ML20003E747
| ML20003E747 | |
| Person / Time | |
|---|---|
| Site: | Shoreham File:Long Island Lighting Company icon.png |
| Issue date: | 02/28/1981 |
| From: | LONG ISLAND LIGHTING CO. |
| To: | |
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| ML20003E745 | List: |
| References | |
| NUDOCS 8104100401 | |
| Download: ML20003E747 (400) | |
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APPENDIX A l LEAD PLANT ACCEPTANCE CRITERIA j (NUREG-0487, APPENDIX D) POSITIONS '
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Revision 4 - February 1981 810410040l,: ,
i Appendix A - Lead Plant Acceptance Criteria (NURM-0 487,
, ,-s Appendix D) Positions l
- 1. LOCA-Related Hydrodynamic Loads A. Pool Swell Loads
- 1. Pool Swell Elevation (modified by Section II.A.2 of NUREG-0487, Supplement No. 1)
NRC Position Shoreham Position Use Pool Swell Analytical Model (PSAM) Acceptable with polytropic exponent of 1.2 to a maximum swell height which is the greater of 1.5 vent submergence or the elevation corresponding to the drywell floor uplift AP per NUREG-0487 Criterion I.A.4. The associated maximum wetwell air compression is used for design assessment.
- 2. Pool Swell Velocity NRC Position Shoreham Position The pool swell velocity used to determine Acceptable
)
s_ impact and drag loads on wetwell components shall consist of the velocity predicted by the pool swell analytical model described in NEDE-21544-P multiplied by a factor of 1.1.
- 3. Impact / Drag Loads on Grating NRC Position Shoreham Position The static drag load, FSS, on grating Acceptable in the pool swell zone of the wetland shall be calculated for grating with open area greater than or equal to 60 percent by forming the product of pressure differential as given on Fig. 4-40 of NEDO-21060, Revision 2 and the total area of the grating. To account for the dynamic nature of the initial loading, the load shall be increased by a multiplier given by:
FD/Fss= 1 + gf1 + (0.0064 Nf) 2; for Wf <2000 in/sec, lO A-1 Revision 4 - February 1981 l
where:
Fo = dynamic load applied to the grating W = width of grating bars, in.
f = natural frequency of lowest mode, H z Fss= static drag load
- 4. Diaphragm Upward Load NRC Position Shoreham Position The maximum upward load, tiPUP on the Acceptable diaphragm shall be calculated by the correlation:
tiPUP = 8.2 - 44 F (psi) 0 $ F 5 0.13 fiPUP = 2.5 (psi)
F > 0.13 F = AB . AP VS VD . (AV) 2 where:
AB = break area AP = net pool area AV = total vent area VS = initial wetwell air space volume VD = drywell volume
- 5. Asymmetric Bubble Load (modified by Section II.A.3 of NUREG-0487, Supplement No. 1)
NRC Position Shoreham Position Use 20 percent of maximum bubble pressure Acceptable statically applied to 1/2 of the submerged boundary.
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- 6. Impact Loads on Small Structures NRC Position Shoreham Position The hydrodynamic loading function that Acceptable characterizes pool impact on small horizontal structures shall have the versed sine shape:
t p(t) =P max (1-cos 2n--)
where:
p = pressure acting on the pro-jected area of the structure Pmax= the temporal maximum of pres-sure acting on the projected area of the structure t = time T = duration of impact For both cylindrical and flat structures, v the maximum pressure Pmaxand pulse duration T will be determined as follows:
(a) The hydrodynamic mass per unit area for impact loading will be obtained from the appropriate correlation for a cylindrical or flat target on Fig. 6-8 of Reference 4.
(b) The impulse will be calculated using the equation NH I I
p
= VX A (32.2)(144) where:
Ip = impulse per unit area, psi-sec MH = hydrodynamic mass per unit area, A lbm/fta, from (a) above V = impact velocity, ft/sec, determined according to
() Section I.A.2 A-3 Revision 4 - February 1981
(c) The pulse duration will be obtained from the equation 0.011W (flat target)
T=
7, ,_;0463D (cylindrical 0 V V > 7 f t/sec target) f V T = 0.0016 V < 7 f t/sec where:
i T = pulse duration, sec D = diameter of cylindrical pipe, f t
- W = width of the flat structure, f t
! V = impact velocity, f t/sec i
(d) The value of Pmax will be obtained using the following equation:
P max
= 2I /T For both cylindrical and flat structures, a margin of 35 percent will be added to the Pmax values (as specified above) to obtain prudent design loads.
The load acceptance criteria, as specified above, corresponds to impact on
, rigid structures. The effect of finite l tiexibility of real structures will be l accounted for in the following manner.
I When performing the structural dynamic
{ analysis, the " rigid body" impact loads will be applied; however, the masses of
, the impacted structures will be adjusted by adding on the hydrodynamic masses of impact. The numerical values of I hydrodynamic masses will be obtained from i the appropriate correlations for I cylindrical and flat structures on Fig. 6-8 of NEDE-13426-P dated August 1975.
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h B. Steam Condensation and Chugging Loads
- 1. Single Vent Lateral Loads NRC Position Shoreham Position l r j The following single vent load will be used.
(a) A static equivalent load of 8.8 kips Acceptable shall be used provided that: -(i) the downcomer is 24 inches in ' diameter; (ii) the downcomer has a dominant natural frequency of 7 Hz or less in its submerged state; and (iii) the downcomer is either unbraced or braced at approximately 8 feet or more above the downcomer exit.
(b) A static equivalent load of 8.8 kips Acceptable multiplied by the ratio of the downcomer natural frequency and 7 Hz shall be used for downcomers with natural frequency greater than 7 Hz but less than or equal to 14 Hz.
This specification may be used only
- if the other restrictions outlined in item (a) above are satisfied.
, (c) If the natural frequency of the Acceptable, except downcomer is above 14 Hz, or if the that MKII Program downcomer is braced at a point Task A.13 load point closer than 8 feet above the definition will vent exit, a dynamic structural be used in lieu calculation of the downcomer re- of that identifi-sponse shall be performed on a ed in NRC Lead plant specific basis. For such a Plant Acceptance calculation, the dynamic load shall Criterion be defined by the equation: I .B .1 (c)
F (t) =F 0 sin ;D<t<T
= 0; for t < 0 and t > T l where 2 mscc <T <10 maec and the impulse I = F (T/r )x2 is 200 lbf-sec. This specification is also subject to restriction (i) listed in item (a) , above. Analyses of downcomer dynamic response to lateral chugging loads shall be performed for all plants during the MKII Long Term Program to provide
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additional confirmation that the static load specification is conservative.
- 2. Multiple Vent Lateral Loads NRC Position Shoreham Position The multiple vent load specification As noted in Sect-shall consist of the load specified on lons 1.2, 4.2.5.1, Fig. 4-10b of NEDE 21061-P, Revision 2 and 8.4, a dyna-multiplied by a factor of 1.26. For mic multivent downcomers with natural frequency greater analysis is in than 7 Hz, an additional multiplier equal progress. This to the ratio of this frequency and 7 Hz analysis super-shall also be applied. sedes the static analysis required by NRC Lead Plant Acceptance Criter-ion I.B.2.
II. SRV-Related Hydrodynamic Loads and Pool Temperature Limit
- 1. Discharge Device NRC Position Shoreham Position The applicants for all MKII facilities Acceptable will be required to commit to the use of quencher type devices.
- 2. Interim Load Specification l
NRC Position Shoreham Position (a) Methodology for Bubble Load Predic- Acceptable tion -
Those applicants that have committed to the use of the KWU "T"
- quencher device, shall use the SRV l air clearing loads based on the l predictions of loads for ramshead air discharges extrapolated to MKII conditions. The methodology for predicting ramshead loads is described in Section 3.2 of NEDO-21061 and NEDE-21061-P Revision 2.
! The applicants for all plants using l the four-arm quencher device shall
, use the methodology for predicting l the quencher discharge loads described in Section 3.3 of NEDO-21051 and NEDE-21061-P Revision 2.
O A-6 Revision 4 - February 1981 l
A final load specification for the quencher devices is under-develognent and shall be based on results from large scale tests which must be complete before plant operation. In plant tests are desirable. However, the need for these inplant tests will be determined as a part of NRC staff *s review of the Applicant's quencher supporting program.
(b) SRV Discharge Load Cases - The fol- Acceptable, except lowing load cases shall be consider- that Load Case 5 ed for design evaluation of contain- is not included ment structure and equipment inside in the ramshead the containment: design assess-ment basis.
- 1. Single valve discharge for Appendix G first and subsequent actuation, demonstrates
- 2. ADS valves discharge; that ramshead
- 3. Two adjaceat valves discharge, sequential ,
- 4. All italvt:s discharge sequent- entry (Load i ially by setpoint group, Case 4) bounds
- 5. All valves discharge simultan- the loads eously and assuming all bubbles actually imposed oscillating in phase. by the quencher device, even if O the quencher loads are applied simul-taneously in phase i The number of valves actuatted subsequently in the multiple valve cases (Cases 2, 4, and 5 above) shall be determined by plant unique analyses.
(c) Bubble Frequency - The forcing func- The ramshead bub-tion used to evaluate the SRV dis- ble frequency is charge load case described in item b deterministic as above shall include a spectrum of described in bubble frequency ranging from 4 to Appendix G.
