ML20002A774
| ML20002A774 | |
| Person / Time | |
|---|---|
| Site: | Waterford  |
| Issue date: | 11/14/1980 |
| From: | Maurin L LOUISIANA POWER & LIGHT CO. |
| To: | Schwencer A Office of Nuclear Reactor Regulation |
| References | |
| Q-3-A29, W3P80-0158, W3P80-158, NUDOCS 8011210511 | |
| Download: ML20002A774 (50) | |
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142 CELAAoNCE STREET MioOLE south P O W E R & L I G H T! P O Box 6008 . NEW ORLEANS. LOUIStANA 70174
. (So41 368-2345 utor:ES Sf STEM November 14, 1980 W3P80-0158 Q-3-A29
_. c El Mr. A. Schwencer :-
Division of Licensing $?
Licensing Branch No. 2 7._ 3 ;,
U. S. Nuclear Regulatory Co= mission ifp: ~
Wachington, D. C. 20555 254 y $
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SUBJECT:
Waterford 3 SES d 3 1 Docket No. 50-382 G ua si Response to FSAR Questien 110.1/
Asymmetric Loads
Dear Mr. Schwencer:
On October 1, 1980, at the draft SER Review Meeting between the NRC MEB Staff, MEP Consultants, LP&L Staff and LP&L's Consultants, LP&L'e consultants pre-sented the basis and the results of an evaluation of the various components /
structures listed in Enclosure 1 to Q. 110.1. At the conclusion of the meeting and as indicated in the " Action Items" of the meeting minutes (c'ated October 24, 1980), LP&L was to provide a written report of the evaluation.
Accordingly, the attached report contains the information presented at the meeting and responses to the MEB Staff / Consultants questions.
Based on the results of the evaluation contained in the enclosed report, LP&L has concluded that all the reactor system components / structures (except fuel assemblies which will be addressed under FSAR Q. 231.2) cited in Q. 110.1 can withstand the combination of loads due to a postulated, but highly un-likely, design basis LOCA and seismic events and that Waterford 3 can operate without undue risk to the health and safety of the public. Also, it is LP&L's judgement that an evaluation using more sophisticated techniques than that contained in the enclosed report would only quantify the safety margins that already exist based on the evaluation contained in the enclosed report. Hence, from a value-impact standpoint LP&L believes that further analysis using more sophisticated techniques is not warranted.
At the conclusion of the draf t SER Review Meeting, the MEB Staff provided oral acceptance of our proposal to assess the adequacy of fuel assemblies subjected 1 to LOCA loads resulting from the controlling break, i.e. a postualted break at l h00l .5 801121o 5g g ih G
Mr. A. Schwencer W3P80-0158 Page 2 the cold leg vessel inlet nozzle. The basis for selecting this break location is contained in the enclosed report. The Waterford 3 fuel assembly assessment will also be made on this basis.
LP&L requests that the staff provides a prompt review of this response to Q. 110.1 and indicate the staff's concurrence based on its review thereof. It is LP&L's intention to include the enclosed report as the final response to Q. 110.1 in the FSAR Amendment scheduled for December, 1980. NRC Staff comments received by December 1, 1980 will be included in this Amendment.
Yours very truly, a .
U J.J -m L. V. Maurin Project Director LVM/aWP/dde Attachment cc: W. M. Stevenson, E. Blake
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i RESPONSE TO FSAR QUESTION 110.1/ ASYMMETRIC LOADS i WATERM)RD STEAM ELECTRIC STATION UNIT 3 i
i DOCKET NO. 50-382 r i
- j. NOVEMBER, 1980 A
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l In February.1979, the NRC, in. Question 110.1, required LP&L to expand Section l ' 3.9.1.5 of the FSAR to more specifically address the consideration of asymmetric loads on reactor coclant system components and supports which could result from postulated reactor coolant pipe breaks within cavities located inside the
! containment. Enclosure 1, attached to the question, described the information required.
LP&L had tried unsuccessfully, on several past occasions, to obtain clarification of this particular question. The. clarification is warranted because of the
. following:
a) The probability of the type of pipe failures required to be assumed for these analyses is very low;
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b) the nature of the analyses required to perform the stated assessment can vary in sophistication to a significant degree so that the time and expenditures required for the performance of said assessment can also vary quite substantially.
c) LP&L has already modified the plant original design to severely limit the area of the breaks that could be postulated to occur in both the reactor and steam generator / pump compartment and has increased the load carrying capacity of the fuel spacer '
grids and guide tubes; and d) an analysis of the response of the primary system to break at locations specified in Enclosure I was performed and reported in Appendix 5.4A of the FSAR, and no formal feedback has been obtained to date from the Commission for that submittal.
On August 13, 1980, LP&L personnel and LP&L consultants met with the NRC-MEB staff to present LP&L's intended approach in responding to Question 110.1.
