NL-08-0869, Vogtle, Units 1 and 2 - Enclosure 8, Westinghouse Electric Company LLC, LTR-CDME-08-043 NP-Attachment, Response to NRC Request for Additional Information Relating to LTR-CDME-08-11 NP-Attachment, Dated March 18, 2008
Text
Vogtle Electric Generating Plant Units 1 and License Amendment Request to Revise Technical Specification Sections 5.5.9, "Steam Generator (SG) Program" and TS "Steam Generator Tube Inspection Report" for Interim Alternate Repair Enclosure Westinghouse Electric Company LLC, L TR-CDME-08-043 NP-Attachment, "Response NRC Request for Additional Information Relating to L TR-CDME-08-11 dated March 18, WESTINGHOUSE NON-PROPRIETARY CLASS LTR*CDME..QS
....3 NP-Attachrnent Response to NRC Request for Additional Information Relating to LTR-eDME-08-11 NP-Attacbment March IS. 200S Westin.house Electric Company P.O. Box MadilOll, PA C 2008 Westinghouse Electric Company All Ri,hlS QUESTIONS RELAl1NG TO STEAM GENERATOR AMENDMENT ON INTERIM ALTERNATE REPAIR The NRC has provided to Wolf Creek Nuclear Operating Corporation (WCNOC) by email dated February 28. 2008 the Request for AdditiOll11 Information (RAJ) to aD interim alternate repair criterion (IARq that requires full-leogth inspection of the steam generator tubes within the tubesheet.
but does not require plugjng tubes if the extent of any cin:umfereadal ClICking observed in that greater than 17 inches from the top of the tubesheet that meets the performanc:e criteria of NEI 97-Q6, Rev. 2. "Steam Generator Program Guidelines," (Reference 1). A total of thirteen RAJ were provided to WCNOC. Four additional RAJ have siDee been provided to Southern Nuclear Operating Company for Vogtle Units 1 and 2. The same four additional RAJ were also provided to Bxelon Generation Company for the Braidwood Nuclear Power Station. The responses to RAJ 6 through 17 are provided below. After adjusting for growth as documented in Refereace
- 2. the allowable cnck sizes in the tube (20n and the weld metal (94' II'C bounding values and they apply for Model DS. Model F. Model 44F and Model S IF steam generators.
The 1.0 inch axial separatioo criterion discussed herein for multiple circumferential cracks also applies to these same model steam geoerators.
The ASME Code stress report relUlts summarized in response to RAJ 9 apply to the Model F steam generator only; however. it has been confirmed that similar resuhs have beea obtained for the Model DS steam generators.
3-7 (LTR*CDME*08-11-P) nuds to provith aU geometry details Q$sumtld in tM weld analysis on pages 7, 9 and 10. fl1re NRC staff does nor wulentDnd the tUSll1Mil weld geomnry based on tM discrusion on 18'$ 7, 9 and 10.) With nSp<<1 to the equation for S.A. near the top of page 10, wltGt is the pa1'tIIfI4lf!r whose WJlue is 0.020 and whIIt is the solillion for "y"'! Response:
The tubc-to-tubesheet weld is modeled in Figure 6-1 below. The tube wall has an inner radius rj and an outer OOius r.. and it is displaced upward [ J FJpre6-1 The equation of a line. relative to the ellipse is: y =mx + b. where 4 the slope = lane. and one point is localed at (r** 0.020). The resulting equation for the line on which the crack grows is: [ "C.t ] Similarly, the equati on of the eUipse. as offset from the origin. is: [ a,c,t ] where [ Simultaneously solving the equations for the line and the ellipse results in the point of their intersection (x. y): Settiog the points so that they are now relative to the original coordinate system gives the point (x'. y'). [ ] Lc,e The surface area of the frustum. S.A., is calculated by the surfaces of revolution technique and is a,C.t [where, the equation for the line can be rewritten U: a,C.e [ s and Thus, d:.c-==cot8 dy a.c:.e [] and the mult is: [] The previous calculation made use of surfaces of revolution (ct varies from 0 to 2*11:) in order to c:a1cuIale the sudKe area of the entire frustum. Now, since the cin:umferential flaw does not subtend a surface completely around the frustum, the equalion must be intepJed over an angle of revolution (ct to In Iddition, as the crack grows along the line of crack propaption.
the y-value is intepated from y' to y'+d*sine, where d is the crack depth. Thus. in this cue, the surface area of the flaw. Ar.is: a.c,e [] the final result of which ..[ ] The suIface area of the circumferential flaw. Arc. is a hybrid of the previous two. The angle of revolution again varies from 0 to 2 K, as in the case of the swface area of the frustum. HoweWI', the y-value varies from y' to y*+d*sin9.
just as in the case of the partially circumferential flaw. Now the integral is: ..t,e [] and the result is: ..c.e [] 6 On paBe 10, the lUs"nwa flaw is said to extend lJ aisttJllCe "a" into this ",urface." Does "surface" ,II/e, to the ollie' ellipse 0' inll4,ellip.
in FiB",e 3-'? FiBure 3-' sUBBeslS it i,jrom the i1UU!' ellipse. Response:
Referring to the frustum pictured in Figure 3-4 on Page 16 ofLTR-CDME-08*11, viewing the frustum from above (loo1cing down) or viewing the frustum from below (looking up), the view obWned is shown in Figure 3*S. The Cl'llCk originates in the bottom of the frustum in Figure 34 and grows upward along the surface depicted.
