w s = 2 0. Hg T iii tiiii.. i i i i ti IllIlIllllllIlll!l!II l!lIIIIIIIIll'I' O .s i.o i.s 2.o

• s'**

Figure 10.3 Forced Vibration Analog Dynamic Load Factors l l l DET-15-020 15 Revision 0 nutagh

l l 4 i I I! IIIIIIIIIIIl 1 I. '8 III 'k il l# liII i .6 l l l l NlVIII IIII 4 = l l l l l l l l\\ /l i I I I III i { l l l l l l Y3Cl l l l l l l 8 IIIIIII IIIilIIIIII II I I IIIIII IIII II II I I I I I I I I I I I I I I I i l i i 0 .5 1.0 1.5 2.0 i i "B / *s Figure 10. 3 Typical Modal Correction Factor 1 DET-15-020 16 Revision 0 nutggh

Question 10.3b Provide justification for the applicability of these factors to "multidegree of freedom" systems since the factors were developed using simple "one degree of freedom" systems. Response to Question 10.3b . ". a.-.e.~.... - - m "y "y.- a c.=..' w.". w"..-...'.".e. c" ^# ..". a t."a n s i t'_ n t." =-.c., "yC ". s t' v-r " ^ ~ =~ -~s e k, 4..a A ,s.4 n .u.. c....C -,- A Ave n .c.~.v . n.e. s s b.a n, n n,A .s v s v o ""yy "." C a #.."., '. b. c ^ ^ " '

• 4 ^ ".a~

^#

• "4 C

s " "ye " " C.C. 4 * # ^ ". ~.. c.". ^ #. " s".'.". C. ,--.v. e .e ~..'.a . v 1 ~, _. _ A a n - = a n: : v a. a A n. w.'.. me ~. C t n.,.. e.v enga., 50 u- --r.-- w- --w. 3.. - - - a g g.: C r. C '". ' ~...'^ A ". ."."A. .# e C ". " "y.' a # _4 ".~ ^- .ce' v. c",."'"."'s.".*. ^# = iyy.- 0 #er"on ^# .# " = o A "w "o. s '; s * * ~..s. =. =. ".. ". s '- "' C ' '- " 3.1

1. "*~

s.;.". c.1 m _.,' e..w ';

1. . a a d C ~.. = "; s * *-.~..

o _" "..C d e..c. " o "y." e s e.". '.- c d '.^ "; u" s i.". y^.' a ueC,,"ee C# The resc.onses Cf each sinc.ie dec.ree of freedCm system - - r "y ". o s c.4 o r. c '- - 2 4.,. ~.k.e m*,1 ." e s "yonse C.# u. e g o..~..e A

• ". e c""

-r oue- --w-. -w b C..,.m. k e.*.. D u..4... c,

• b.e s u.r... 2 4 ^ ". " " v^ C a s s,

~.C C' a ' ^ o " " o C - 4 ^ ".. a C

• c." s w

v. y. -eyl4pA %.e .v a,. s y C,s e-vbtm,.;,pA

e n n..

a. L'e'..o, 4.,.,..v e 1.

a.. 9

-D ,wa 2-n m n .w. .9 e -v ~ o. A ". ^w h .C.,u "y "y " e s s _4C ". o# c.' C ". s 4 ". ;~.I e # e y~~."c c c#.#"GedC ".1 s V, s *e

1. ^"

Cw.amOer f rec.uenCV.. As CisCusseC in O n.e rescC"se tO C.uestiCC 10.3a, '.h e ~~^ d =.1 r _" _ a C - d o." _# a C

• o. s a _- a A e ". e.' C "eed u s _i.. ~

^ sinc.le Cecree c: IreeCCm systems and are CCmc.atible :or 3, s e .4. 4he- ..v C 2 _, s t. e -, C s. ; ; n,.,..... u e _-.. o C s _4.,. C e. _4n s ...e wm..oA 4s w - e .r CCmpletely linear. DET-15-020 17 Revision 0 nut _ech.