At this meeting, LP&L indicated that several of the components listed in Enclosure -1 to Q.110.1 have already been assessed by either specific considera-tion of asymmetric loads or in the design of the component / structure. LP&L further indicated that it intended, to the extent deemed technically feasible, to demonstrate the adequacy of the remaining components / structures by comparison type analyses with other similar plants which had been previously analyzed.
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Table 1 summarizes the status of the assessment of each component / structure requested in Question 110.1. This brief summary report further expands on the contents of Table 1 and describes the rationale and/or analytical approach employed in the assessment by comparison.
Assessment of Reactor Vessel, Vessel Supports, and Biological Shield Wall The analysis reported in Appendix 5.4A of the FSAR has demonstrated the adequacy of the reactor vessel supports to withstand the asymmetric loads resulting from the 350 in2 cold leg vessel inlet and 100 in2 hot leg vessel outlet guillotine breaks. The response of the vessel supports for o.ther breaks needed no evaluation since it is clear that the absence of external aysametric loads for those breaks, postulated to occur outside the reactor vessel cavity, would result in lower overall loads since (a) thrust loads remain the same or are less for cold leg and hot leg breaks respectively outside the reactor. cavity as they are for the cold leg and hot leg breaks inside the cavity, and (b) internal asymmetric loads are not significantly affected by the size of the break area alone but rather depend on the combination of the break area and the break opening time.
Sensitivity studies performed for instance, for cold leg breaks, have shown that a full area guillotine break, requiring approximately 30 msee to fully open, a 600 in2 guillotine break requiring 18 msec to open, and a 200 in2 guillotine break requiring 8 msee to open, result in internal asymmetric loads differing only by a few percent. Figure i shows the behavior of the break area developed as a function of time following an assumed instantaneously occurring severance of the cold leg pipe a the vessel inlet nozzle. This figure shows that to each break area there corresponds an opening time. Table 2 reports the opening time required to achieve the breaks postulated in the RCS of Waterford 3 which are also tabulated in Table 6.2-1 of the FSAR.
3 The same analyris has demonstrat.d of course the adequacy of the reactor pressure vesrel to withstand those loads. At the same time, the adequacy of the biological shield is insured by design since the wall has been designed to withstand loads in excess of those transmitted by the reactor vessel supports in combination with the asymmetric pressure loads across the wall as described in Section 6.2.1.2 of the FSAR, and other normal operation and seismic loads.
Assessment of Steam Generator and Pump Compartment Wall.
The adequacy of the compartment wall surrounding the reactor coolant pumps and steam generator is also assured by design. Refer to FSAR Sections 3.8.3.3.1 and 3.8.3.3.2 for ine structural design and Section 6.2.1.2 for the subcompartment analyses providing the asymmetric loads across the wall for the postulated breaks (see Table 6.2.-1_ for the breaks postulated in the compartments).
Assessment of Steam Generator. Reactor Coolant Pump and Their Supports The asymmetric pressure histories reported for the breaks postulated in FSAR Section 6.2.1.2 have been used to determine the asymmetric loads across the steam generators and reactor coolant pumps. These components have been designed to accept these loads. The adequacy of the steam generator and reactor coolant pump supports is also assured by design. Refer to FSAR Sections 3.6.2.3 and Appendix 3.6A for the design of those supports and restraints.
Assessment of Primary Coolant Piping The adequacy of the reactor coolant piping under LOCA conditions had been i
verified when piping rupture restraints limiting the break areas were backfitted into the original design (refer to FSAR Section 3.6.2.3 and Appendix 3.6A).
The Waterford 3 design, being very stallar to that of St. Lucie 1, and both i
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l i having seismic' loads which'are an-order of magnitude less than the allowable E loads, and the St.' Lucia'1.RCS piping having been demonstrated to be adequate j iby' plant specific analysis for combined LOCA and seismic loads, then it is
' udged that the Waterford 3 RCS. piping should also be adequate since its j
i expected stresses are lower than'those of St. Lucie 1 for the following reasons: -(a) design basis break areas in Waterford 3 are smaller than break areas of St.~ Lucie 1 at the corresponding locations, hence asymmetric -
effects are smaller, and (b) vessel motions, resulting from asymmetric loads
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within the-cavity in Waterford 3 are significantly smaller than the corresponding vessel motions in St. ~ Lucie l'(refer to Figures 2, 3, and 4). These two results stem from the.significant difference in the forces across-the vessel generated
{- by.the pressures in the reactor cavities. Table 3 compares the seismic moments i
j ~at points in the RCS main piping for St. Lucie 1 and Waterford 3. Figure 7
- shows the horizontal and-vertical forces acting across the vessel for various i plants with reactor cavities which are not necessarily very similar. The 2
force computed for-the Waterford 3 break area-of 350 in in the cold leg from the pressures reported in the FSAR Section 6.2.1.2 is seen to be considerably 4
1 less than that for the generic plant full break, and to essentially fall in line with the trend of lateral. force vs. break area shown in Figure 7.