That is what the crack in Figure 3-S is attemptiDI to show. The crack originates at the point (x', y') in die filSt figure provided to answer Question 6. What was the assumed flow st,ess fo' the weld mDterial?
What was the basis fo, .,.eting this "alue? Response:
The weld is an autogenous weld; no filler metal is used. The flow assUDlll:d for the weld bead is the same IS that of the tube (base) metal, which was taken from Westinghouse WCAP-I2S22 (Refereuce 3). This is a conservative assumption since the Alloy 182 weld metal used for the tubesheet clad is stroogerthID the base metal of the tubing. Manufacturer's specjficalioos l for Alloy 182 and Alloy 82 weld metal indicate that the yield strength rIDges from [ ]II,C.. and the ultimate tensile strength ranges from [ ]&.U The flow stress (O.s*(Sv+SUT>>
then ranaes from [ ]IU This range of values is higher than the flow stress used in the tube ligament lIDIlysis
[ ]1,C,8 LTR*CDME-o'-209-P (Rqere1U:e
- 5) stale, tMt the tube-to-tubesheet welds we,e cksisned DIId onalyUd as primary press",e bountUIry in tlCcordimce with the ,equi,eme1lt8 of Section 111 oj the ASME Code. PFOIIide a lunurrmy oj the Code OIUIlysis.
illclutlinl the calclllatea maximwn sires, and applicable Code stress limit. Response:
GeperaJ Summery of ASME Code Stresa Report Besuks aelaliye to the lARC The existing Model F steam generator tube end weld (TEW) analysis used an axisymmetric finite element model (FEM) to estimate the 8b'CSS state of the weld material.
The assumptions in the weld analysis (Reference
- 2) closely resemble the assumptions in the IARC (LTR-CDME-08.11-P).
For example, in the Model F FEM analysis there is [ ]L." 'Ibis result is similar to the [ ]a.c. plane cited in LTR-CDMB-08-11-P when the different weld surfaces are compared (i.e., the flat plane chosen in the Model F FEM geometry versus the elliptical plane used in LTR-CDMB-08-1l-P).
Therefore, the results described for the limiring weld ligament in OS-ll-P are reasonable.
In addition.
the stress results contained in WNET-153, Vol. 6 (Reference
- 6) for a I FAX from Samuel D. ICailC1'.
P.E.,ofInc:o Alloys lnl'I.Inc.
We1diDI Products Co. da18d August 31, 15199 to Karan K. Oupca ofWestinlhouse NEE-Pensacola.
7 Model DS steam generator are bounded by those contained in the Model F steam generator report (Reference 4). Weld Geometry Model Figure 9-1 shows the configuration of Ihe weld as modeled in the Code stress aualysis.
This is a conservative idealization of the actual weld bead. whicb is approximately an [ ]o,c.c The inlerfacing elements to the weld have beeIIldded to Figure 9-1 for clarity. II,C,C The average actual height of the weld bead wu delennined by destructive examination of 10 faClOl'y welds and was found to be [ r" The modeJed heipt of the weld was conservatively sel at[ rc.e To maximize the load applied to the weld, since the dominant loading is tubesheet deflection, a "stiff" tube of [ rAe wall waS assumed. Stress The resules of the stress analysis are contained in Table 9-1 for the limiting section of weld 8 T.ble9-1 Test Note: Pili is the primary membrane stress intensity The design primuy membrane stress intensity is based on the design pressure differential of ( ]",c,l and an iSOlhermaltemperature of [ lOU from the Equipment Specification.
r ned' and Loading Conditions Tbere lR four soun:es of applied loads on the weld material: Deformation imposed by the tubesheel motion (taken at the center of the tUbeaheel.
assuming no resaraint from the divider p1ale, to maximize the tubesbeet deflection).
This is the most significant of the loads.
pressure differences. Local temperature gradients.
Shown to be "trivial" in the Code stress analysis. IsOlhennal temperature.
Local temperature gradients lR very small. (Exception:
Non-ductile failure evaJuation.)
Weld residual stress is not considered because it is stated to be inlrigDific:ant compared to the operating loads. This is because the ASME Code stress report analysis assumes that there is [ The end cap loads and fatigue results for the tube end weld were evaluated for several ASME Code defined conditions lIS specified in the Equipment Specification for the Model F steam generator.
The condilions in the analysis included: Design Condition Normat and Upset Conditions Emergency Conditions Fautted Conditions Test Conditions 9
Material The materials used in the PEA model
- Tubesheet Ligament:
SA-SOB CI 2a
- Tube: SB-163 (Code Case 1484)
- Tubesheet Cladding:
Inconcl Weld See the tables below for a detailed description of the appropriate data from the applicable Code year. 10 --------------_._
.._-_ ..
TAlLa 4-1 HAmlAL PlOPBl\T1U VS. nMPB8ATl'U
'01 SA-508-CL.