11 Hz.
- 3. Pool Temperature Specification NRC Position Snoreham Position (a) Pool Temperature Limit - The sup- Acceptable, ex-pression pool local temperature cept that local shall not exceed 2000F for all plant pool temperature A-7 Revision 4 - February 1981
transients involving SRV operations. limit of 2000F The applicants are required to pro- is considered vide an in plant test data base for applicable only
, establishing the difference between when SRV dis-local and bulk pool temperature. The charge mass
! definition of local and bulk pool flux exceeds I .amperature is provided in the fol- 40 lba/fta sec
} lowing section. The applicants are as described l also required to provide plant in Section 3.4 i unique analyses for pool temperature
, responses to transients involving SRV operations to demonstrate that the plants will operate within the 3
limit.
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, (b) Local and Bulk Temperature - Local Acceptable f temperature is defined as the water
, temperature in the vicinity of the quencher device. For practical
\ purpose, measurement from the l temperature sensors, which are
- located on the containment wall in j the sector containing the discharge device and at the same elevation of l the discharge device, can be used as local temperature.
[ Bulk temperature, on the other hand, l is a calculated temperature which l
assumes the pool as a uniform heat
- sink. Bulk temperature is calculated on the basis of energy j
and mass released from the primary system through the safety / relief valves following the plant transients.
- 4. Suppression Pool Temperature Monitoring System NRC Position Shoreham Position The suppression pool temperature monitoring system is required to ensure that the plant is always operated within t the tecnnical specification lbnits. It l 1s our position that the applicants I should meet the following general requirements for the design of this system:
- 1. Redundant sensors shall be provided Acceptable at each monitoring location.
O A-8 Revision 4 - February 1981
- 2. The total number of monitoring loca- Acceptable tions shall not be less than eight.
Monitoring locations shall be dis-tributed. evenly around the pool.
4 3.. Sensors shall De installed auff1- Acceptable ciently below the minimum technical specification water level to assure that the sensor properly monitors pool temperature.
- 4. Pool temperature shall be monitored Acceptable on recorders in the control room.
Two sensors from each sensor group shall be recorded. The difference between measurement reading and actual local temperature shall be within 20F.
- 5. Instrument set points for alarm shall Acceptable be established so that the primary system can be shutdown and depressurized to less than 200 psia before the suppression pool temperature reaches the temperature limit as specified in 3 (a) .
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- 6. All ' sensors shall be designed to Acceptable
, seismic Category I, quality group B l and energized from onsite emergency
- power supplies.
ci III. LOCA/SRV - Submerged Structure Loads A. LOCA/SRV Jet Ioads
- 1. LOCA/SRV-Ramshead Jet Loads NRC Position Shoreham Position LOCA related vent jet loads and ramshead jet loads shall be calculated based on the methods described in NEDE-21730, subject to the following constraints and i modifications: i (a) Acceleration drag at the time the Acceptable jet first enc (.atnters the structure must be multiplied by the factor l
6Va 1+ CD ^x*i A-9 Revision 4 - February 1981
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where Va , CD , AX are the acceleration volume, drag coefficient and projected area of .
the structure as defined in NEDE- l 21730 and ni is the vent exit l radius.
(b) Forces in the vicinity of the front Acceptable shall be computed on the basis of an acceleration and standard drag (Formulas 2-12 and 2-13 of NEDE-21730). The local velocity, U= ,
and acceleration, U. are to be computed conservatively by the methods of NEDE-21471 from the potential function:
-3 cos e 4 = -- U j VW 8r r2 where r and 6 are the spherical coordinates from the jet front center with 0 measured from the jet direction, Uj is the jet front velocity from NEDE-21730 and Vw is the initial volume of water in the vent.
(c) After the last fluid particle has Acceptable reached the jet front a spherical vortex continues propagating. The drag on structures in its vicinity can be bounded by using the flow field from the formula for j# above with Uj as the jet front velocity from NEDE-21730 at time t = tf.
NRC Position Shoreham Position Quencher jet loads are expected to be Acceptable small. This load may be neglected for those structures located outside of a i sphere (or cylinder of 5 f oot radius) l circumscribed around the quencher arms.
! If there are holes in the end cap on the l quencher, the radius of this sphere l should De increased by 10 hole diameters.
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Confirmation of this assumption must be provided in the Long Term Program.
%/ B. LOCA/SRV Air Bubble Loads 1 LOCA Air Bubble Loads- (modified by Section II.C.2 of NUREG-0487, Supplement No. 1)
NRC Position Shoreham Position The methodology for computation of submerged structure loads during the air clearing phase of a postulated IDCA shall be based on an analytical model of the bubole charging process and drag calculations of NEDE-21471 until the bubbles coalesce. After bubble contact, the pool swell analytical model, together with the drag computation procedure of NEDE-21471 shall be used. Use of this methodology shall be subject to the
-following constraints and modifications:
(a) A conservative estimate of bubble Acceptable asymmetry shall be added by increasing accelerations and velocities computed in step 12 of
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Section 2.2 of NEDE-21730 by 10
, percent. If the alternative steps SA, 12A, and 13A are used, the acceleration drag shall be directly increased by 10 percent while the standard drag shall be increased by 20 percent.
(b) The drag coefficient Cp for the Acceptable I standard drag contribution in steps 13 or 13A and 15 of Section 2.2 and step 3 of Section 2.3 of NEDE-21730 may not be taken directly from the steady state coefficients of Table 2-3.
The method presented in Zimmer Nuclear Power Station FSAR for defining standard drag in unsteady tiow (Attachment 1.K, Amendment 99, Revision 61, dated September 28, 1979) is acceptable except that:
(1) CH = CM -1 in the formula for FA
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(2) CL = appropriate value or 1.6 for noncylindrical structures (3) Standard drag coefficients for O pool swell and oscillating air bubbles should be based on data for structures with sharp edges.
(c) The equivalent uniform flow velocity Acceptaole velocity and acceleration for any structure or structural segment shall be taken as the maximum values "seen" by that structure, not the j value at the geometric center. An exception to this position is that
! " segmented" structures with 1.0 51.5 I may be assessed using the velocity at the geometric center.
(d) Drag forces on submerged structures Acceptable may be computed independently of each other (as presented in NEDE-21730) for structures sufficiently far from each other so that
', interference effects are negligible.
i Interference effects are expected to be insignificant when two structures &
are separated by more than three W characteristic dimensions of the larger one. Por structures that are closer together, either detailed analysis of interference effects
- must be performed or a conservative mulciplication of both the
! acceleration and drag forces by a
' factor of four must be performed.
(e) A specific example of interference Acceptable which must be accounted for is the blockage presented to the motion of l the water slug during pool swell due l a the presence of downcomer bracing systems. If significant blockage relative to the net pool area exists, the standard drag j coefficients shall be modified for i
I tnis effect by conventional methods (see Pankhurst and Holder " Wind Tunnel Technique," Chapter 8, Pitman
& Sons, Ltd., London, 1952).
1 (f) Formula 2-23 of NEDE-21730 shall be Acceptable A-12 Revision 4 - February 1981 L
modified by replacing MH bY 'FB VA where VA is obtained from Tables 2-1 O and 2.
- 2. Ramshead Air Bubble Loads NRC Position Shoreham Position The methodology for computation of submerged structure loads during the- air clearing phase of a ramshead as described in NEDE-21471 shall be used, subject to the following constraints and modifications:
(a) Standard drag shall not be- Acceptable neglected without first estimating its order of magnitude. The relative importance of standard to acceleration drag depends on the size of the structure, size of the bubble, and the distance from the bubble.. The importance of standard drag can be estimated using the equation:
F R S
M [P max C'D .
min [Rmin)2
=f1 \
F ( p, / Tr C ( r j 3
M where:
FgM = maximum standard drag:
FAM = maximum acceleration drag C'o = cycle-averaged effective drag ,
coefficient; d = diameter of a cylindrical structure; Rmin= minimum bubble radius; r = distance from bubble center to the structure; and O
A-13 Revision 4 - February 1981
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/ max) max fl l = 8/3 for < 30.
j i P. / p. -
(b) The constraints and modifi- Acceptable cations described above in Section III.B.1, for LOCA air bubbles shall also be applied i to the calculation of ramshead air bubble loads as appropriate.
- 3. Quencher Air Bubble Loads NRC Position ,Shoreham Position No procedure or analyses for air clearing Acceptable loads on submerged structures for quencher SRV operation has been presented j by the Mark II Owners Group. These loads may be computed by the same basic
,! methodology as used for the ramshead device subject to the modification of the l source strength as substantiated by
- experimental data. An interim submerged structure load due to four-arm quenchers shall be determined using the source strength derived from bubble pressure calculated by the methods of NEDO-21061 Revision 2 and NEDE-21061-P for a four j arm quencher.
t Associated loads for the T-quencher may be computed on the basis of the ramshead methodology and bubble pressure described in NEDO-21061, Revision 2 and NEDE-21061-P. Ilowever, the bubble shall be assumed to be located at the center of the quencher device with bubble radius equal to the radius of the quencher. The bubble pressure of the quencher may be assumed to be 25 percent of the bubble pressure calculated by the ramshead methodology.