' Assessment of Reactor' Internals
! Motions of the core barrel, the reactor vessel, as wellas the relative' motion
. :between core b'arrel and vessel end hydraulic loads, are the important parameters
' in assessing the adequacy-of~the reactor internals. The relative motion between i core barrel and' vessel-represents the_new element in the assessment since it had been ignored in the analyses reported in Section 3.9.2, which had utilized very conservative hydrodynamic asymmetric loads,'but had held the reactor vessel l
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' A comparison of the vessel motions figures 2 and 3) resulting from the most significant of the postulated breaks; namely, the break at the vessel inlet nozzle, obtained for Waterford 3 to those computed for_similar plants; i.e.,
l- the generic CE plant (Calvert Cliffs, Reference 1)) and the St. Lucie 1 plant shows that the Waterford 3 vessel deflections are significantly smaller. This is due in-large part to the fact that the break area in Waterford 3 is significantly smaller by design,'than the corresponding areas in the other two plants. Another.significant difference besides the smaller magnitude of the vessel deflections in Waterford 3 is that the period of oscillation is ;
longer. Corresponding accelerations are thus much lower.in Waterford 3 than the generic plant or St. Lucie.l. The vessel deflections of Waterford 3 are further compared with those calculated for a plant subjected to a cold leg nozzle inlet break area of 188 in2. The comparison shows similar displacement time histories in magnitude and frequency for the Waterford 3 and the plant with 188 in2 break area. The slightly larger displacement in the plant with the smaller break area resulted from the larger gaps between the reactor vessel nozzle pads and the supports. The break area at the reactor vessel outlet nozzle for Waterford 3 (100 in2) is slightly smaller than the corresponding break for the generic CE plant (135 in 2). It can be expected therefore that the motions of the reactor vessel resulting from that hot leg break will be similar in both amplitude and periodicity for both plants but that the Waterford 3 amplitudes will be somewhat lower. For the generic CE plant (Reference 1) vessel motions re- ing from the hot leg break at-the reactor vessel have been computed to be , a than 1/3 of those resulting from the cold leg break at the vessel.
Since corresponding cold leg break vessel motions for Waterford 3 are essentially 1/2 of the generic plant motions, the cold leg break inside the cavity in Waterford 3 will produce ves,sel motions which are approximately 30 percent
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6 larger than the motions produced by the hot leg break. The hot leg break outside the reactor cavity (600 in2) would result in 90 percent of the vessel motions produced by the hot leg break inside the cavity.
As indicated above, the vessel motions for Waterford 3 are less than half of the corresponding motions of the generic plant (see Figures 2 and 3) and have about the same frequency content initially but subsequently they exhibit a longer period. The internal asymmetric hydraulic forces, across the barrel, are essentially unaffected by the break size and can thus be considered nearly
.the same in both plants. Sensitivity studies have shown that internal asymmetric hydraulic loads only decrease about 4 percent when the break is reduced from full area opening in 30 msee to an ar'ea of 188 in2~1n opening in 8 msec. The relative motion between the vessel and the core barrel for Waterford 3 vill be slightly smaller than that for the generic plant on the basis of slightly smaller internal hydraulic loads and the absolute values of core barrel and in-phase vessel motions for Waterford 3 being about half that for the generic plant. The in-phase nature of the barrel and vessel motion is confirmed by Figure 5, applicable to the plant with a break area of 188 in 2, It can therefore be expected that reaction forces on the Waterford 3 internals will be less than the corresponding forces for a cold leg break in the CE reference plant; whereas, they will be nearly the same for hot leg breaks 1
since the break areas in both plants are approximately identical (100 square :
inches inside the cavity for Waterford 3 and 135 square inches for the generic plant). The hot leg break outside the cavity is larger (600 in ),2 however, it gives rise to no cavity pressure. Hence, a significant load had the break been inside the cavity, which would contribute to the motion of the vessel, is absent.
7 Since the internals had been analyzed..for-a full. hot leg break .
with a fixed vessel, and since the vessel motions in the absence of cavity pressure are expected to be less (about 90%) than corresponding motions when cavity pressures are present, the determining hot leg break would be the hot leg break within the cavity.
In the generic plant analysis it has been found that the cold leg break in the cavity results in more severe loading on the internals.
Table 6 ccmpares the internals for Waterford 3 with those of the generic plant.
As can be seen, the Waterfo'rd 3. design, with one ex7eption, unimportant for this evaluation, is capable of witustanding higher loads than the generic plant design. The additional margin has been computed by considering the relative capacity of each element to accommodate bending (both meridional and circumferential) hoop stresses, local and global shear, and buckling; thus assuming the minimum capacity irrespective of the loading condition.