Question 10.4 Provide justification for not considering the effect of bending moments in column analysis using interaction for-mulae. Response to Question 10.4 Ccnsideration of column bending mcments using the interaction formula is necessary when ccmpressive stresses er column buckling is a concern. Since the suppression chamber support columns for Fermi 2 are heavily reinforced and braced continuously'alcng their length,'as shcwn in Figure 2-2.1-7 of the PUAR, the effects of buckling are negligible. Further-more, bending =cments in the support columns are small since the columns are cermitted to slide horizcntally at their base. The support column loads shcwn in Table 2-2.5-4 are substantially below the allowable compressive loads for the support eclumn wide flange sections with cover plates. DET-15-020 18 Revision 0 l nutech ENGarvEERS

Question 10.5 With regard to the suppression chamber columns, provide justification and/or additional information to indicate why a nonlinear time history analysis was not per fo rmed as required by the criteria when net tensile forces are produced in the columns. Table 2-2. 5-2 o f PUA report indicates that net tensile forces are produced in the columns. Response _to Question 10.5 The criteria requirements for performing a non-linear time history analysis are applicable for plants in which the suppression chamber and its supports are nct anchored to the basemat. Such a condition would result in gross non-linear behavior if uplift loads exceeded the weight of the suppression chamber and contained water. The Fermi 2 suppression chamber is fully anchored to the basema at each mitered joint column and saddle base plate location, as shown in PUAR Figures 2-2.1-7 and 2-2.1-3. Although tensile forces are produced in the column and saddle supports, the tensile forces are less than the allcwable anchorage capacity of the support system, as shown in PCAR Table 2-2.5-4. The requirements for a non-linear analysis, therefore, need not be evaluated for Fermi 2 since the suppression chamber is fully anchored to the basemat and the effects of non-linearities on the overall suppression chamber response have been minimized. DET-15-020 19 Revision 0 nut.e. c..h

Question 10.6 Provide justification for using two different temperatures, 173 F for suppression chamber and vertical support systems and 100 F for the base plate of the support system, for cal-culating the respective allowable stresses. Response to Question 10.6 The allowable stresses for the suppression chamber and its vertical supports are conservatively determined at 173 F since this is the maximum temperature specified for any LOCA event, as shown in PUAR Figures 2-2.2-4 through 2-2.2-6. The ~ allowable stresses for the vertical support system base plates P are determined at 100 F which bounds the maximum temperatures of the base plate expected during the specified events. There may be long term conditions which result in higher base plate temperatures, however base plate temperatures higher than 100 F expected to occur during times of peak transfer of are not hydrodynamic loads to the suppression chamber vertical support system. Furthermore, the allowable stresses at 100 F and 173 F are not significantly different. DET-15-020 20 Revision 0 nut _ech.

Question 11.1 Provide justification for using SRSS method to combine the SSE and LOCA responses for SRV piping analysis instead of the absolute sum or cumulative distribution function approaches as required by the criteria. Respon_se_.to Question 11.1 m_ The method of cc=bining responses due to L^CA and SSE loads .o-c=.e, _43.4, ~ d a_ s - _' " e d c -ec._^.. C# ~..".e._'_=.. o y, 3 - Unique Analysis Report (PUAR) is based on 1:RC document 2:UREG-0,' 3,', .c.e ". 4 s _i ^. ', ".'f.o.i.^dc'o^3.v

1. ^-

C... " '... _4.. ~ ".v.=..._' c C a_ s ~ v...e e.=. ", v. 3.. -,. " _1 4 s ". e d 4.. .'4= v,, _ 4. :.0. ".'.".a_ o _' 3 _4.. = ' _sa"m.a ^# " C o.r." - O,' e: . " ". _ _ _ = _ _ , - - ~ _ ~ --.,. e -~ ccmbination of responses due to LCCA and SSE within the reactor coolant pressure boundarv. usinc. the SRSS tec hnic.ue. The curren: Revision 1 has e:: tended the application of this combination technique to include ASME Section III, Class 1,2, and 3 systems, components and supports. As described in Revision 1, use of the ep.es me,.4-oe y-m.y, 4a ^# e,' e _- ' =... s- - es a. .-excea_d=..ca - - - = " _4 _' _4 _ v_ ^ ~ y or higher. Since the rermi 2 SRV piping is Class 2, the use of the SRSS method is acceptable based on Revision 1 of :UREG-04S4. DET-15-020 21 Revision 0 l nut _ec_h.