- 2. r=.. ._--.--TC TO a x 10 6 E x 10-6 a-Sy Su (atu/hr-ft--') (ft 2/br) (lnlla--') (pd) (kal) (k8l) (ksl) ;--._-----------.-, a.cte ...... TC -
Conductivity TD -Thee..l Diffu81vlty a _ H.an CoefUcieat of BxpanslClft ,olna frOll 70-' to todicateel t.,.rature. E -Hodulus of 11**ticlt, S. Des. Str** Iat**lty 8y
- Yield Strensth Su
- Ultl..te Strenath
'IOU IfATIlIAL PIOPBIlTIIS va. TDl'QATl'1I 1'01 S1-16:! (COde Clle ......__* _0 " Syax 10'TD S. (\tat)(kli)(in/:ln--F)(ft 2/hr)I tt B x 10-6 Su (Btu!br-ft-*') (pal) (ItsU ; Ite,e !'J-Ie
- TIlerlllll1 ClJIIduc:tivity TP
- Ther_1 Dlffusivtt7 Cl *Hun CDefficient of bpandon loinl fr_ 70-' to inelieated teMperature.
P. Hodulu. of Blut:Lclty 2 s.
- Deai;D S tr... lnt..at ty S7 -Yield Strenatb Su
- DItt_t. Stree.t..
The thermal properties and the elastic modulus of the cladding are assumed to be the same as those for lhe tube. Tbcrma1 Analysis The thennal analysis considered a boundinl transient for Normal and Upset conditions.
Inadvertent RCS Depressurization.
For this transient.
the muimum calculated differeace between the nodes repIeRnted in tbe FHA model is [ ]II,U! It was concluded that the [ Method of Analysis The analysis was performed with an axisymmetric fmite element analysis in tile WECAN computer program with a very fiue nodal mesh in the weld area and its interfaces with the tube and the tubesheet clld. The elements consisted of [ ]"",,c Applied 10MIs were due to deformation imposed by the tubesheet
- motion, secondary pressure differences.
local temperature gnutients, and isothermal CalculaJed S!Jlw, The foJlowing tables arc reproductions of the tables included ill the code stress IDIlysis for the tube end weld. Table 7-5 shows that the [ ]a.c.c The section numbers in Table 7-5 correspond to the section numbers in the model description figure above. In order to demonstrate acceptability, [ jl,C.I 13
u c.i* .;......... . C".... ... I':.... .lit:;' Jl i! ..*M ...
- 111 ,*,.. ... .. III*... II ! 8 ... =*... I..I ... .. ... I ...... I... ! .;= t.. GIl i I ...*I I Ii.. i t; Ii ..II GIl I! I ".I N us .!f .... CIIIJ D :l.......*...,; .. 1-0;... ........C".. OIl .. "'t/, 13..:2.5 :.-... *I..... *!.l! ...*.... ..II: I*M..*B .. S! II I.. ::" ......**u ....g..*..14
,;... 1:2;I;;!e C"...,. Ii ..... . 'Ii.. : :loS .. J ... VI I II , .. ... .. -.......... u III ... ..C/ll..... = .-4 .. J*:01 OJ J:. *! ... uo UI ... .. D:..... ... t: D... u K... .;"'!i..a=o'I.':
- =8 *u.. ....I*.. Ii....l:; .. .I ... .* I 0. II.. ...... I... ...... ", .. *:0: .:I*..:If I.. N *" 0 lOt' *D Zlr.. .. Ii.... ..u VI 15 TABU 7-5 AlII) war Pll1HAKt PUJS RCOIIDU! SDISS IftIRsm Lacat10ll
- eP L + 'b + CUi) Allowable L1II1 t (:1> Lc.,e Section nUmbers ale identified in the figure included with tbe Weld Geometry Description, above. AJI transients creetiag primary*plus-sccondary stress intensity ranges greater than 38m are evaluated inelastically.
16 Summary of Fatigue Usage from Code Stress Analysis of the Tube End Weld: lLC,e The point of maximum usa. factor, where [ ]...... is the most likely fatip Cl'1lCk initiation point. although the usap is still less than 1.0. Non-Ductjle Failure BValuation The nzthods of evaluating non-duetile fililure are [ 10. Regarding the weld repair criterion: A detlJiled IInll Q1I/lly&i.f (e.g., finite elmrent) would be ...cted to reveal a much compla Itrell It au than tlUll tulllmlltl in the licenue'l antdy&il.
which may impIJct tM llUly loctlliotLs for crack initiation tmd dinletfon of cracl: propagation.
In addbion, the dornintmt Itrell" for crack initl4lion and crack growth m4y inl10lw residwJl Itru,el in atldition to operDlional Itnnel. Abo, fltrwl may Ittm bftn introduced dlUinl weld fabrictJlion.
Thu, the 35-tlegree conical "pltme" i& not 1M only pitIM within which cracks may initiate Dnd grow. OM 1rypotherkal crack plDM, which tlJ'1HlIrs more limiting than the one tl$lumed by the liceruee, is the cylindrical "plane" UfiMtl by the upandetI tube oilier dU:uneter where tile weld is in a ltale of IMar. A&&uming a flow Itnl&! of 63.7 lsi and an qJ.ctive weld dqtIa of 0.035 ;IIChe, (tl$ IIrbwn in LTR-CDME-05*209*P, Figure 2.1), the NRC naJf elrimatel that the required.circuniferenrlalliglJlMnt to resist Dn end cap load of 16J71b is greaur than 180 tUgnl" (wWaout allowancel).
Addreu tMle collcern, and provide a tUta/led junijication for why the lubmitted alla]ylu is con.fervOlive.
17 Response:
Weld residual stress (WRS) was not considered since there is no definitive basis for any value used. Both the original Wolf Creek code stress analysis and a more recent code stress analysis for different models of steam generarors dismiss teaidual stresses in the weld as negligible.