A comparison of the Waterford 3 internals with the generic plant internals indicates that generally where relatively low margins to design values had been computed for the generic plant, the Waterford 3-internals have at least 11 percent more capacity, with much larger margins in most of the internals components. These
! added margins, coupled with the fact that loads are expected to be lower, as a result of the limited break areas, assure the acceptability of Waterford 3 internals under the combined LOCA and seismic loads.
The fuel alignment plate is the only component whose capacity in Waterford 3 is less than that for the generic plant. Stresses in that plate were not critical for the generic plant, it is expected that similar conclusions can also be drawn for-Waterford 3.
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C l-8 i Assessment of CEDMs Control rod insertion for break areas exceeding 0.5 sq. ft. is not required due
, to NSSS design. Hence, it is only necessary to demonstrate that the CEDMs retain their integrity.
. (Reference 1) and St. Lucia 1, with the exception that in the latter to plants, j all nozzles are of equal. length;'whereas Waterford 3 CEDM nozzles have unequal 4
length resulting in all CEDMs tops being at the same elevation. This difference causes the CEDMs -differing in length from those analyzed for the CE generic plant and St. Lucie 1 (referred to hereinafter as the short nozzle CEDMs) to respond to lower frequencies on a mode by mode comparison than those previously analyzed.
Table 4 lists the frequencies of the-first seven (7) modes for both the shortest and longest.Waterford 3 CEDMs.
In Waterford 3 the amplitude of the Reactor Vessel Head motions (which provide the excitation to the CEDMs) are one-half of those of the generic CE plant for the_ cold leg break and the amplitudes of the motions for the hot leg break are another 30 percent lower as erplained earlier. For the cold leg break, the dominant frequency of the driving force is about 9Hz. This'is sufficiently close to possible modal resonances for any length CEDMs to be concerned with
-amplification'of motion. For the generic plant and for St. Lucie 1, the dominant frequency of the driving motion is about 17 Hz. On a purely elastic basis,.long-term excitation of their CEDM (all short) could be expected to result in two-fold amplification of motion as measured at the center of mass.
Results of an elasto-plastic analysis revealed instead an amplification of approximately 25 percent. This is attributed.co plasticity in the nozzle as well as short time application of the input'octions. A similar behavior i
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. 9 can be expected for Waterford 3, except that a higher amplification would be predicted on an elastic basis (factor of three for the longest nozzle). The expected amplification from an elast-plastic analysis would thus be of the
! order of 85 percent. Since the amplitudes of the driving displacements in Waterford 3 are half of those of St. Lucie 1, the maximum moment that can be-expected for Waterford 3 is less than that computed for St. Lucie 1, which was 173,000 in-lbs., and is expected to be approximately 160,000 in-lbs.*
The ultimate moment that the Waterford 3 nozzle can carry is app.roximately i
356,000 in-lbs. and has been computed using the approach of Gerber (Reference 2),
with minimum ASME property values for the nozzle material. To assure that the pressure boundary is not violated, the criterion is used from the ASME Code, Appendix F, that the integrity of the pressure boundary is assured if the applied loads do not exceed 70 percent of the plastic instability load. Since the maxicum LOCA moment for cold leg breaks in Waterford 3 is well below this 70 percent criterion (250,000 in-lbs.), it can be stated that the Waterford 3 CEDMs are adequate for cold leg breaks. The marimum seismic moment that the
< nozzle sees is 104,000 in-lbs. The absolute addition of these moments would thus slightly exceed the 70 percent criterion. Since the absolute addition of-moments is a conservative manner of combining LOCA and seismic induced loads
! and typically the nozzle material exhibits more-strength than the minimum ASME values, it is expected that the CEDMs will withstand the combined load.
1
- based'on Dynamic Amplification Factor = where f is the driving 1 - C.f_)2 frequency and fn the modal fu frequency.
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10 For hot leg breaks, the dominant frequency of the driving force is about 16-17 Hz, hence, it is not near any resonance with Waterford 3 CEDMs and further it would excite higher modes than the cold leg break. Since the amplitude of the driving motionn for the hot leg break are even less than those for the cold leg break, the adequacy of the CEDMs for this break is assured.
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RCS Attached Piping Assessment For the ECCS lines, an assessment of the adequacy of the lines was initially made by comparing the routing of the Waterford 3 lines with that of the ECCS line which had been analyzed in St.-Lucie 1. That analysis had shown that some plasticity might occur in the vicinity of the first elbow after the nozzle. Further, it had shown that one can conservatively estimate the bending moment at that location by use of the simple equation M = 3 6 EI/L 2 where 6 is the displacement at the nozzle, M the maximum bending moment, L the length (the length of the pipe section between the nozzle and the elbow), E the Young's Modulus, and I the moment of inertia.