Question 11.2 Provide justification for using Markl's equation for fatigue analysis of SRV piping instead of the SN curve given in ASME Code Section III, Division 1 Appendices. Response to Question 11.2 The methodology for evaluating Fermi 2 SRV piping fatigue was originally presented to the NRC during a meeting in June of 1981 and was followed by a letter submitted to the NRC shortly thereafter (Re fe rence 11.2-1). Section 3.9.3 of Supplement 1 of the Fermi 2 Safety Evaluation Report (NUREG-0798) refer-ences the required fatigue evaluation and the proposed method-ology. Since the SRV piping is a Class 2 system, the approach out-lined in the presentation and letter was to evaluate fatigue using ASME Class 2 piping rules as a guideline. The proposed methods included extension of the Class 2 equations and curves used for thermal fatigue evaluation to include all cyclic loads. A comparison of the extended Class 2 method to a Class 1 fatigue analysis was also provided which showed that the two methods yield similar results. Using the proposed methods, a fatigue usage factor is determined for each of the cyclic loadings. For Mark I LOCA related loads, estimates of total stress cycles during plant life would be determined and associated fatigue usage would be calculated. Since only very conservative estimates of the number of SRV DET-15-020 22 Revision 0 i 4

discharge related stress cycles were available, the approach proposed that SRV actuations would be monitored to assure that the allowable fatigue usage was not exceeded. Following Detroit Edison Company's commitment to the NRC to perform an SRV piping fatigue evaluation, the matter was dis-cussed between the NRC and the Mark I Owner's Group. These discussions resulted in a commitment by the Mark I utilities to perform fatigue evaluations for SRV piping in the torus and for torus attached piping systems as part of the plant unique analyses. Discussions among Mark I Owners and their AE's followed and a task force was formed to develop a generic approach for fatigue evaluation. The approach agreed upon was a method which extended the Class 2 piping fatigue rules similar to the methods initially proposed for Fermi. Refinements to the proposed Fermi 2 methods which have been incorporated into the generic approach consist of the following: o Fatigue usage is evaluated based on considering critical loading combinations instead of on an individual load basis. DET-15-020 23 Revision 0 nut.ech.

o Total cumulative fatigue usage for all cyclic loadings is calculated in lieu of monitoring SRV actuations. o The allowable number of stress cycles is determined by using Markl's equation (Reference 11.2-2) in lieu of the Class 2 thermal fatigue equation basis. (Markl's equation forms the basis for class 2 piping fatigue and was used in developing the Class 2 piping Stress Intensification Factors). o Actual stress cycles for a given response time-history are converted into equivalent full stress cycles using the methodology defined in Section NC-3611. (e) (3) of the Code. The SRV piping f atigue evaluation per formed for Fermi-2 and documented in Volume 5 of the PUAR includes the extended Class 2 approach originally proposed for Fermi-2 and incor-porates the additional refinements included in the generic Mark I approach. The refinements result in a more practical, comprehensive method of evaluation for fatigue. Re fe rence 11.2 Detroit Edison Company Letter EF2-53,824 to NRC dated June 22, 1981. Reference 11.2 Markl, A.R.C., " Fatigue Test of Piping []({f()()h) Components", Transactions ASME, Volume 74. ENO4PoEERS DET-15-020 24 hufAMiwct3

Question 11.3 Provide justification for not using Equation 11 of ASME Code Section III, Subsection NB for calculating the fatigue stresses, and explain the method used. Response to Question 11.3 Justification for not using Equation 11 of ASME Code Section III, Subsection NB (Class 1 Piping) is provided in the response to Question 11.2. Equation 11 of Subsection NC (Class 2 Piping) of the ASME Code provides combination methods for thermal and other sustained loads used in evalu-ating for fatigue. The methods applied in the Fermi Plant Unique Analysis Report extended traditional usage of Equation 11 to combination of stresses due to dynamic cyclic loadings, using the same method of absolute summation of stresses. DET-15-020 25 Revision 0 nutech EPeG4NEERiin