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O APPENDIX B CONTAINMENT STRUCTURE DESIGN MARGIN O
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TABLE OF CONTENPS O Section Title Page B
1.0 INTRODUCTION
B-1 B2.0 CRITICAL DESIGN SECTIONS B-2 B3.0 INTERACTION DIAGRAMS B-3 B4.0 INDIVIDUAL IDAD EFFECTS B-4 B5.0 LOAD COMBINATIONS B-5 B6.0 DESIGN MARGINS FOR LOAD COMBINATIONS B-6 B7.0 SENSITIVITY OF CONTAINMENT DESIGN MARGIN B-7 TO SRV AND CHUGGING LOADS B
8.0 CONCLUSION
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B-i Revision 4 - February 1981
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LIST OF TABLES Table Title l B4-1 Load ComDinations and Load Factors i
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l LIST OF FIGURES O Fiqure Title l
B2-1 Primary Containment Meridional Bending Moment - SRV l l Sequential Firing j i
B2-2 Primary Containment Meridional Axial Load - SRV l Sequential Firing B2-3 Primary Containment Meridional Bending Moment - DBA Pressure and Temperature B2-4 Primary Containment Meridional Axial Load -
DBA Pressure and Temperature B2-5 Critical Locations B3-1 Containment and Mat Internal Ioads b3 -2 Flexure and Axial Load B3-3 Interaction Diagram - Base of Primary Containment 83-4 Interaction Diagram Basemat - Just Outside Pedestal B4-1A Individual Load Effects - Interaction Diagram - Base (O,) of Primary Containment B4-1B Individual Load Effects - Interaction Diagram - Base l of Primary Containment B4-1C Individual Load Effects - Interaction Diagram - Base l of Primary Containment I
l B4-2A Individual Load Effects -
Interaction Diagram -
l Basemat - Just Outside Pedestal B4-2B Individual Load Effects -
Interaction Diagram -
l Basemat - Just Outside Pedestal B4-2C Individual Load Effects -
Interaction Diagram -
l Basemat - Just Outside Pedestal B6-1 Load Combinations - Using SRSS - Interaction Diagram - l Base of Primary Containment B6-2 Critical Load Combination - Using SRSS - Interaction l Diagram - Base of Primary Containment B6-3 Critical Load Combination - Using SRSS - Interaction l Diagram - Basemat - Just Outside Pedestal v/
B-iii Revision 4 - February 1981 l
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LIST OF FIGURES (Cbnt)
Fiqure Title B6 -4 Critical Load Combination -
Using Absolute Sum -
Interaction Diagram - Base of Primary Containment l B6-5 Critical Load Combination -
Using Absolute Sum -
I Interaction Diagram - Basemat - Just Outside Pedestal l B7-1 Sensitivity Study - Using SRSS - Interaction Diagram -
Base of Primary Containment l
l B7-2 Sensitivity Study - Using SRSS - Interaction Diagram -
Basemat - Just Outside Pedestal l
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B7-3 Sensitivity Study - Using Absolute Sum - Interaction Diagram - Base of Primary Containment B7-4 Sensitivity Study - Using Absolute Sum - Interaction Diagram - Basemat - Just Outside Pedestal l
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B-iv Revision 4 - February 1981
B 1.0 INTRODUCTICN Tnis study has been prepared to demonai. rate the capability of the containment structure to acc amodate arbitrary increases in the suppression pool hydrodynamic loads. Such increases could result from changes in load specifications or from analytical considerations such as fluid-structure interaction effects. The results of this study supplement the Design Assessment Report (DAR) which concluded that there is aufficient design margin for l the containment structure to sustain load combinations which include the effects of safety / relief valve (SRV) discharge and loss-of-coolant accident (LOCA) loads.
To study the eftects of assumed upward variations in SRV and chugging loads on the containment structure, critical design sections are selected at those locations having the minimum reserve capacity to sustain the additional increases. At these critical sections, detailed investigations are conducted into the relation of internal loads to ultimate capacity for all design loads and load combinations.
The information presented in this study substantiates the conclusion that the containment structure has sufficient design margin to accommodate substantial increases in SRV and chugging loads. In the most critical regions within the containment, there is sufficient reserve capacity to accommodate more than a 100 percent increase in SRV and chugging loads.
O B-1 Revision 3 - November 1978
l B2.0 CRITICAL DESIGN SECTIONS In order to study the sensitivity of the containment design to upward variations in SRV and chugging loads the design sections are selected at those locations which have the minmum reserve capacity to sustain the additional increases.
l Figures B2-1 through 4 illustrate how the containment Dending moment and axial load vary in the vertical direction for a safety / relief valve (SRV) sequential discharge and design basis accident (DBA) condition, respectively. It can be seen from these figures that the bending moment attenuates rapidly with distance from the base of the containment. Similiar behavior was found to exist in the basemat (i.e., rapid attenuation of bending moment away from the intersection with the pedestal and containment) . The information illustrated on those figures and the arrangement of containment reinforcement steel support the conclusion that the critical areas for this assessment occur at the basemat r.o containment and basemat to pedestal intersections.
l Figure B2-5 shows the critical containment and basemat design sections studied.
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B-2 Revision 4 - February 1981
B3.0 INTERACTION DIAGRAMS O The capability for a particular reinforced concrete section to sustain internal bending moment and axial loads without exceeding the allowable concrete and steel stresses can be conventionally established through the use of an interaction diagram.
Figure B3-1 illustrates the internal loads that act upon sections l isolated from the primary containment and basemat. The bending moment and axial load components are shown on Fig. B3-2 acting l
upon a typical reinforced concrete section. Any combination of axial load and bending moment which is found to lie on the curve entitled Ultimate Capacity would indicate that the reinforcing steel and/or concrete have reached their respective allowable stress / strain. The curve entitled Allowable Capacity is established by multiplying the ultimate loads by the capacity reduction factor (pt) . Therefore any point, such as point A, which lies inside this diagram will support that combination of axial load and moment and maintain stress levels within allowable limits . The distance between points such as A and B on the diagram is a measure of the reserve strength or design margin inherent in the section. Design margin as defined here, therefore, is the ratio of allowable load to calculated load.
Interaction diagrams for the critical design sections (Section B2.0) are shown on Fig. B3-3 for the base of the primary containment and on Fig. B3-4 for the basemat, just outside the pedestal. Whs.n using the interaction diagram O margin, to assess the region of interest, in which all the design load combinations lie, will be enlarged for clarity.
design l
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B-3 Revision 4 - February 1981 l
l B4.0 INDIVIDUAL LOAD EFFECTS Tne individual loads which comprise the load equations shown in Table B4-1 are compared in order to determine their relative h magnitude and therefore provide a means of assessing the effects of the SRV and chugging events.
l Figures B4-1A through 1C show the enlarged region of interest from the interaction diagram for the base of the primary containment (Fig. B3-3) . 'Ihe figures not only illustrate the amount of reserve strength for the design section, if each load had been acting alone, but also provide a pictorial means of comparing the relative magnitudes of dead load, intermediate break accident (IBA) , DBA, pipe rupture loads, and earthquake loads with that of SRV and chugging loads. Note on the figures that the static loads such as dead load ani gradual increase in DBA pressure and temperature appear as r'%gle points on the diagrams while the dynamic events such as SRV and earthquake appear as a block representing positive and negative variations in load.
A comparison of these figures indicates that the SRV and chugging loads are relatively small when compared to other loads appearing in the same load combination. Additionally, if the SRV and l chugging loads were acting alone there would be considerable reserve strength for the design section.
l Figures B4-2A through 2C show the enlarged region of interest from the interaction diagram for the basemat, just outside the l pedestal (Fig . B3-4) . As can be seen from the figures, a i
comparison similiar to that for the primary containment discussed
! above exists and the effects of SRV and chugging are small.
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B-4 Revision 4 - February 1981
B5.0 LOAD COMBINATIONS The load combinations for the containment and internal concrete structures including suppression pool loads are presented in Table 5-2 of the DFFR Rev. 2 and in Table B4-1 of this study. l The internal loads resulting from the SRV case of an all valve discharge are used in load combination equations (1) , (2) , (3) ,
and (6) since, as explained in Section 5, it provides the most severe state of stress for the containment. For the evaluation of the containment and internal structures, absolute sum method is used instead of square root of the sum of the squares (SRSS) method as discussed in Section 6.
Ioad Equations (4), (5) , and (7) of Table B4-1 include the combination of chugging and automatic depressurizing system (ADS) effects. However, the combination of these two events has also been found to be less severe than the case of an all valve discharge. Therefore, for conservatism, the results for these equations will be presented based on an all SRV simultaneous discharge.
Ioad Equations (4a) , (Sa) , and (7a) of Table B4-1 include the l combination of chugging and single SRV effects. However, this combination has been found to produce less severe results than l
those based on the internal loads resulting from a three adjacent valve simultaneous discharge (Case 5 of Section 5) . l O
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B-5 Revision 4 - FeDruary 1981 i
l B6.0 DESIGN MARGINS FOR LOAD COMBINATIONS The individual load effects described in Section B4.0 are combined in accordance with the load cczabinations delineated in l Table B4-1. In each of these load equations, the directions of the peak transient internal loads are assumed in such a manner as to produce the most critical conditions for stress calculations.
For independent, short duration, vibratory loads, such as seismic and SRV discharge the individual loads are combined by both the realistic SRSS method and the most conservative absolute sum method. Results based en SRSS are presented on Figs. B6-1 through 3. Absolute sum results are presented on Figs. B6-4 and B6-5.
Figure B6-1 shows how each of the ten load equations are positioned within the interaction diagram for the base of the primary containment. Recall that the rectangular block represents all positive and negative variations in the bending moment and axial load components. It can be seen from the figure that the minimum reserve strength (or design margin) is obtained from load Equation (7a).
l The critical loading Equation (7a) is isolated on Fig. B6-2 and the design margin is determined as 1.15. It should be noted that this margin is written against the allowable capacity and a l
further increase in load can be- sustained before reaching the l ultimate capacity of the section.