For instance, the maxtr.um bending stress computed by the above formula for St. Lucie 1 is 22,800 pai which compares very well with the 19,280 psi computed by time history elasto-plastic analysis -# that line. That analysis had also ,
I shown that everywhere else in the line stesses, in combination with seismic l l
stresses vere within allowable limits (for St. Lucie 1). I In Waterford 3, the maximum bending moment has been computed on the basis i
that nozzle displacements would be half or less than the St. Lucie 1_ nozzle l
[ displacements (since vesrel displacements are one-half or less). The maximum l
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11 bending moments computed for two ECCS lines; one a 12 inch and one a 14 inch nominal line, are 3.2 x 106 in-lbs. and 4.1 x 106 in-lbs., respectively and they occur at-the first elbow near the nozzle at the RCS pipe.
Because of plasticity, consideration to simultaneous seismic loads has been given as follows: the maximum bending moment due to LOCA and that due to e
seismic have been added absolutely and their sum has been compared to 70 percent of the ultimate capacity of the elbow to carry moment without collapse.
For the two lines examined, the total bending moment (LOCA and seismic) is less than 70 percent of the collapse moment. Hence, it can be concluded that the ECCS lines wilf retain their integrity and function during a simultaneous LOCA and seismic event. Appendix A provides additional details on the method ,
of comparative analyses used in this evaluation.
The calculated moments are compared with 70 percent of the maximum bending moment carrying capability of those elbows computed by the methodology of Gerber (Reference 2), which are 5.6 x 10 6 in-lbs. and 6.55 x 106 in-lbs., respectively for the 12 in. and 14 in, pipes scaled by the B Index 2
in the ASME Code, B =1.817.
2 This simpler method of computing the capability of the elbow to sustain an applied moment had been compared to a finite element analysis method employed in Reference 1, and was found to give excellent agreement. For instance, the simple method predicts maximum collapse moment for the elbow of an ECCS line in ;
1 St. Lucie 1 to be 5.94 x 106 in-lbs., whereas the finite element method of Reference.1,'for a virtually identical elbow computed it to be 5.5 x 106 in-lbs.
A question was raised during the draf t SER review meeting (between MEB Staff, MEB Consultants, LP&L's Staff and LP&L Consultants) in New York of October 1, 1980, as to whether the seismic supports of the RCS attached lines would be
12 able to withstand the reaction forces resulting from LOCA-induced motions.
This had not been analyzed for Waterford 3 since a simple analysis had been employed to determine the bending moment in the pipe. An elasto-plastic analysis hts been-performed since, on the basis of actual Waterford 3 vessel displacements which are assumed to adequately represent the ECCS nozzle displacements. ,
Table 5 lists the reaction loads computed for Waterford 3 for LOCA motions, together with a description of the type of support /restaints employed for the restraints. Seismic loads and thermal loads are also listed. The overall loads computed by summing thermal and the LOCA and seismic loads combined in SRSS fashion is compared to the capacity of the support / restraint to failure. From Table 5 it can be concluded that the supports / restraints are capable of withstanding loads resulting from a combination of postulated LOCA and seismic events.
This analysis, conducted for the 14 inch line only provides confirmation of the conservatism of the assessment of the first elbow bending moments by use of the simple formula given on page 10 . For this line the bending moment computed by elasto-plastic detailed analysis is 7.8 x 105 in-lbs. which can be compared with a calculated 4.1 x 10 6in-lbs. determined by the simple formula. Since the maximum bending moment conservatively calculated by the simple formula are considerably less than 70 percent of the ultimate carrying capability of the elbow, functiomdity of the ECCS lines is assured.
Assessment of Fuel Assemblies As indicated in Table 1 and at the draft SER review meeting with the MEB Staff /
Consultants on October 1, 1980 the adequacy of Waterford 3 fuel assemblies will
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I 13 be provided in LP&L's response to FSAR Q. 231.2, as Q. 231.2 specifically addresses the ability of fuel assemblies to withstand combined seismic and LOCA mechanical loads. However, as indicated at the October 1, 1980 meeting, the Waterford 3 fuel assembly evaluation for LOCA will be limited to the cold leg vessel inlet-break, since all prior analyses of fuel assemblies (generic plant, St. Lucie 1) have indicated that.the cold leg brcak at the vessel inlet nozzle produces maximum LOCA loads on the assemblies. Appendix B provides a basis for arriving at the above conclusion.
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References:
! (1) Calvert Cliffs Nuclear Power Plant Units Nos. 1 and 2, Dockat Nos.
50-317 and 50-318, Reactor Coolant System Asymmetric Loads Evaluation Program - Final Report.