Development of credible residual stresses using FBA methods is extIemely difficult, particularly for small welds like the tube-end weld. A comprehensive test propun involving deep/shaDow hole drilling, or finite element analyses which include the birthing of elements under very high to simulate the welding process would be required in order to develop a value for use. Verification of finite element WRS analysis results by deeptshaJlow hole drilling can only be acc:omplished for larger volumes of weld metal as removal of cores of trepanned material is required.
For small volumes of weld metal, verifICation of the finite element analysis is much more difficult and thus. the WRS values assumed are more uncertain.
In the ASME Code stress analyses, the operating loads on the weld are characterized as overshadowing any effects of WRS. Current development of residual suess models (lmpUblished) for consideration as a Code Case indicate dw the stress on the inner diameter of the tube is compressive, and not conducive to crack opening. The WRS values lJSed as the basis of the modeling were taken from the heat affected zone (HAZ) of stainless steeJ welds; therefore, the actual WRS profile may be differenL The profile is tensile in some lU'eU aod compressive in others (only tensile components of WRS have a deleterious effect). Consideration ofWRS further compliclres the analysis, but does not necessarily add any conservatism.
The weld region is not in a state of pure shear. 11lere are tensile loads as well as the pressure acline on the face of the weld exposed. to primary coolant. Therefare, the limits for pure shear (ASME BclPV Code Section m. NB-3227.2) are not considered to apply. Thus, the ASME code is satisfied with respect to pure shear. The shear plane used in the !ARC weld ligament calcula&ioD was OI'lly used to calculate the shear component of the stress state. This is consistent with the original Wolf Creek code stress analysis in which shear was nat explicitly considered, and the shear plane identifJed was not found to be the limiting plane. The most likely C1'8Ck iaidadon point, due to fatigue usage, was on a plane extending from the weld root almost normal to the face of the weld. A recent code stress analysis for another plant did consider pure shear explicitly and determined that the weld region is not in a state of pure shear, thus supporting the WCNOC stress analysis.
This report definitively stated that the pure shear limit of 3227.2 (O.6S,J does not apply. The crack opening performed in the weld region for the Wolf Creek IARC was assumed to open due to maximum principal stress, which is tensile, and flow stress was chosen as the limiting strength parameter.
While reviewing the Wolf Creek lARC repon, it was found that the component stresses, which genente the principal stresses, were not being recalculated as the flaw grew. The correction to this problem (see below), which is documented in Reference 7, changed the bounding requital remaining ligament for partially circumferentiaJ flaws in the weld region to [ ]II,U (not adjusting for growth) from the approximately
[ }8U origiaally reported in LTR.o>ME411 P-Attaehment (reference Table 3-3). The value of [ }o.c.e supersedes the old value of [ ]"'" Westinghouse believes that these corrections make the cOl'lSideratiOl'l of the flaw area in the left hand side of the farce balance equations correct. 18 The normal stress component was: LC,e [ ] The nonnal stress component now is: a,e.e [ ] The shear stress reported in the Wolf Creek IARe was: a,C,l: [ ] The shear sbeSS component, IDltil the flaw breaches the weld root is now: a,C,e b is the semi-minor axis (0.014 inch). This is due to the &hem-palh being uninterrupted until that point. After breaching the weld root, there is a lack of a sIre5S path. The shear stress at that ooint, is: 1,C,e J propOIed. and
db not interaction of
flaws which may 1M in proximiry .* lVCiol sqaratiOIl of or two UIbe Addrell this tlIUl identih tury nvisiOfU which may 1M rweMd to tM tube repair criteritJ mul the maximum acceptable weld flaw Response:
In order to ascertain how far apart cracks must be in order to be considered to respond independently to an applied rar field stress. a fracture mechanics appl'OllCh was undertaken.
The assumed case was [ 19 lOU Therefore.
a conservative estimate of me distanc:e necessary to prevent the interac:tiOD between c:rac:ks is [ JIU and is equal to 1.0 inch. Il is also worthy to note that 1.0 inc:h. whic:h is between 1 and 2 tube diameters, bounds the 0.5 inch rault contained in the ASME Boiler and Pressure Vessel Code.Section XI. Attic:le IW A-JOOO. "c,e Figure 11*1. ladi'fidual Steam Geaentor Results lor tile DIMaac:e Nec:e8lary for ay, to Equal (f
..e.e npre 11-2. Comblaed StuDt Generator Raultl for the DIItaace for 0" to Equ. (J The impact of Ihe crack separation analysis is summarized below. Refer to Figura 11-3 through 11-5 for explanations of the crack geometries aod combinations of crack-Ub indications considered in the aoalysis.
Table 11-1 is a summary of the text description of the crack separalion analysis impacts. The details described in Table 11-1 epply only to the portion of the tube within the tubesheet 17 inches below the top of the tubesheet (1TS-17 inches). An Indusay Peer Review was conducted on March 12,2008 at the Westinghouse WaltzMiU Site with the PUIpOSe of Mviewing the Fall 'JJXJ7 Catawba Unit 2 cold leg tube end indications to establish whether the repcIted indications are in the tube material or the weld material.
A COOseDSUS was Mached that the 2fX11 Catawba Unit 2 cold leg indications most likely exist within the tube naterial.