I l A similiar behavior was found to exist for the basemat and l Fig. B6-3 shows the critical load Equation (7a) and a design margin of 1.30. Again, as can ba seen from the figure, the design margin is even larger if the load components are measured against the ultimate capacity of the section.
Margins of safety are recomputed based on the absolute sum method l of combining dynamic load ef fects. Figures B6-4 and 5 present results at the base of the primary containment and basemat just outside pedestal respectively for the governing load combination 7-a. Even when this most conservative combination method is used, positive design margins are maintained.
O B-6 Revision 4 - February 1981
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<9,W;p
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I?.'+4'}p II/// / ,0 kf!g ,
//4,g%
~
4 - , , - , -
TEST TARGET (MT-3) 1.0 lp a na J '4
~
E.EtE @'2.2
. I.l tcm L'l2.0- .,
l.25 1.4 1.6 i
I
! g i
1 5
)
w p% + sh AM/$p>p
<'>O W ,4,y
+;
7/ 4npg g
j
- p r
/fff g///l:#%>*\> /
//
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45<,
\ ,
\ IMAGE EVALUATlON NNNN TEST TARGET (MT-3) l.0 E a Ed i != M g=n
- t in l-l ["2 lllllE l.8 l.25 1.4 1.6 i
t
/
4 6" * . I i
i i
f l e$r # # /4%
6/>4 y
%') '43 a <$
~
O +$.fpw? V
B7.0 SENSITIVITY OF CONTAINMENT DESIGN MARGIN TO SRV AND l CHUGGING LOADS The previous sections have described the SRV and chugging loads, load combinatioas, methods of analysis, and the concrete containment capability to sustain load.
This section will demonstrate the ability of the containment design to accommodate arbitrary increases in the SRV and chugging loads. Both SRSS and absolute sum methods of combining dynamic loads are considered.
Figure B7-1 shows the effect of doubling the SRV and chugging l loads in the critical load Equations (7) and (7a) at the base of the primary containment when the SRSS combination method is used.
It can be seen graphically from Fig. B7-1 that a significant l increase in SRV and chugging load decreases only slightly the reserve strength and thus the design margin. In fact, a 600 to 700 percent increase in the SRV and chugging loads is required before reaching the allowable capacity of the section.
Figure B7-2 shows the effect of doubling the SRV and chugging at l the basemat location, again, when the SRSS combination method is used. It can be seen from Fig. B7-2 that even with a doubling of l the SRV and chugging loads the reduction in reserve strength is small. The SRV load in Equation (7) could be increased by 400 percent while the chugging load in Equation (7a) could be increased without exceeding the allowable capacity of the design y section by 1,000 percent.
Figures B7-3 and 4 present similar results for the most conservative absolute sum method of dynamic load combination.
l m.v e n in this case the minimum capability for incr, ease is 200 percent for chugging loads at the base of the primary containment.
It can be concluded from the above discussion that the concrete containment can accommodate a significant increase in hydrodynamic loads before reaching its design allowable capacity and further increases within the ultimate capacity.
O B-7 Revision 4 - February 1981
l B
8.0 CONCLUSION
S As previously stated, this study has been prepared in order to determine the capability of the concrete containment structure to sustain arbitrary increases in the suppression pool loads.
It has clearly been shown that there is sufficient design margin inherent in the containment structure to accommodate more than a 100 percent increase in SRV and chugging loads.
This capability is more than enough to cover possible fluid-ctructure interaction effects.
Throughout the study, design conservatisms have been introduced so that the actual reserve capacity in the containment structure is larger than indicated.
In summary, the internal loads in the contalrunent structure resulting from the SRV and chugging events are relatively small and have little impact on the structural integrity of the containment structure.
O i
O B-8 Revision 4 - February 1981
( 5
'b \
i i
TABLE B4-1 IDAD COMBINATIONS AND IDAD FAC10RS LOAD 3
E COND. Q L F ( T, gu k b SS EB EA IA EA Er E i
, 1 Nonnal j w/o Temp 1.4 1.7 1.0 1.0 - - - - - - - - -
1.5
! 2 Normal I w/ Temp 1.0 1.3 1.0 1.0 1.0 1.0 - - - - - - -
1.3 l
j 3 Normal j Ser. Env. 1.0 1.0 1.0 1.0 1.0 1.0 1.25 - - - - - -
1.25 i 4 Abnormal 1.0 1.0 1.0 - - - - -
1.25 -
1.0 1.0 -
1.25 na Abnormal 1.0 1.0 1.0 - - - - - -
1.25 1.0 1.0 -
1.0c a >
l 5 Abnormal 1 Serv. Env. 1.0 1.0 1.0 - - -
1.1 -
1.1 -
1.0 1.0 -
1.1 Sa Abnormal
- Serv. Env. 1.0 1.0 1.0 - - -
1.1 - -
1.1 1.0 1.0 -
1.0'*8
- j. 6 Normal i,
Ext. Env. 1.0 1.0 1.0 1.0 1.0 1.0 -
1.0 - - - - -
1.0 i
7 Abnormal Ext. Env. 1.0 1.0 1.0 - - - -
1.0' 1.0 -
1.0 1.0 1.0 1.0 7a Abnormal Ext. Env. 1.0 1.0 -1.0 - - - -
1.0 -
1.0 1.0 1.0 1.0 1.0(*)
l >
LOAD DESCRIPPION 4
D = Dead Ioads Eo = Operating Basis Earthquake L = Live. Loads Egg a Sate Shutdown Earthquake F = Prestressing Loads PB = SBA or ILA Pressure Load l To = Operating Temperature Loads P3 = DBA IDCA Pressure Load Ro = Operating Pape beactions TA
= Pipe' Break Temperature Load
! Po = Operating Pressure Ioads RA = Pape Break Temperature Peaction Loads s
SRV = Oatety/kellet Valve Loads kr = Reaction and Jet Forces Associated with the Pipe Break
(*> Single valve actuation 1 of-1 Revision 4 - FeDruary 1981
REFUELING BELLOWS SEAL
\
('", , '
M STAR
,\
TRUSS ~i OUTER BELLOWS SEAL
\ ,
Q
- c. s
". 4 l .\ O J.; PR: MARY CONTAIN M ENT
.N O 4 .
t '.
?:
-, o 5- l l
& o y O
.., o c ,
4
'1 D
O
> i O I
-s O
.f I I l A I I r~I
.c : ~ - .>~~ :). -
,.:." '. r: : -100 -75 -50 -25 0 25 50 75 10 0 Mc -(FT K/FT)
REACTOR PRESSURE k VESSEL NO E:
FIG. 82 -I t u b N Tfi E iR cu uFE REN T: A tty ) PRIM ARY CONTAINMENT MERIDIONAL (m') BENDING MOMENT-V'-
SRV SEQUENTIAL FIRING SMCre H AM NUCLE AR PCwER STATION-UNIT I Pt ANT CESIGN ASSESSMENT FOR SRV AND LOCA LCADS REVISION 4-FEBRUARY 1981
RE.cuELING BELLOWS SEAL s
.} OUTER BELLOWS SEAL
& '* ~- .O f '
STAR TRUSS Ia '-
l U O
- PRIMARY RJ6 CONTAINMENT
~
. O i.
4 . O r:
() .i l .j O l l ..
i
. r.
- 4 0 1.: O ,
c .
b O ,
O
.. O
^~y O O
i; O O
I .
.'A' ()
E I I I i 9 v
i j i ! i
>;, :.6 : . .
-L. -yd
. ,9 . . _40 -30 -20 -10 0 10 20 30 40
~~
- ,- . 7: 4: . 7;. ;,:
N 33(K/FT)
N TE; '
FIG. B2-2 MAXIMUM INTERNAL LOADS
( MAX. IN TIME AND CIRCUMF ERENTIALLY) PRIMARY CONTAINMENT MERIDIONAL AXIAL LOAD-SRV SEQUENTI AL FIRING l SHOREHAM NUCL.r AR POWER STATION-UNIT I PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 4-FEBRUARY 1981
REFUELING BELLOWS SEAL CUTER BELLOWS SEAL w 3;*. O STAR TRUSS 8 .~ . .
I ' O d, ()
.~ 0 PRIMARY
~n CONTAINM ENT O
f4 c ..
O
,d l
M O
- p -
O
~s fO) .
- s O
el O O
O l l J o i i ! <2 i
- r . . -
- e '. A: r. *' ~;J.A- ~i..~,
. 3...;2 . -500
.. -300 -10 0 O 10 0 300 500
. ; u .. ~ : .a .
MSS (FT-K/FT) i REACTOR PRESSURE i
( VESSEL i
l NOTE:
l FIG. B2- 3 l '"' ""#' ' #
PRIM ARY CONTAINMENT MERIDIONAL BENDING MOMENT-i pd DBA PRESSURE AND TEMPERATURE I SHOREHAM NUCLEAR POWER STATION-UNIT I l
PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 4-FEBRUARY 1981 l ._. __._ . --
REFUELING BELLOWS SEAL C
OUTER BELLOWS SEAL
,' a STAR 4 '.
TRUSS ,j
- f. O
'.M O
PRIMARY CONTAIN MEN T o 6 '.
.4
.. O a:.
l V. O i:
O O O 4., -
O O
~}-
- h. O O
q ~-
0 l .h O
-i O i[- ' O i I I l o- 1 I I I
~ p* -9: * . r.r. j* ;: . : 2 l
0 50 l50 200
.-r. r:.-
>; .;. .:i 9:.::.