Calvert Cliffs 1 and 2, Fort Calhoun, Millstone 2 and Palisades t
(2) Gerber, T. L., " Plastic Deformation of Piping due to Pipe Whip Load-ing", ASME Paper 74-NE-1, 1974.
(3) Combustion Engineering Topical Report, CENPD-168, " Design Basis Pipe Breaks", July 1975.
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TABLE 1: Assessment of Structures / Component of Question 110.1 .
Compon:nt/ Structure ^***'""*" "* "* ' " * ***"#* ""*
Status Basis
- Reactor Pressure Vessel Complete Plant Specific Analyses FSAR App 5.4A
'Ste:m Generators "
Design Cg Design Reports
, Rczctor Coolant Pumps " " "
. ftreter Vessel Supports Plant Specific Analyses FSAR App 5.4A
, Steam Generator Supports "
Design FSAR Sect 3.6.2.3 and App. 3.6A Rzector Coolant Pump Supports " " " "
) Biological Shield, Wall " "
FSAR Sect 3.8.3.3.1, 3.8.3a.2 and 6.2.1.2
- Stctm Cen. , R C Pump iCompartment Wall " " " "
!RCS Main Piping Considered Plant Specific Analyses , FSAR Sect 3.6.2.3 Waterford had Complete and Reference to.other App. 3.5A been analyzed Plant Analyses (Reference 1) for all pos-tulated breake 3 w/o consider-ation of asya-
, metric loads agdfora350 in cold Igg and 100 in let leg guillotine break at the reactor vessel nozzle for asym-metric loads also (see text).
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TABLE 1: Assessment of Structures /Componets of Question 110.1 (Continued) -
Component / Structure Assessment Evaluation Reference Comments
. Status Basis RCS Attached Piping Considered Simplified comparison (See Text)
(FCCS, etc.) Ceneri: Plant Complete analyses to previously St. Lucie 1 Analyses were for analyzed plants Docket 50-335 both cold and hot Reference 1 leg guillotine -
breaks RCS Attached Piping Supports n n (See Text) and Restraints -
CEIMS -"
Comparison to previously Reference 1 "
analyzed CEDH's St Lucie 1 Docket 50-335 (See Text)
Racetor Internals "
Comparison to analyses Reference 1 "
of similar plants (See Text)
Fual Will be addressed in response to Q 231.2 See footnote (See Text) and cover letter i
Prior analyses done for similar CE plants (with 14x14 fuel however) have indicated that the cold leg vessel inlet break is the determining break for fuel assessment.
Comparative analyses indicate that.there should be no problem with coolability of the core.
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i TABLE 2 1
3 Break Opening Times for RCS Postulated Breaks l
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Reactor Cavity MinimumOpeningz'y (from CENPD-168) -
' 2 100 in hot leg guillotine break 20 msec. 350 in2 discharge leg guillotine 6 maec.
i Steam Generator Compartment i
2 600 in hot leg guillotine 14 maec.
430 in suction leg guillotine 11 maec.
2 592 in suction leg guillotine 17 msec. 480 in discharge leg guillotine 28 maec.
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TABLE 4: Modal Frequencies (Hz) of Waterford 3 CEDMs l t
! Mode No.. Shortest In gest 1 3.4 2.7 ,
2 5.1 4.9 3 11.5 10.9 4 13.5 11.8 5 13.8 13.6 i f
6 23.8 15.0 I
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TABLE 5 Waterford .3._,ECCS Line Reaction Forces on Supports / Restraints (kips)
Themal - ,
Support / Restraint LOCA SSE (Normal) Total Capacity (Refer to Figure 6 for location of Supports / Restraints)
RCSR-92 0 19.2 o~ 19.2 91.0 (Snubber)
RCRR-94 11.0/-4.7 17.8 .929 21.8 26.0 (Strut)
RCRR-292 2.1/- 1.5 16.4 9.382 25.9 29.2 (Strut)
RCRR-100 1. 3 /- 1.1 23.1 6.008 29.1 35.3 (Strut)
RCRR-183 2.0/- 1.9 5.0 .212 5.6 27.0 (Strut)
RCRR-184 0 31.5 6 31.5 91.0 (Snubber) 4 4 ._ RCSR-186 0 27.3 0 27.3 91.0 (Snubber)
RCRR 293 1.1/-0 .
12.2 3.292 15.5 27.0 (Strut)
Total = Thermal + (LOCA + DBE )
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Appendices I l l
APPENDII A
. WATERFORD 3 SAFETT INJECTION LINE STRESSES DUE TO MOTION OF REACTOR The stresses developed in the Waterford SES 3 Safety Injection Lines are estimated by comparison with a previous analysis. Table A-1 provides such a comparison.