However, some of the indications extend close enough to the tube end that the possibility that the flaws do e",rend into the weld could not be ruled out. Therefore, in order to address the potential for cracking in the tube weld in parallel to crack-like indicatioD8 in the tube. the more limiting ligament size of [ lOU (including the adjusunent for growth) for the weld is used to esaablish the allowable crack size in the tube for cracks less than 1.0 from the tube end. Crack-like indicationi in a tube: If any circumfeMntial crack-like indication in the tube exceeds 203", plug the tube. If there is more than one circumferential crack-lie indication in a tube, and no single crack angle ex.ceeds 2030, and the minimum axial distance of separation between the crack-like indications is 21 greater than or equal to 1.00 inch, then the maxinwm crack angle is used to describe the flaw and the tube remains in service. If there is more tban one circumferential crack-like indication in a tube, and no single crack angle exceeds 203-, and abe minimum axial distance of separation between the crack-like indications is less than 1.00 inch. and the non-overlapping sum of the crack angles plus the overlapped crllCk angle is less than or equal to 203 0 , the tube may remain in service. If there is more thaD one circumferential crack-like indication in a tube, and no single cnlCk angle exceeds 203-, and the minimum axial distaDc:c of separation between the crack-like indicatioas is less than 1.00 inch, and the non-overtapping sum of the crack angles plus the overlapped crack angle is greater than 203°, plug the tube. Crack-like indications in a tube less than 1.0 inch from the tube end: If there an! one or more cracks in the tube that an! each less than or equal to 94-, and there is a minimum axial separatiOD dilllmlCe between the tube end and the tube cracks of less than 1.00 inch, and the non-overlapping sum of the tube crack &Dgles plus the overlapped crack angle is less than or equal to 94 0 , the tube may remain in service. If there is a crack-like indication in the weld less than or equal to 94° and there are one or more cracks in the tube that are each less than or equal to 94°, and there is a minimum axial separation distance between the tube end and the tube cracks of less than 1.00 inch, and the non-overJapping sum of the tube crKk angles plus the overlapped crack angle is gruter than 94°, plug the tube.
Table 11-1: SummarY of Crack SeDantion AIUlIYsIs and Interadlonl Multiple Cracks? Max. Crack Angle in Tube fP* Max. Crack Angle in Weld J* a Min. SeparatIon Distance.
L Required Action Case
Dearees (0) Degrees CO) inch 1 No > 203 I No Crack I Nt A Plug Tube 2 Yes 8,
.6IJ.6.
S 203 No Crack Cracks do not inlerlct.
Report max. crack angle less than 203°. Leave in Servic:e.
3 Yes 8,+6IJ+t.t.
oS 203 No Crack <1.00 Sum of total non-overlapping cnck angle plus overlap an2Jc less than 2OJO. LeaVe in Servic:e.
4 Yes
> 203 No Crack <1.00 Sum of total non-overlapping crack angle plus overlap auJe areater than 20JO. Plul Tube. 5
- Yes 94 Possible Crock in Weld < 1.00' Sum of total non-overiapping crack angle plus overlap angle Jess than 94°. Cracks in weld and tube do Intel1lCt.
Leave in Servic:e.
6 Yes 8,+6]+4+<<>
94 Possible Crack in Weld < 1.00' Sum of toeal non-overlapplng crack angle plus overlap angle greater than 94°. Cracks in weld and lUbe do interact.
P1uJt Tube. 1. See Figures Il-3. 11-4 and 11-5 for tube crack angle and weld crack angle definition.
- 2. 6. is the sum of any remaining crack angles the first two c:ntek-lilte indications.
For example. the statement:
9.r oS 203° is equivalent to writing: + 9J + 8J +... 203°. 3. Separation dislanc:e.
L. is measured from the tube end. 23 Fipn 11-3: Tube CrKk Gee.etry Jl1pre 11-4: Tube.nd Weld Crack AqIe Meuunment L I Tube, (J Weld, a Fipre 11-5: Axial Sepanllon DlsCaace Between Weld and Tube Crack-Ute IDdluCioDS The technical support docummt for the interim ARC amendmenl dtNs not make it dear hoM' licensees will elllure they S(llufy tM t1Ccitk,., induced lealuJle pelformtulCe criteria.
Describe til.
to be u.red to the accident leakage peifomronce criteria is rMt. Include in this (a) huw I"akage from sources other than tM lower 4*incMs of the rube wUl be oddress,d (in tM context of "nslU'ing tM performtulCe critflrln is met), and (b) 1u1w leaJcagefromflaws (if CIIIJ) in tM lower4-ilu:Ms oJtM tube will be thtermined
(,.g-, thtermining th"leakagejrom
<<IClt.jlaw; nwltipl,ing 1M normal operating I'd rat, by a specificJtlCtor).
Response:
The Modified B*1eakage analysis in the IARC Ieport calculates the ratio of undegraded Clevice length determined by eddy current inspection to die length of undegraded crevice required to meet the design basis accident analysis primary-to-sccondlry lealcage analysis assumption for the limiting design basis accident.
By definition of the IARC. 17 inches from the top of the tubesheet is the available undegraded crevice length because conflI1Ded CJ'IICking in this length will require the tube to be plugged. Both the pressure difference ratio and the length of crevice during normal operating and design basis accident are factored in the margin determination.