3: -200 -t 50 -l00 -50 l00 N3s(K/FT)
(
l l
l I-j NOTE:
FIGURE B2-4 INTERNAL LOADS.
PRIMARY CCNTAINMENT MERIDIONAL O
V AXIAL LOAD -
DBA PRESSU:tE AND TEMPERATURE SHOREHAM NUCLEAR POWER STATION-UNIT I PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 4-FEBRUARY 1981
O f
m
- 41.70FT. *
+---I I.lO FT.--*
/
/ .'f,.g ,
'i '
- v. n
- \. if
. 6.. SUPPRESSION POOL 4. /
TOP OF POO,L ,,
w ^ ^:_ ::::_ : _-
^
SUPPORT ~ ~
PRIMARY
~ "
PEDESTAL ~ CONTAINMENT
't-
! !A -
l ,
l l - E L.17 '-O" . " . '
! 1.P i.'
Y. 12" COVER, SLAB- INSIDE T CONTAINMENT l INSIDE )- - EL 9-O CONTAINMENT [ DESIGN SECTION PEDESTAL M ,
N @4 g*
, .. ..".'4
.,A; b
\ -EL. 8'-O',
\
! ,Y 1I
[ ,
'.5,.
-A. .
. ?
g .
D' l
! BASEMAT DES;uN SECTION CRITICAL LOCATIONS l FIG. B2- 5 CRITICAL LOCATIONS
- SHOREHAM NUCLEAR POWER STATKtJ-UNtT 1 l PLANT OESIGN ASSESSMENT FOR SRV AND LOCA L(MDS REVISION 4-FEBRUARY 1981
O
. Nss n
Os _
\ OUTSIDE OF PRIMARY CONTAINMENT rr y
\/
N' _ds Os " M ss y
N ss q
[ TOP OF MAT G
- O + M ss M ir
~
N A ja N ss N TT FIG. B 3-1 CONTAINMENT AND MAT INTERNAL LOADS
\ SHOREHAM NUCLEAR POWER STATION-UNIT 1 PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 4-FEBRUARY 1981 l
. , . . . . - - . . . . . - - . , . . - . . - . - , . - . - , . . - . . , , . . - - _ . . , , ~ - . - . . . _ . - . - - - - .
TYPICAL REINFORCED CONCRETE SECTION CRACKED CONCRETE REINFORCING f STEEL F_jlg,gi,f _ J _} ,
T 3 7-1 l i
E M9:'?. i::
) P { AXI AL
' . LOAD)
M (MOMENT)
INTERACTION CURVE & STRESS-STRAIN CURVE I YP
% ALLOWABLE STEEL N CAPACITY m f,c p \
- CONCRETE
\ z A B F ULTIMATE u M AX. US ABLE CAPACITY
- S TR AIN O
M ST R AIN Ou, DESIGN MARGIN ALLOWABLE LOAD DM = CALCULATED LOAD 65
=_
OA l
4 FIG. 83 -2 FLEXURE AND AXIAL LOAD
- b l
SHOREHAM NUCLEA R POWER STATION - UN6T I PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS I
REVISION 4-FEBRUARY 1981 i
, _ _ - . _ . . - ~ _ . _ - . . .
wo8 l O
~a8 l crdgam 9seH<-
x7 / l
\ /
/
\ /
\ N/
\ O I
S / br5g=m
\ $, I S / OD2ed E
% R 7 P /
\ M N
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/
F N / I I K /
,N (
/
D
~ A /
O
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- L o8
- A I
I
/
, X /
A
/
O - -
/
/ - - -
b8o 4g -
i53 8o 8o0 ' 8o
/
/ $ @_2o rogm2" \
/ - ) ,x 'nH~ \.
/ -
\
/ .o8 N
/
i
\
// O N
\
\
\
omo 9 I
\
/ S z4- 4 N g -
// / l E T
\
\ .
/ /
/(\
1o8 I 3o' $ l u
- N2>oqOz 9>c)%>K!
O W>m om o3 2 E>2 OOzq> _z %zH soAIC 2Rrgm my9 wYdO $=2
'C5 Mw3z >MgwE5,8 am<$a s9 sM*
mQ $, j=E3 G$
, : ,, < ;i l ,
5000 --
ULTIMATE CAPACI Y. ,
3 coon ALLOWABLE
/ NCAPACITY
./ N
/ -
/
[ o mo - -
5 a
\
N
/ 8 fE \
/-
J g
a.
o \
( w \. I
< ,ow s
N
- O Ns /
i i , , s , i i i 8000 m 4000 'zooc 2000 40m / . sooo BEPOING MOMENT (FT-K/FT)
N /
/
/' N N /
REGION OF INTEREST .,oon _
6
-2000 FIG. B3-4 INTERACTION DIAGRAM O BASEMAT-JUST OUTSIDE. PEDESTAL SHOREHAM NUCLEAR POWER STATION. UNIT 1 PLANT DESIGN ASSESSMENT FoR SRV AND LDCA LOADS REVISION 4-FEBRUARY 198f
SM -
N 250 -
\
- DEAD LOAD
\
CHUGGING k LOCA VENT CLEARING 'l t r -n 7 I t I f / f 1
' '" ' 1500 2000 2500 Moo
-500 500 1000 BENDING MOMENT (FT-K/FT) /
e l8A LOAD
-250 7
[ '
- DBA LOAD /
s i 3 /
e /
3 -500 -
/
ULTIMATE CAPACITY g /
a 5
-/
I 750 -
O x ALLOWABLE CAPACITY
/
'x -ioco N
/'
/
N /
\ /
\ /
\/*^
-1250 -
-isoo -
FIG. B4 -I A INDIVIDUAL LOAD EFFECTS-
" ~
RECTANGLES REPRESENT 1
[)
y VARtATIONS ON DYNAWl0 LOAos BASE OF PRIMARY CONTAINMENT SHOREHAM NUCLE AR PCWER STATION
. , , , . . _ _ , _,..y.,
soo
}-
(}
PIPE RUPTURE
'I 250 -
CBE F j-SSE \
// I r 4-I
, , , , , 1 . ,
-s::o Lp soo sooo -isoo zoooj 2500 sooo BENDING MOMENT (rT-K/FT) /
/
/
. - 2S0 -
/
C /
t /
5 /
e /
o j
9
-soo -
,/
/
.g ULTIMATE CAPACITY W /
- i /
g -7so -
/
ALLOWABLE CAPACITY, /
\ /
\ /
\ - /
-i j
\ /
\ /
\ /
\* /
-izso e
-isoo -
FIG. B4 -lB REcTuotEs neppESENT t INDIVIDUAt. LOAD EFFECTS-vaRiaticNs iN ormavic LcAoS. 'NTERACTION DIAGRAM-BASE OF PRIMARY CONTAINMENT SHoREMAW NUCLEAR Po*ER STATION-UNIT I PLANT CESMiN ASSESSMENT FoR SRV AND LCCA LCACS Rf.VtSICN 4 TEBRUARY 1988
500 -
)
N
- 250 -
SRV- ALL SEQUEtiTIAL SRV- 3 ADACENT IN PHASE \
SRV - ALL SIMULTANEOUS }
I
, rd +=P"
, , , , / , ,
-500 500 sooO 1500- 2000 / /2500 3000 BENDING MC tlF'!T (FT-K /FT ) 7
/
/
- 250 -
/
F
/
t /
5 /
e /
o /
' -500 -
f.
9 /
y ULTIMATE CAPACITY y /
a /
2 /'
x -750 -
7 ALLOWABLE CAPACITY
\ /
\, /
-seco - /
\ /
'\ \ /
/
/
\ /
\* /
-1250 -
t
-1500 -
(
marts FIG. 84-IC
' ONI[T$s'ro"fe"[u'a'v"Ic [cios. INDIVIDUAL LOAD EFFECTS-INTERACTION DIAGRAM-O
- r. sav pawsstao tonos.
BASE OF PRIMARY CONTAINMENT SHOREMAW NLCLCAR PC'*ER STATICN- UMT I PLANT CE54N ASSESSMENT roR SR v AND LCCA LCAOS J
REVISICl4 4-FEBRU ARY 198i
, , . . _ . . - . . ~ . . , - - - - . . . . ~ _ - - . _ - - -
C\
Q.
/
- OHUGGING '
[
l [3~
[LOCA VENTI
, f i^o ' ^o CLEARING
/ ,,
1000 ' 2000 3000 4000 5000 6000 BENDING MOMENT (FT-K/FT) [
/
_ zoo -.
/
/
/
s
\
.yo _
/
f ALLOWABLE
\ CAPACITY ULTIMATE CAPACITY N
l Ov $
5 \\ /
o 800 -
7
/
b /
5 h
-o00 w
-2000 -
e FIG. B4 - 2 A NOTE INDIVIDUAL LOAD EFFECTS-RECTANGLES REPRESENT 1 INTERACTION DIAGRAM-VARtATIONS N DYNAMIC LOOS BASEMAT-JUST OUTSIDE PEDESTAL
'% SHOREHAM NUCLEAR POWER STATKN-UNIT 1 PLANT DESIGN ASSESSMENT FCR SRV AND LOCA LOADS REVISION 4-FEBRUARY 1981 i
I
,,w..- -
,c.,__, , - - - , _ , ,
-. -,m..~. , ,. ~ , , , . , _ - , . . . , - - .-...--.--,---...w. -. , ,
O 4
/
PtFE RtPTURE [
/
[ce'= /
. , _' . ' . [ SSE
, . . . /- / , .