Question 11.4 Provide justification and reference for the maximum-stress cycle factors given in Table 5-2.4-4 of the PUA report. Response to Question 11.4 See the response to Question 11.2 for a description of the methodology for evaluating Fermi-2 SRV ? ping fatigue. The i basis for developing R factors used to determine maximum equivalent full stress cycles is derived from the Class 2 piping thermal fatigue techniques defined in Section NC-3611. 2 (e) (3) of the Ccdc. R factors for individual dynamic cyclic loadings also take into account consideration o f loading characteristics such as frequency, time-history and random phasing of load components. DET-15-020 26 Revision 0 IIbl k () ENGerdeLEAS

Question 11.5 Provide the magnitudes of the dynamic load factors used in Tables 5-3.2-1 to 5-3.2-3 of the PUA report and the justification. Response to Question 11.5 The dynamic icad factors (DLF) included in the leads specified in Tables 5-3.2-1 through 5-3.2-3 cf the PCAR are summarized in the attached Table 11.5-1. The loading functicns for water jet impingement loads, T-quencher thrust loads, and T-quencher end cap thrust loads are defined as rectangular pulse loadings. The maximum DLF s ecified by standard structural dynamics handbcoks

or this type of load function is 2.0.

DLF 's f or SRV air bubble dr ag loads were determined using Monticello in-plant test data as permitted by SUREG-0661. The criteria states that actual measured pressure wave forms determined in tests may be used to develop a maximum structural amplification for reacnant conditions. Using the measured Monticello pressure waveforms, a maximum DLF of 3.0 at resonant conditions was developed and is used for structures whose natural frequency is within the 4.0 to 14.0 He frequency range of the SRV air bubble drag loads. For structures whose natural frequency is well above the maximum air bubble drag load frequency a DLF of nutech DET-15-020 27

===a

Revision 0

1. -a

'a-~#a . 0 4s c nse-su'.4'e:' 'ead. " " e 's'.a' ' 'i ---n-- -- s of the T-quencher and its supper:s and the subr.erged SRV - ~ o v. e. .,. e...,.. ;...v.. at-b u ,o 5, u - - - - c - -..u..n n e.. :: a

a-

.,a p.;.n, -.-o n ~~-- load frequency, as shown in FUAE Figures 5-3.4-3 through 2-2.4-o. nutech DET-15-020 28 Revision 0 am

e f 4 i ,s 4 m Table 11.5-1 e PUAR Dynamic Load Factors for SRV Piping, T-cuench6rs, and T-cuencher Supports r" hi ,\\. 4 f 4 PUAR Table (D. 1 Dynamic.oac. actor L: Load Type Number n SRV water jet impinge-5-3.2-1 2.0 ment -) ,p s '5 4 SRV' air-bubble ' drag 5-3.2-1 and - 2.0 5 -3. ^.;-3 4 I kg I T quencher'and end cap thrust load +< 5-3.2-2 2.0 . s- .~ ,. -. s 4 s, d**, W _.g i s / ~ .. ? ~ ~ ' s t _ ? = k ( ~ ,g s. / DET-15-020 29 Revision 0 e s nutggh eP.

Question 11.6 Pro' vide the results of the analysis of bolted or welded connectior.c asscciated with the SRV piping. Response to Question 11.6 The tables in the prJAR which ccntain analysis results for wetwell SRV piping major support connections and welds are shcwn in the attached Table 11.6-1. The referenced tables show that SRV piping wetwell support connection stresses are within allowable limits. DET-15-020 30 Revision 0 nut.e_c_h.

Table 11.6-1 PUAR Table References for Wetwell SRV Piping Support Connections and Welds Support PUAR Table Number Connection / Weld Stress Vent Line-SRV 3-2.5-6 Piping Penetration and Welds SRV Piping Vent 5-245-5 Line and Vent Header Supports Ramshead and 5-3.5-4 T-quencher Arm Supports and Welds J I DET-15-020 31 i Revision 0 nutg,gh i