Refening to Table 4-5 of the IARC Ieport. the limiting design basis accident for WOOS is a postulalCd steam line break (SLB) event. Referring to Table 4-2 of the IARC report. it is calculated that { ]1oU of undegnded crevice length is required to preclude exceeding the SLB accident lIJIa1ysis leak rate assumption of 0.25 gpDL This conesponds to a safety factor of approximately
[ ]a.c.. in terms of the ratio of non-dcgraded crevice as confinned by eddy cunalt inspection (17 inches) to the crevice length calculated using the D'AIcy equation necesury to preclude exceeding the SLB accident analysis leakage assumption
[ ]o,c.. Therefcn.
the maximum leakage rafe that would OCCID' during a postulated SLB event from cracks occurring 17 inches below the top of the tubcsheet is calculated to be [ ]o,c.. from the faulted SO. This provides a margin of [ F on leakage rate for other soun:es of accident-iocluced leakage. The table below shows the available margin for leakage sources other than the tubcsheet based on the IARC method for calculating the estimated leakage for which a bounding zero.conuct-pressum
...alue of loss coefficient.
based on the a...aiJable test data. is used. 25 Table 12-1: CalcalatioD of A"rlIIIable MaraiD for Leakap Soarc:a Other ......D lu the Tubelbeet
])grille the UmltIna PIaat DesIp BMis AceldeDt (DBA) Plant NOP Leak limiting Plant DBA Leak Limit LRequired farDDA Safety Margin DBA Leak Margin Available Limit DBA I,C, e -The to Question 13 (following) further clarifies the methodology for satisfying the induced leakage performance criteria.
For the underlying assumptions of the IARC -no contact pressure between the lUbe and the tubeahcct in the hydraulic cxpansion region -the diSCUS$ion above shows that significant margins exist over the length of the crevice required in the 17 inch speD below the top of the tubesheet.
However, a conservative factor of 2.5 will be applied to that part of the observed normal operating leakage that caBnOl be associated with degradation mechanisms outside the tubesheet expansion region to calculate the accident-induced leakqe from the lUbesbcet region. The resulting calculated accident-induced leakage will be added to the predicted leakage from other degradation mechanisms that have been detected in the: 50s that have the potential to result in accident-induced leakage for evaluation against the accident-induced leakage performance criteria.
J 3. The proposed "modified B*" approach relies to S01M utellt on an IUSumed, con.rumt Vallie of 10$$ coefficient, btued 011 a 10tW!r hound of the data. 71U.r colltrrutl with the "nomiMl B." approach wlUch. in il$ltItutform (IU we understantl it) is no, directly impaaed by 1M IUIwned value of losl coeJliciellt since this vtJlue is IUlunwa to 1H C01l$kInt with increlUUaB colltaet pre$$ure lMtween the ruM tlnd .,heet. Given the tJmOlIIIl of timI for the NRC stqff to review the interim ARC. the NRC Ittljfwill not be able to mtJlce a conclwion tIS to whether 1M tUSIII1Ied value of 10$$ coeJficienz in the "modified B*" approach is cOlUerlltlliw.
However, the NRC staff Juu performed 101M eval/ltltiOlls regarding the potenliDI for the normal opertJtinr ktJk rau to increase II1IIkr steam line break conditions wing various Wlluel oj(INO,IlsuJ utermin4iijrom the "nominal B*" approach (which doel not rely on an tUsumed lIQlue of losl coefficient).
With 26
and ncognjzing tIIsoci4Ied with S01M oj previous H*/B*
it would that a factor of 2.5 nasonably bounds 1M
in tMI would in goinB from normal to
conditions.
Discuss your plans to modify your proposal to that tM kak tblring normal opertJlion Uor j/Ilws in 1M 4-incMs oj tube) will incntlU by a factor of 2.5 under
condilions.
rIb NRC stJl/f makes two oburvaDoru in to indlUtry
- 11. First. tM NRC staff tJrat 1M ratio oj tM allowed tlCcitknt and the
is only 2.5 for Wolf which is to 1M factor of 2.5 (TJaU ratio is 3.5 for Vogtle and 5 for ByrorrlBraidwood).
ThU is not an atypical sinuJtion as is discrused in NRC RIS 2()(J7-20.
17w limit in 1M If:chnicalsp<<ijications can neHr tllllAmed to tMt accUknt will be within what is tIIsllmMl.
in tM acdtUnt analysis.
f:Wn if tM limit is zero. For uomple. part through wall flaws in tM span which an not
normal conditions may pop through wall and l.ale accidnat conditions.
For cracu in 1M span which an leaking under nomtDl conditions. ratio of SLB to normal operating can substDntilJUy tluua 2.5 tkpending on tM oflM crack. It is tM re.rponsibilily to mrun that 1M occident limit.r tlT'f: 1Mt through implmtSltation oj an eff-etiN SG program. i,u:luding an tusellmDIt of cury
that may occur in oj iu implications for
.....r DCcilknt conditions (based on considerations such til past iMp<<:tkJn rflSWU tmil operational twe.s.mumts.
Itlsimilar
- plants, Second,
- NRC stqff is not of any to datI: from tM region for tM cltus of planu. and there Unle to that this .ritUltlion wiU cJumge significantly in 1M nf:JCt 18 months. Thw. tM NRC stairs appro<<h discw.d is 1I0t to any signijicant impactfor 1M
nli4ffrom tlal: tubl: repair criteria in tM lower 4-incMs oj tM tube.} Response:
The proposed ratio of 2.5 of the SLB to NOP leakage is conservative from the perspcc:tive of pedicled SLB leak rate from a postulated flaw below TIS-17 incbes based on the analysis below. Based on the D'Arcy Model for flow in an axial porous medium, if no value for loss coefficient is usmned, the incrase in predicted leakage from the tubesheet region would be lower than thai determined by using a flCtor of 2.5 and also than that provided in the !ARC justificalion.