- *m i + oco . 2xo xco 4wo soco/ soco EENOWG WCWENT (FT-4/FT) /- ~
/-.
/
- 2x -
/
n /
t /
3 /
g . .cc - /
\ Au.OWABLE camcity
/
f
\t g /
o LTtMATE C#W. CITY
_a gN z \ /
o
% /
5 / '
5
~
\ /
.eaa -
% /
& y y
K %l /
.toco -
.axL i
FIG. B4 2B
<T E INDMDUAL (DAD EFFECTS-
- cisvArs w.)=cscv: 2 INTERACYtON DIAGRAM-u=>aTo<s m omw,c t:4:s BASEMAT-JUST OUTSIDE PEDESTAL
$M.EhLW **JCLEAst 8 Cwt 1t Su!Os -thit 1 PLANT 2$Gs ASSESSMENT POR SW ANO LCCA LCES i
s
! REVIS20% 4-FE8EU AU 1986 4
Y
}
-l
~i
/
20c - f, SR/- ALL SEQtENTIAL. /~
7 5AV-3SRV-ALL ADJACENT IN PMASE
$!MULTANEOUS /
/
w . / ,,
40W i I ooo roco 3:co 4000 y a BENDING MOMENT (FT-K/FT)
/
/
-zoo -
f
/
/
/
\ 4co -
/
\ /
O -
x( eta =r1x^ N' / JIMATE CAPACITY t
5
\ / .
\ /
o - *% -
\ /
8
' N /
%/
9 y .cco -
I #
.J 5
X 4 . zoo -
i i
l l.
l i
FIG,. B4 - 2C l-l setts INDIVIDUAL LOAD EFFECTS -
!!IE$U'E$NS,Yd2,.s INTERACTION O!AGRAM-l e sw uvs-tu ecs. BASEMAT-JUST OUTSIDE PEDESTAL l .
SKPEhW 6 W $'ATiCg. TNT 1 [
P.A%T oESGN AS!ti$WEhT FCR $F# UC LOCA Cc t
REV!SiON 4-FEBRUARY 39 81
O f
,_ / Soo -
)
OAD EQ.I 6
/ 1f
\
250 -
g .
4 \
I I / 3 \-
/ h
, , ,/ ./~ ' , / , ,
_500 500 ' 7' / Ao 2000 2500 3000 b I
[
[4A / BENDING MOMENT (FT-K /FT) f ,1 E
.230 _
l L I /p/-7^ /
/
i ! /
5 /
g .sw -
/
m ALLOWABLE CAPACITY [ CITY O a
< .750 -
/
\
\ /
\ _
/
\ /
\\//
3250 -
-t500 -
NOTE.
RECTANGLES REP 9ESENT 1 vARrAn0NS IN DYNAMC LOADS FIG. B6 - 1 LOAD COMBINATIONS-USING SRSS-INTERACTION DIAGRAM- '
]
j BASE OF PRIMARY CONTAINMENT SHOREWLW NUCLEAR power STATON-UNIT I PLANT MSIGN ASSESSMENT FCR SRV AND LOCA LOADS REVISON 4-FEBRUARY 1991
. . . _ . ~ . . .. - . . _ _ , - . .- , . --
/
l I
l 500 -
}
\ -
25o _
\-
\.
. \
04D EQ 7A
, O , , f -, , , , ,
500 500 '000 I '600 2000 / 2500 3000 BENDING MOMENT
/ (FT-K/FT ) -
f n
250 -
l s
[ A E '
C o
< / '
9 z 50o ._
a /
5 /
ALLOWABLE CAPACITY y 750 - LTIMATE CAPACITY 5 e
i N
[
\ /
N -
/
N /
\ /
' .\ [ ALLO *ABLE LOADS , g ,
DESIGN MARGIN = CALCULATED LOADS OA
-f 250 -
CESIGN MARGIN 8 LIS
-1500 -
FIG; B 6-2 CRITICAL LOAD COMBINATION-USING SRSS-INTERACTION DIAGFtAM--
O d BASE OF PRIMARY CONTAINMENT SHOREHAM llUCLEAR POWER STAil0N-UNIT 1 PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS
- = REVISION 4-FEBRUARY 1981
._-. ., , . . , . . . - . . , . . . - . - - , . . . . . , . - . _ _ - - _ _ ~ , . , - ~ . - . - - - _ , . , . - _ . . . . . . _ . - . . . . , . . . . , , . . ~ , _ . . . - . . . _ _ . . , - . ,
O
/
m-
/
ri.OAD EQ.7A
/.
I 1 , , /
-I000 2000 3000 4000 5 6000 (FT-K/FT)
~
Bf -C
. zoo -
/
7
/
400 -
\ /
'\ /
I N /
$ g AU.OWABLE CAPt4 TTY /
)
y /
ULTIMATE CAPACITY a
$ -800 - f l
A 4
%/
-1000 -
DESIGN MARGIN: ALLOWABLE LOADS , OB
-4 000 -
CALCULATED LCA0s CA DES 4N MARGIN s 1.30 FIG. 86 -3 CRITICAL (JOAD COMBINATION-USING SRSS -
INTERACTION DIAGRAM-BASEMAT-JUST OUTSIDE PEDESTAL '
b SHOREMAW NUCLEAR POWER STATION UN1T 1 PLANT DESIGN ASSES $UENT FOR SRV ANO LOCA LOADS .
4 REV15'ON 4-FEB RUARY l981
\
500 -
)
\
- 25o -
\
\
\
OAD EQ 7A
, O , , . , , f , ,
500 500 1000 1500 2000 / - 2500 3000
-/ BENDING MOMENT (FT-K/FT )
250 -
~ /
5 /
Q ^/ B g c Z 500 -
8 /
5 /
O :
g 750 ALLOWABLE CAPACITY
/
LTIMATE CAPACITY N -
\ /
\
l
\ /
\ -
/
N /
\ /
A OAD ,0 CESIGN MARGIN
-I2SO -
DESIGN MARGIN
- 1.06
-1500 -
FIG. B6 -4 i
CRITICAL LOAD COMBINATION-USING ABSOLUTE SUM -
('g INTERACTION DI AGRAM-l V BASE OF PRIMARY CONTAINMENT SHOREHAM NUCLEAR POWER STATION-UNIT 1 PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 4 FEBRUARY 198i
- m. .
l
/
M-j D EQ.7A
/
, , 0 ,
, j /
sy
..wo exo xoo .wo
- 47 MOMENT.
<mc
.A Bf C (FT- K/FT) f
. zoo
/
/
.vo -
/
\ /
\ /
- N /
h ( ALLOWABLE CAPACITY [
O i e
5
-ex -
'x N N /
e> >'c c >c>r<
A
- t
%/
-iew i
l ocs,cu u. sis. Auf eABLE LCAOS , g[
-120c0 -
CALCULATED LCA0s CA l OES4N tAARGN = 1.15 i
[
FIG. B6-5 CRITICAL LOAD COMBINATION-USING ABSOLUTE SUM-INTERACTION DIAGRAM-3 BASEMAT-JUST OUTSIDE PEDESTAL j SHOREHAM NUCLE AR POWER STATION - UNIT I PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 4-FE9mJAnv 198t
M
(% "
500 -
)
\
250 -
EQ 7 WITH SP/ DOUBLED \
-\
EQ 7
- }
1000 1500 2000 2500 3000 500 f f t f f
/' t t
./ BENDING MCMENT -
(FT-K /FT)
'/
/ .
EQ.7 WITH SdV.
250 -
INCREASED 7 TIMES
-/-
EQ. 7A f 7
/ EQ.7A WITH CHUGGING-INCREASED 6 TIMES.
EQ. 7A WITH CHUGGING DOUBLED !
p -soo -
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TABLE OF COlfrENTS O Section Title Page C
1.0 INTRODUCTION
C-1 C2.0 FSI EFFECTS ON SUPPRESSION POOL IDAD DEFINITIONS C-2 C2.1 FSI Effects on SRV Loads C-2 C2.1.1 Ramshead Device C-2 C2.1.2 T-Quencher Device C-2 C2.2 FSI Effects on LOCA Loads C-3 C2.2.1 Chugging C-3 C2.2.2 Condensation Oscillations C-4 l C3.0 FSI EFFECTS ON STRUCTURAL RESPONSE C-5 l C3.1 Considerations in the' Design Assessment Report C-5 l C3.2 Mark II Generic Results C-6 l C3.3 Shoreham Plant Specific Results C-6 l C3.3.1 Method of Analysis C-6 l C3.3.2 Response to SRV Loads C-7 l C3.3.3 Response to LOCA Loads C-8 l C
4.0 CONCLUSION
S C-9 l CS.O REFERENCES C-10 l O
C-i Revision 4 - February 1981
LIST OF FIGURES Figure Title C3-1 Fluid-Structure Interaction Model C3-2 SRV 10 Hz Ramshead Ioad - All Valve Simultaneous Discharge Radial Displacement Profile Primary Containment C3-3 SRV 10 Hz hamshead Load - All Valve Simultaneous Discharge Vertical Displacement Profile Primary Structures C3-4 SRV 10 Hz Ramshead Load - All Valve Simultaneous Discharge Vertical Acceleraticn Profile Primary Structures C3-3 SRV 10 Hz Ramshead Load - 3 Adjacent Valve Simultaneous Discharge Horizontal Displacement and Acceleration Profiles Primary Containment C3-0 SRV 10 Hz Ramshead Ioad - All Valve Simultaneous Discharge Vertical Amplified Response Spectra Top of Peactor Support Pedestal C3-7 SRV 10 Hz Ramshead Inad - All Valve Simultaneous Discharge Vertical Amplified Response Spectra Primary Containment at Elevation of Stabilizer C3-8 SRV 10 Hz Ramshead Load - 3 Adjacent Valve Simultaneous Discharge Horizontal Amplified Response Spectra Primary Containment at Elevation of Stabilizer C3-9 Chugging - 20 Hz Uniform Radial Displacement Profile Primary Containment C3-10 Chugging - 20 Hz Unif orm Vertical Displacement Profile Primary Structures C3-11 Chugging - 20 Hz Uniform Vertical Acceleration Profile Primary Structures C3-12 Chugging - 20 Hz Asymmetric Horizontal Displacement and Acceleration Profiles Primary Cont?inment O
C-ii Revision 4 - February 1981
LIST OF FIGURES (CONT'D)
Fiqure Title C3-13 Chugging - 20 Hz Uniform Vertical Amplified Response Spectra s
Top of Reactor Support Pedestal C3-14 Chugging - 20 Hz Uniform Vertical Amplified Response Spectra Primary Containment at Elevation of Stabilizer C3-15 Chugging - 20 Hz Asymmetric Horizontal Amplified Response Spectra
- Primary Containment at Elevation of Stabilizer a
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C-iii Revision 4 - February 1981
APPENDIX C FLUID-STRUCTURE INTERACTION (FSI)
C1.0 ItfrRODUCTION The Shoreham Design Assessment Report (DAR) for safety reilef valve (SRV) and loss-of-coolant accident (LOCA) loads discusses the hydrodynamic suppression pool phenomena and tne resulting structural response. The ef fects of fluid-structure interaction (FSI) on the plant assesament, which have not been explicitly considered previously, are discussed in detail here, as they apply to the Shoreham design basis loads.