Question 12.1 Provide justification for the method of lumping additional fluid masses along the ringbeam, quencher beam (Page 2-2.10 3 o f PUA report), submerged length of SRV piping, T-Quencher and supports (Page 5-3.49 of PUA report) as indicated in the PUA report. Resoonse to Question 12.1 The hydrodynamic masses used for evaluating submerged structures are calculated using the relationships con-tained in PUAR Table 1-4.1-1 which are taken from LDR Table 4.3.4-1 (Reference 12.1-1). For the SRV piping, ramshead, T-quencher arms, 6 in. diameter lateral sup-port members, and 20 in. diameter lateral support beam, the hydrodynamic mass equations for a circular cyclinder were used. The hydrodynamic mass for the T quencher arm ring plate supports are calculated using the equation for a circular disk. For the ring beam and vertical quencher support beam, the hydrodynamic mass is calculated using the equation for a plate in the lateral direction and an 1-beam in the vertical direction. Reference 12.1 General Electric Report NEDO-21888, Revision 2, " Mark I Containment' Program Load Definition Report", dated December, 1981. DET-15-020 32 g)(Jggg[)[] Revision 0

3 Question 12.2 Provide-justification for not considering the loads indicated in Table 1 which are required in the analysis ~ according to NUREG-0661. Response to Question 12._2 { All loads specified by NUPIG-0661 were addressed in the PUAR. The loads identified in Table 1 which are not included in 4 i i Table 1-4.3-1 of the PUAR can be categorized as being negligible, not applicable to the Fermi plant, or are considered in the 4 i analysis. The loads circled in Table 1 are discussed in the paragraphs which follow. i 4.3.5 Froth Impingement. The froth impingement loads on the i t

• orus shell are negligible as indicated in LDR t

Section 4-3.5-1. The torus support system will i also have negligible effects due to the froth j impingement. For SRV piping, the portion below the vent header is protected from pool swell impact 1 j loads by the vent header deflector. The portion f below the vent line experiences negligible loads due to froth impingement. ) 4.3.8 LOCA Bubble Drag. The vent header support columns are i { the only structures above the bottcm of the downcomers 1 f and below the normal water level. The-LOCA bubble ] drag loads on these columns are contained in l PUAR Table 3-2.2-9. f i Md DET-15-020 33 Revision 0 _-.4--- -r-.- _,----r-

4.5.3 Chuc.c.inc. Vent Svstem Leads. The chuc. c. inc. loads en the main vent and vent header were considered in the analysis and are contained in PUAR Table 3-2.2-19. 5.2.5 T-Quencher Air Bubble Drag. The SR7 air bubble drag loads en these structures were considered in the PCAR. The SRV air bubble drag loads on the downccmers are given in PUAR Table 3-2.2-22. The SRV air bubble drag loads on the T-Quencher and the SRV piping are given in PUAR Table 5-3.2-3, and the S RV air bubble drag loads on the T-Quencher supports and vent header support columns are given in PUAR Tables 5-3.2-1 and 3-2.2-23, respectively. 5.2.6 Thrust loads on T-Quencher arms. The thrust loads on T-Quencher arms are given in PUAR Table 5-3.2-2. 5.2.7 SRVDL Environmental Temperatures. The SRV discharge line environmental temperature loads are discussed on page 5-3.17 of the PUAR. 5.3 Ramshead Loads. Ramshead loads are not applicable for Fermi 2 since the SRV Lines are ecuipped with the T-Quencher discharge devices rather than the ramsheads. DET-15-020 34 Revision 0 i nutech ENGarwEEftS

Question 12.3 Provide information on the analysis of the attachment welds of regions connecting the internal structures to the torus shell, indicating whether the criteria require-ments have been satisfied. Response to Question 12.3 The internal structure attachment welds to the tcrus shell have been evaluated in accordance with the criteria requirements. The attached Table 12.3-1 shows the most highly stressed catwalk and menorail support pad plate attachment welds to the torus shell. The load cc=binations for which the welds are evaluated are presented in PUAR Table 4-2.2-2. The welds are evaluated using the ASME Ccde criteria contained in Subsection NE for Class MC ccmponents. As can be seen frcm the attached table, the internal structure attachment weld stresses are within allcwable limits. DET-15-020 35 Revision 0 nut.e_c_h.

l I l l l l Table 12.3-1 Catwalk and Menorail Support Pad Plate Weld Stresses Item Calculated Stress Allcwable Stress Calculated (ksi) (ksi) Allowable Catwalk 3.46 15 01 0.23 Pad Plate Weld Monorail 0.49 15.01 0.03 Pad Plate Weld 1 1 DET-15-020 36 00 Revision 0 - _ _ - _ _ -}}