For example, usume that both the loss coefficiem aDd the length of porous medium surrounding a tube above a postulaled crack are constaI'lt during both normal operating (NOP) and stellD line break (SLB) conditions.
The crevice below the neutral axis of the tubesheet will be tighter during accident conditions even if no aedit is taken for thermal lockup between the tube and the tubesbect due to increased pressure differential across lhe tube. If the pressure differential across lhe tube at SLB conditions is discoumed, the resulting condition is still an increase in cont8Ct pressure due to structural deflections and roIations.
Thus, there is no basis to assume a lower loss coefficient It SLB condition than at NOP condition.
Further. lhe viscosity during a SLB acciclent would be higher, due to the reduced temperatures in the crevice. Therefore, the assumption of a constant value for loss coefficient is, in fact. the worst case, and 27 is reasonable and conservative for the IARC because the flow resistance is expec:ted to increase during B postulated SLB event below 17 inches from the top of tile tubesheet.
Following the assumptions described in Question 13 (above), the D'Arcy Model becomes: Q= 6p R R=pKl 1CfIlQP =KsL.... K I NClP ... l sLa '" 17 in "'I This Ulumption forces the estimated increase in leakase to be a factor based on the ratio of differential preasures and the ratio of the applicable viscosities only. For the Wolf Creek S1eaID generators, the viscosity of the fluid dming NOP conditions is approximately 1.75xlO" Ibf-seclin 2 and during SLB is approximately 2.66xlO" Ibf-seclin 2* The pressure djfferential (AI)
- p..1 -Pas:> for Wolf Creek during NOP is 1443 psig and the pressure differential during SLB is 2560 psig. Substitution of these values into the D'key Model gives. 2560 Q u = Ie 9.624e81 Kl 2.66e -6(Kl) 1443 QN01' ==8.245l!81 KI 1.75l! -6(Kl) Qu =
Xl =9.624l!8!
= QNOl' 8.24Se8 Kl 8.245d I Using the D'Arcy Model to calculate the estimated increase in leakage during SLB yields a result of approximately 1.17. This is less thin the conservative utios wlUch range from 2 to 6 as reported in the tARC descripdon and the 2.5 factor propoaed by the NRC staff. For integrity assessments, the ratio of 2.5 will be used in the completion of both the condition monitoring (eM) and operational assessment (OA) upon implementation of the !ARC. For ex8Il1ple, for the CM assessment, the coqxment of leakage from the lower 4 inches for the most limiting steam generator durinS the prior cycle of operation will be multiplied by a factor of 2.5 and added to the total leakage from any other source and compared to the allowable accident analysis leakage assumption.
For the OA. the difference in leakage from the allowable limit during the limiting design buis accident minus the leakage from the other sources wiD be divided by 2.5 and compared to the observed leakage. An administrative limit wiIJ be estabUshed to not exceed the calculated value. It is not planned to modify the existing !ARC report. but, as noted above. a constant multiplier of 2.5 wiD be used in CM and OA evaluations to calculate SLB leakage from the lower 4 inches. 28 The mathemmical constDnIll' has omitt<<1from the fint oftM Mar the top of fHJ84 8 and the at th, bouom
- 8. It i.r not ckar if this i.r a typographical error, or if fr 1uu bun purpo.reftJly 1/tM omission is uplDin. Response:
Two typographical errors have been identified in the left band side of the equalioas for force ba1aDce for the partial circumfaential flaw in the steam generator tube wall and the partially circumferential, wall flaw in the steam generator tube waD an Page 8 of LTR-CDME-oB-ll P-AtblChmeDt.
A factor of 1t was omitted in each equatiOD in the report but not in the actual calculations.
1be calculation results are not affected by the typographical errors. The wt tema of tM at the bottom of 8 includes the parelllheticaJ (r,}+r/).
The stoff that this .should be (r,,2-r;'J). It is not clear if this il a rypogrop1tical error, or if the radii are illlentiolJlJlly beinB SlUJllneti.
1/
pkase explain why the sqruued radii shmrld be SIUMIM and not sllbtmctd.
Response:
Westinghouse agrees that the plus sign (+) should indeed be a minus sign (.). The error is typographical and did not affect the calculations.
The last tenD in the force balance equation for the putially circumferential, through-wall flaw in the steam generator tube conWDS a a x (112) x (r.2+ r?) x 69 tenn on the right hand side oftbe equation.
That should read a x (lf2) x (r 0 2- x Explain why it I.r Mcelsary to subtract AI (tuft' of tM flaw) from SA. ($wface area of tM f'nutum) in tM first tema of tM balance 0'1 page 10. (1M SUfI! believes that this term should be Response:
The area of the flaw must be subtracted from the sunace area of the frustum when calculating the force balance because that area is no longer contiguous and c8llDOt react to the applied stress. In other words, the flaw area is no lODger available ro the principal stress, but, is instead loaded by the internal pressure.