TABLE A-1 Stresses in ECCS Lines , Plant St. Lucia 1 , Waterford 3 Line 1-5-1 (File 3-7-9) 1RC12-40RL2B 1RC14-45RL2 0.D. 12.750 inches 12.750 inches 14.000 inches I.D. 10.126 inches 10.270 inches 11.636 inches Pressure 2235 PSI 2235 PSI 2235 PSI
#Press,ure 10,860 PSI
- 11',490 PSI 13,240 PSI Maximum Dynamic 2/3 inch *1/3 inch 1/3 inch Displacement .
Length af First 10.73 ft. 7.00 ft. 7.07 ft.
.,. Appropriate Section Bending (TACA only) 19 8I 2
26,830 PSI 2 28,890 PSI 2 p U Equation 9 3 = 45,000 PSI = 60,000 PSI = 66,000 PSI Since the maximum stresses calculated are at or above the 3 S limit of 48,000 psi, it is necessary to demon _s.st.; tate the igezEityllun.C319.nality_of these lines. To do this, the maximum bending moment wh_ich can be sustained by the pipe is computed and compared to bending moment actually developed in the lines. Gerber'spe mav4=um sustainable bending moment.is taken to be 70 percent of result; the actual bending moment is computed by a simp,11fied method explained later in this appendix.) Table A-2 summarizes such a comparison. I 1 i from PLAST ! 2 l Calculated from Simplified Method for Computation of Pipe Stress due to Specified Displacements presented in the subsequent pages. 8"I.IECPress + 2*# Bend I 1
**d 32 " 1*017
TABLE A-2 Maxi:num Load Carrying Capability of First Elbow in ECCS Pipe vs. Computed Moments f Plant St. Lucie 1 Waterford 3 Line 1-B-1 (File 3-F-9) IRC12-40RL2B IRC14-45RL2 t/ R 0.206 0.194 0.169 o 9* 0.309 0.292 0.253 b*/R, 0.0484 0.0458 0.0362 n 3 (40o 2) IN 93.3x10[PS. 3M 3 3h ' 3% (R, -Rg) 165.9 158.8 189.9 ( 0*/R,) 0.695 0.691 0.671 6 6 6 max 5.94 x 10 in-lb 5.6 x 10 in-lb 6.55 x 10 in-lb 7 max 4.15 x 106 in-lb 6 3.93 x 10 in-lb 4.58 x 10 in-lb bending 2.8 x 10 0 in-lb 0 3.2 x 10 in-lb 4.1 x 10 'n-lb i
. Seismic 6.25 x 10 5 in-lb 5
3.98 x 10 in-lb 2.76 x 10 in-lb
+ Computed by the method of Gerber- -for straight pipes, and scales down by the B 2 f*** # i" '9"**1 " 3'
I Simplified Method for Computation of' Pipe Stress due to Specified Displacements I Consider a piping line as shown in Figure A-1. ~GTv~en a specified dis ~ place-ment history at node 2 the question is to determine the mav4== s mding stress developed at any location in the line. Normally one finds such answers through the use of a sophisticated computer analysis. This sectioii"- describes a quick method for determining the approximate value of the bending stress. At present the analysis will be restricted to high frequency displacement schedules; that is to say, it is assumed that the displacement function does not excite one of the natural vibration modes of the piping system (which are generally low frequency)., An important consequence of this i assumption is that the largest displacements and largest bending stresses i
-occur near the forced node. Physically this is a result of the progressive attenuation of the displacements at greater distances from the forced node, the attenuation being due to the inertia of the system.
For example, consider the two piping segments shown in Figure' A-z, E the_ first line the inertia of the valve attenuates the lateral displacement field as long as one does not excite the natural frequency of that segment. In the second line the axial inertia of the pipe segment between nodes 3 and 4 attenuates the I-I displacements, though it will have little affect on the T-Y displacements. In terms of analysis, the not result of the above assumptions is that each straight pipe segnant can be analyzed independently for =v4== bending stress in terms of maximum lateral deflection.
~
In general, the pipe segment will displace as shown in figure X 3 7 the goal _ is to determine the = vd - bending moment (4 or M.,), given a set of displacements T , 0 , 6 and 9.,). To accurately determine this relation-shiponemusthahedktai$edinf5rmationabouttheeffectivelateraland torsional spring race of the remainder of the pipe line. While possible, such detailed considerations are inconsistent with the order of the approximations already made; but for an approximate result one can use the i t formula l ' 1
. l
' 2 M = 36EI/L as illustrated by the two examples of Figure A-4. Here M is the = v4 mum banding moment, 6 is the prescribed displacement, L is the length of the pipe section, E is Young's modulus, and I is the second moment of inertia of the pipe cross-section. ; l Accepting this one can then calculate the bending stress from a = Mc/I, where e is the -outside radius of the pipe section. Combining this with the above equation, one finds for the mavimum banding stress 1
l B-2 a = 36cI/L 2 . As a means for checking the validity of the above approximation, bending stress calculated according to the above formula is compared to the results of a PLAST computer analysis of a safety injection line (St. Lucie No.1. File 3-F-9) . Using the first appropriate line segment, similar to that between nodes 6 and 34 in Figure M,ona finds an approximate stress of 22,800 PSI as compared to 19,280 PSI computed by PLAST. t
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l l l t APPENDIX B JUSTIFICATION OF COLD LIG INLET BREAK AS,THE DETERMINING BREAK FOR FUEL, Analyses performed for the generic CE plant (Reference (1)), as well as i specific plants like Fort Calhoun, have indicated that the determining break insofar as fuel analysis is concerned, is a cold leg guillotine break at the reactor vessel inlet nozzle. This conclusion is based on the fact that the response of the 'uel for these plants has been analyzed for both a full area guillgtine break at the reactor vessel inlet nozzle sud a limited area (135 in ) guillotine break at the reactor vessel outlet
- nozzle, and that the former break has been found to be limiting break.