FSI effects considered herein are of two types. First is the possible modification of the pressure history from the source (SRV discharge device or downcomer vent) to the pool boundary due to motisn of the pool boundary walls. This FSI effect is or concern from the point of view of determining the appropriate wall load specifications. Extensive research in this area has
- been done recently, the results of which are discussed in Section C2.0.
The second type of FSI effect considered is that of the influence of the suppression pool water on the structural 2:esponse to the load. The issue here is the proper representat'.on of the water s in the structural model. The results of both Marx II generic and Shoreham plant specific studies of this effect are presented in Section C3.0. The plant specific studies address both axisymmetric and asymmetric SRV and chugging loads.
O C-1 Revision 4 - February 1981
C2.0 FSI EFFECTS ON SUPPRESSION POOL LOAD DEFINITIONS C2.1 FSI Effects on SRV Loads C2.1.1 Ramshead Device Fluid-structure interaction effects on measured wall pressure data during SRV actuation with a ramshead discharge device have been assessed in a report issued by General Electric Co. (GE) ( a ) .
The results are not directly applicable to the Shoreham assessment because of the use of a T quencher discharge device instead of a ramshead device, but they are presented here for intormation purposes.
The referenced report contains an investigation of FSI effects in the Monticello Mark I containment. Three analytical models of the coupled fluid-structure system are developed and usad to predict wall pressure response to an idealized bubble forcing tunction inside the fluid. Wall response is predicted for a range of containment diameter to thickness ratios (rigid and flexible containments). A camparison of predicted wall pressures for a flexible containment with those measured experimentally in the Monticello test show the same important characteristics, thereby demonstrating the adequacy of the analytical techniques.
Finally, a comparison of results from the flex 1 Die containment cases with the rigid containment results indicates no significant FSI ef fects.
It is therefore concluded that the measured wall pressures in the Monticello tests contain insignificant FSI effects and furthermore this test data can therefore be used to benchmark the Mark II ramshead methodology.
C2.1.2 T-Quencher Device Tne T quencher discharge device and associated wall pressure time histories have been discussed in Section 3 of the DAR. When the tests from which this data was obtained were performed, efforts were made to minimize fluid-structure interaction effects by attaching the pressure transducers to very rigid walls. An assessment of FSI ef f ects in tne T quencher test pressure traces has been conducted by Kraftwerk Unionta) (KWU) and is summarized in this section.
The basic approach was to compare power spectral censity tunctions developed from the early stage or the pressure traces with those developed from longer durations to determine if any structural response effects appear after the bubble pressure amplitude has been damped out. Power spectral density functions were rirst developed for the first 0.6 seconds of the measured pressure traces. The predominant bubble frequency can clearli be identified in these results to be between 6 and 8 Hz. Power spectral density functions were ther developed for the same pressure traces out carried out for a longer duration such that the bubble pressure amplitudes had been damped out. The same h C-2 Revision 4 - February 1981
bubble frequency is still found to-be dominant. If the measured traces had been influenced by structural response, this would
()
(m) haec 1ppeared in a comparison of the two power spectral- density functions for each pressure trace.
These results demonstrate the fact that there is no FS1 effect in tlun measured wall pressure histories for the T quencher device and therefore it is appropriate to use them as pure Dubole pressure forcing functions on the wetted suppression pool boundaries.
C2.2 FSI Effects on LOCA Loads This section addresses FSI effect on the Shoreham design basis chugging and CO loads, as described in Section 4. Any FSI effects related to revised steam condensatisn loads derived from the new 4T CO tests will be addressed in Appendix N.
C2.2.1 Chuqqing The chugging load specification used for the Shoreham Design Assessment, as described in Section'4 of the DAR was based on data from the temporary tall tank test (4T) facility. Tne effects of fluid-structure interaction on this test data has neen investigated and the results have been presented by GE(3). l Following is a brief review of the metnods and results of the study.
m
) The first step was the experimental determination of the 4T f acility natural frequencies and mcde shapes. These results were then used to qualify an analytical model of the system.
Simulated test chugs were also performed and the results were compared to analytical predictions. Finally, the analytical model was used to determine the effects of the 4T wall flexibility on the measured response.
Extensive analysis and assessment of the results has led to a fundamental understanding of chugging in the 4T facility. Three important conclusions are drawn from the results of. this ef fort.
The first conclusion is that the measured pressure histories have been influenced by the 4T f acility structural characteristics.
The nature of this influence is such that the chug pressure amplitude is reintorced and that a system response in the form of a sinusoidal like ringout is included in the measured pressure
, traces. The second conclusion is that the effect of actual fluic structure interaction (moditication of the wall pressure due to movement of the wall / fluid interface) is small - on the order of 10 percent. The third conclusion is most important trom the point ot view of the plant assessment. That is, even though the 4T results have been influenced by the system flexibility, .the effect is to result in conservative predictions of structural response when the data is applied directly as pool boundary loads.
O C-3 Revision 4 - February 1981
l l
l C2.2.2 Condensation oscillations Condensation oscillation (CO) loads are predominantly of low frequency content. 7n fact, the Shoreham design basis CO load has frequency content only in the range of 2 to 7 fiz .
Studies (*) of the 4T facility have shown that in this frequency range FSI has virtually no effect on measured pool boundary response.
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O C-4 Revision 4 - February 1981
C3.0 FLI EFFECTS ON STRUCTURAL RESPONSE C
C3.1 Considerations in the Design Assessment Report Tne primary manner in which the suppression pool water affects the response of the containment structures to dynamic loads is through its inertial effects. This has always been considered and accounted for in the structural analysis for suppression pool loads by selectively lumping the water mass at nodal points in the mathematical model. Any additional effects of actual fluid-structure interaction were considered to be secondary and as such would be adequately covered by a number of conservatisms in the overall design assessment.
Recent analy3is, discussed in the following sections, has confirmed the original assumption that fluid-structure interaction effects are of a secondary nature. Conservatisms in various atages of the design assessment have been discussed in tne appropriate sections of the DAR. The most signiticant ones are restated here.
- 1. Conservative asuumptions and methods have been used in the structural analyses for all original londs which must be considered with the suppression pool loads.
- 2. All suppression pool load specifications used in the m design assessment have been conservative.
4
- 3. A conservative measure of structural damping (less than 4 percent damping in the entire frequency range of interest) has been utilized in all containment structure analyses for hydrodynamic loads. No damping was used for the assessment of primary structures for SRV loads.
- 4. Resultant forces in the containment structures due to suppression pool loads are small as compared to those due to the original design loads. Therefore, margins of safety are not sensitive to variations in the new loads i (see Appendix B) .
- 5. Maximum axial loads , moments , and shears for each dynamic load are assumed to act srmultaneously (i .e . ,
added by absolute sua) when computing member stresses.
- 6. Conservative damping values ar6 used in subsystem analyses.
- 7. Amplified response spectra resulting from the ,
suppression pool loads are peak broadened by a full 15 percent.
- 8. Design assessments are based on minimum specified
() material strengths.
C-5 Revision 4 - February 1981
l 1
Hnen' all of these features are considered, especially collet:1vely, it is evident that more than enough conservatism exists in the design assessment to cover secondary effects such h;I as fluid-structure interactica -
C3.2 Mark II Generic Result . l A generic study of the effects of fluid-structure interaction on Mark II containment res': .a has been conducted by Burns and Roe for GE(5). Three ca.egories of Mark II containments were analyzed - prestressed concrete, reinforced concrete, and steel.
Response wa s compared from the dynamic analysis of the coupled tluid-structure system and from the system in the absence of water. Some significant differences were found and plant specific FSI analyses were recommended. An extensive FSI analysis has been performed for Shoreham and the results of it form the basis for conclusions regarding the importance of FSI in Shoreham. It is noted for information purposes only that the largest ef fect found for Shoreham in the generic study was a 35 percent increase in the pool wall displacement. As stated above, this was compared to results obtained from an analysis which did not even include the inertia effects of the suppression pool water.