- 17. Explain the of the matMlfJQtical constant P, preuure) rather than P (3JJP or 4800 psi) on the eqllDlions on pages 8 and 10. The uplanali01l on page I I is not ndficimt t.11Ul to the stqffto be incorner.
Response:
It remains Westinghouse's position that it is conservative and correct to use an intemaJ pressure of 2250 psi an the crack flank to calculate an acceptable remaining ligament for crack-Jike indications that may be present in the tube and weld. However, at the NRC staff's request, the allowable Jipment sizes for the tube and me weld were recaJculated assuming a 4800 psi pt'eSSure on the crack flank. The revised values for remaining ligament for the tube and the weld are [ ]..... (inchKling an adjuatment for growth) rapectivcly.
29 ------------------------_._---_.
__.
For completeness.
a sununary of the Westinghouse position on the justification for the use of an internal preslUte of mo psi is provided below. A SO tube is a thick-waH cylinder.
This is consistent with the ASMB Code stress analysis of the steam generator tubing. Roark (Reference
- 8) defines a thin-wall cylinder as a cylinder with an inside radius to thickness ratio (RIa) pater than 10. For the Model F tube. RIt =8.8. therefore.
the tube is considered a thick-wall cylinder.
Reference 9 provides the equation of axial stress in the thick wall cylinder as: Where P is an active extemalload (for tbis case = 0) PI is the internal pressure pz is the external pressUJe Q is die inside radius b is abe outside radius The second term in the equation.1i(.t 2 ) , goes to zero because the applied extemallOid in this case 1lb -a is zero. The equation is conservatively simplified by assuming the is nesJigible.
Making tbis assumption conservative since retainin, the tenn would reduce the wal calculated stress au' The equatioo is reduced to let PI equal the pteSSUle differential lip. This is consistent with the equation in example 11.2 of Reference
- 9. This eq..tioo. and the following limitat:ions.
are echoed in Roark (Reference
- 8) Table 13.5, Case l.b. The final equation for the calculation of stresS due to the end cap load becomes Calculation of the end cap laid using this form of the equation is inherently conservative.
The limitation of the equation for axial stress in the thick-waIl cylinder due to end cap load. and for the stress equltions in the cylinder, is that the section of interest is far removed from the end caps (RefereDCe 9). Consequently.
the stress in the degraded section of the cyliDder is increased by the reduced wall. but the end cap load remains constaDt.
Calculating the end cap load for the thick-wall cylinder using the 30 degraded wall thickness is equivalent to assuming that the wall thickness for the entire tube is the same as for the degraded local section. It is the Westinghouse position that the load on the crack flank should be calculated separately from the end cap load. This is based on the fact that the end cap load already takes into account any varialion in the cross section of the tube. The underlying usumption for the IARC is that all circumferentiaJ ClaCks detected are lOOCll through wall over the entin: indicated length. The Westinghouse crevice pressure test da1a (Refcmx:e (0) shows that the pressure in the crevice external to the tube in the immediate area of the penetration is the same as the intemal pressure; therefore.
there is no differential pres5un: at that location and 3Ap equals zero. The existins analysis conservatively applies the entin: prilDlJy side pn:sSUfe to the crack face. 1'hae is no operating ccmdition thal justifies using triple the primary pressun: differential on the crack face and the required safety by the ASME Code for this situation (classification u secondary stress) would imply a safety factor of 1.0 on any primary side presS1R. Finally, the calculated on the depaded section are compared to the flow stress wbich is very conservative for this situation.
The condition of inten:st is one of pure axial separatiClll under the usumption of the !ARC, i.e., no axial friction forces between the tube and the tubesheet, but the tubesheet is (RSent in close contact to prevent bending forces. For pure axial separation, it is appropriate to lIIIe the ultimate stralgth of the material, since no beading can occur and burst is not possible due to the constraiDt provided by the tubesheet.
31 Relerenas: NEI-97-06, Rev. 2, "Steam Generator Prop'am Guidelines," May 200s. LTR-eDME.Q8-11.
'1nterim Alternate Repair Criterion (ARC) for Cracks in the Lower Region of the Tubesheet Expansion Zone." January 31, 2008. WCAP-12S22, "lnconel Alloy 600 Tubing-Material Bunt and Strength Properties," January 1990. WNET-IBO (Proprietary), Volume II, Rev. O. "Model F Steam Generator Stress Report," Westinghouse Electric, Pittsburgh, PA, September 1980. L TR-eDME-OS-209-P."Steam Generator Tube Alternate Repair Criteria for the Portion of the Tube Within the Tubesheet at the Wolf Creek Generating Station," January 2006. WNET-lS3 (Proprietary).
Volume 6, Rev. 0, Model DS Steam Generator Stress Report," Westinghouse Electric.
Pittsburgh.
PA. December 1981. CN-COME-08-4, Rev. I, ..Structural Evaluation of the Minimum Circumferential Ligament Required as P.-t of the WCNOC !ARC," Man:h 2008. Roark's Formulas for Stress and Strain, Warren C. YOWIg and Richard G. Budynas, Seventh Edition. McGraw-Hili, 2002. Advanced Mechanics of Materials, Arthur P. Boresi and Richlrd J. Schmidt, Sixth Edition, John Wiley and Sons, 2003. 10. STD-MC-06-11-P, Rev. I, "Pressure Profile Measurements During Tube-to-Tubesheet Leakage Tests of Hydraulically Expanded Steam Genemor Tubing," August 30, 2007. 32