Another result of these analysis has been that the beam-column effect, due to concurrent lateral and av4al loading is also more pronounced for inalt break, but that this effect does not significantly increase the
! maximum fuel bundle stresses. (A dynamic beam-column analysis was performed for the plants of Reference (1) to determine any additional bending stresses and stability of the fuel assembly due to concurrent lateral and axial loading, and that analysis had also shown that the beam-column effects are more sensitive to lateral bending moments than to axial forces.)
The lateral banding moments on the fuel assembly due to the input excitation of the core support plate, fuel alighnent plate, and core shroud displacement time histories have been found to be significantly larger for the reactor vessel inlet breaks than for the reactor vessel outlet break. These analyses have been performed for a 14 x 14 fuel and therefore detailed stress results hawie limited applicability to Waterford 3. However, some of the conclusions reached for thcse plants have a direct bearing on Waterford 3 fuel assemblies also: a) The determining load condition on fuel is the lateral bending moments due to the excitation of the core support place, fuel alignment plate and core shroud lateral displacement time histories. These were largest in the generic plant (1) for the inlet break; however, the inlet break in the plant is a area break (350).in{ull areathe In both break whereas generic plant andininWaterford Waterford, itthe is a 2limited outlet break area is limited._ For the generic plant this break is 135 in while in Waterford 3 that break area is 100 in 2 , l Figures B-1 and B-2 compare the lateral motions of tba generic plant I vessel for cold and hot leg breaks. Significant displacements only occur in the x-direction (parallel to the hot leg direction) for hot leg break, so only that direction is shown.
I
. . j B-2 i i The motions of the vessel, together with the internal asy m tric !
j hydraulic load determine the lateral motions of the core support plate, fuel alignment plate, and core shroud. The vessel motions for the inlet break are clearly mach larger than those for a hot leg break for the generic CE plant. l For Waterford 3, the vessel motions (lateral) are essentially. half ! of those of the gener{c plant (See Figures B-3(a) and B-3(b) for an inlet break of 350 in ). ThemotionsofgheWaterfordvesselresult- l
- ing from a hot leg outlet break of2 100 in would be very similar to those of the generic plant (135 in ). However, the amplitudes of these motions would still be approximately230 to 40 percent less than the
, corresponding motions for the 350 in cold leg inlet guillotine break. In addition, the hot leg break would produce essencially no asyumetric internal hydraulic loads across the core barrel. Hence, it can be concluded that the inlet break is the determining break for establish- ' ing the largest lateral bending moments on the fuel bundle for Waterford ' also. b) Axial loads can be higher for the hot leg breaks, however, the generic - plant analysis has indicated that they are not as significant as lateral bending loads in determining fuel stability and bending stresses. ! , This, coupled with the fact that the beam-column analyses performed for the generic plant indicated neglig 91e additional effects on the fuel bundle; stresses, over thNa predic.:ed from the lateral loading analysis, is indicative of the colt. leg break being the determining break for the , fuel assessment. - ! c) For axial loads, the hot leg breaks would provide the dominant logds. The Waterford 3 plant has two postulate hot leg breaks, a 100 in breakatthevesseloutletanda600in{breakattheS.G. inlet. (Lateral Vessel motions, resulting from either brgaks, would be consider-ablysmallerthan-ghoseresultingfromthe350in cold les vessel inlet break. The 600 in breakwoulfproduceabout90percentofthevessel motions produced by the 100 in break.) These two breaks require different times to open. In any case, Waterford 3 has been analyzed l for the pure axial loads resulting from a full area hot leg break (the 1 so-called core bounce analysis), and the fuel was shown to be adequate for that bysak which groduced axial loads that are larger than either the 100 in or 600 in hot leg breaks. It is therefore concludad that the cold les vessel inlet break is the determining break for demonstrating the adequacy of Waterford 3 fuel assembliei. l l E . _ _ . _ _ _ _ _
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