C3.3 Shoreham Plant Specitic Results C3.3.1 Method of Analysis Methods of analysis of the containment structures for suppression l pool loads nave been discussed in detail in Section 2.4.2. All assumptions and analytical techniques described therein have been utilized in the fluid-structure interaction analysis, except as regards the treatment of the suppression pool water. In the design assessment the prinary effect of the water was accounted for oy including its mass in the structural model. The finid-structure interaction analysis has been performed by represent 1Rg tne suppression pool water by a detailed finite element discretization. The fluid elements utilized have Deen converted trom solid elasticity finite elements by a technique applicable l to the modeling of an inviscid, compressible fluid (6 ). A sketch or the mathematical model, including primary structures, (i.e.
the reactor pressure vmsel (RPV) , and the suoporting soil) , as well as the suppression pool water is shown on Fig. C3-1.
Four representative hydrodynamic Acad events were analyzed using the detailed fluid-structure interaction model. They were the ic11owing:
- 1. 10 Hz SRV ramsnead load - simultaneous discharge of all valves,
- 2. 10 Hz SRV ramshead load - simultaneous discharge of three adjacent valves,
- 3. 20 Hz uniform chugging, and C-6 Revision 4 - February 1981
- 4. 20 Hz asymmetric chugging.
The first event was chosen because it is comparable to that which produced the largest FSI effect for Shoreham in the generic study.
described in Section ~C3.2.. The third case was selected as a representative chugging load. Chugging loads rather than CO loads are used in this study oecause, of the LOCA loads, they have the greater potential for the results to be influenced by FSI. No signiticant FSI effect due to chugging was identificd in the generic study, however, Cases 2 and 4, asymmetric loadings, were chosen for completeness of this study. For all four of these events, all results (i . e . , displacements, accelerations, forces, and amplitied laccmee spectra (ARS)) are compared to these obtained in the original design assessment analysis.
C3.3.2 Resoons qto SRV Loads The evaluation of the effects of fluid-structure interaction on structural response to SRU loads is based on the 10 Hz SRV ramshead simultaneous discre;ge load definition. As mentioned earlier, this load definition was selected for this study because it is comparable to that which resulted in the largest FSI ei'ect for Shoreham in the generic study.
In that study radial displacement of the primary containment was used as the basis of evaluation of FSI effects. Figure C3-2 presents the same ccxnparison for the Shorenam plant unique results. The predominant overall structural response to an all V valve discharge is in the vertical direction. Profiles or verticai displacement and acceleration of the primary structures with and without FSI are presented on Figs. C3-3 and 4. The l predominant response to the asymmetric load is in the horizontal direction. Horizontal displacement ano acceleration proriles ror the primary containment with and without FSI are presented on Fig. C3-5. l As can be seen trom these results, FSI effects produce no ma3or changes in the structural response. Tne largest ditferences at points of maximum structural response are on the order of 10 to 20 percent. Overall, the diffsrences are even less. The degree of eff ect also applies to resultant interaal loads in the primary structures. This is more than covered by the ract that SRV results based on the extremely conservative case of no structural damping was used f or the assessment of primary structures.
Amplitled response spectra comparisons have also Deen developed for both axisymmetric and asymmetric SRV loads. Representative results are presented on Figs. C3-6 through 8. Due to the l relative complexity of ARS, they could be expected to show the least favorable ccxnparison. However, even here the agreement is excellent, confirming the secondary nature >f FSI on structural response to SRV loads.
C-7 Revision 4 - February 1981 L
C3.3.3 Response to LOCA Loads As discussed above, chugging loads have a greater potential for the response to be influnced by FSI than do CO loads and are therefore used in this study. The 20 Hz chugging load specification was chosen arbitrarily to evaluate the effects of FSI on structural response to chugging loads. All of the comparisons discussed or presented for SRV loads c. re also presented here for chugging loads and the same trends are l observed. Figure C3-9 presents the important comparison of maximum radial displacement of the primary containment with and without the detailed representation of the suppression pool water. Profiles of vertical displacement and acceleration of the primary structures due to uniform chugging are shown on Figs. C3-10 and 11. Figure C3-12 presents a comparison of horizontal displacements and accelerations of the primary containment due to asymmetric chugging. As for SRV loads, there is no major difference in response. Maximum local variations in structural response - displacements, acceleration, and forces - are on the order of 10 to 20 percent.
Finally, comparisons of ARS due to unitorm and asymmetric l chugging are presented on Figs. C3-13 through 15. As for SRV loads, the differences are insignificant, demonstrating the
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O O
C-8 Revision 4 - February 1981
C
4.0 CONCLUSION
S
(' All aspects of fluid-structure interaction effects on the Shoreham plant assessment have been considered in this report.
The fundamental conclusion is that FSI effects which were not explicity addressed in the DAR are either insignificant or adequately covered by many other conservatisms in the evaluatica.
This applies to both the influence of the structure on measured wall loads as well as the effect on the fluid on structural response to the specified wall loads.
Following is a more detailed summary of conclusions:
- 1. The PSI effect on the wall pressures measured for SRV discharge with a ramshead device at the Monticello facility is insignificant.
- 2. Because FSI is insignificant in the Monticello tests, those results can be used to benchmark the Marx II ramshead methodology.
- 3. There is no FST 2ect in the wall pressure tracers used in the T quencher load specification.
- 4. There is an FSI ef fect in the chugging test data in the form of system ringout.
- 5. The system ringout in the 4T chugging data has a conservative influence on the structural response when the measured traces are applied as wall loads on the containment.
- 6. FSI effects on structural response to suppression pool loads must be considered on a plant specific basis.
- 7. Shoreham plant specific FSI analysis demonstrates that the effects on structural response are of a secondary
, nature.
l
- 8. The secondary effects of FSI on structural response are j more than covered by numerous conservatisms in tne design assessment.
i C-9 Revision 4 - February 1981
CS.0 REFERENCES
- 1. " Fluid-Structure Interaction Analysis of Ramshead Safety Relief Valve Discharge Device in a Mark I Steel Containment Torus," NEDE-23834, June 1978.
- 2. "Thermo-Hydraulic Quencher Design of the Safety Relief System," KWU Report R14-25/1978.
- 3. " Summary of 4T Fluid-Structure Interaction Studies," NEDE-23710, October 1977.
- 4. "Cc.aensation Oscillation (CO) Load Data for LaSalle,"
prepared by G.E. and S. Levy, Inc., submitted to the NRC on the LaSalle docket July 11, 1980.
- 5. " Evaluation of Fluid-Structure Interaction Effects on BWR Mabk II Containment Structures," NEDE-21936-P, Jul'1 1978.
- 6. A.J. Kalinowsky, " Transmission of Shock Waves into Submerged Fluid Filled Vessels," ASME Conference on FSI Phenomena in Pressure Vessel and Piping Systems, TVP-TB-026, 1977.
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O C-10 Revision 4 - February 1981
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FIG. C3-14 CHUGGING-20 HZ UNIFORM VERTICAL .
AMPLIFIED RESPONSE SPECTRA PRIMARY CONTAINMENT AT ELEVATION OF STABILIZER SHOREHAM NUCLEAR POWER STATION -UNIT 1 PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 4-FEBRUARY 1981
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l APPENDIX D i
RESPONSE TO NRC QUESTIONS l
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l Revision al - February 1981
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4 PREFACE Appendix D contains a list of NRC questions which were issued l suosequent to the original issue of this report. The detailed responses to the questions have been incorporated in the DAR j either directly or by reference to generic snhmittals.
Tables D-1, D-2, and D-3 provide a keyword index to the questions of June 23, 1976, January 17, 1977, and May 29, 1978, respectively.
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D-i Revision 4 - February 1981 I
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TFLBLE D-1 l QUESTIONS OF JUNE 23, 1976 O 3 Number QuestionO) Keyword Index to Questions 020.1 Condensation Oscillations 020.2 Single Downcomer Horizontal 4.3 Condensa-tion Loads 020.3 Multiple Downcomer Borizontal Condensa-tion Ioads 020.4 Downcomer Horizontal Condensation Load Time-Bistory 020.5 Pool Swell Surface Velocity 020.6 4T Test Parameter Matrix 020.7 4T Test 020.8 Pool Swell Waves 020.9 Load Mitigation in Pool 020.10 Plant Specific Application of Pool Swell 020.11 Discrepancy in Figure ';.4-18 Identifica-tion 020.12 Fluid Velocity for Drag Loads 020.13 Bubble Pressure 020.14 Fallback Loads 020.15 Impact Ioad Lesign FWins 020.16 Estimated Pool Swell Parameter Correla-tions l 020.17 Table of Loads l
020.18 Load Combination Time-Histories 020.19 Quencher Data Multiple Regression Analysis 020.20 Quencher Data Base O
1 of 2 Revision 4 - February 1981
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TABLE D-1 (CONT 'D)
O Ntznber Questionta) Keyword Inder to Ouestions 020.21 Quencher Design Loads 020.22 Large Break with SRV Actuation 020.23 Asymmetric Loads 020.24 4T Test Data 020.25 Air Tests 020.26 Primary and Secondary Loads 130.1 SRV Loads 130.2 Load Combination Time-History 130.3 Load Co.tination Probabilities 130.4 Soil Modeling 130.5 Liner and Anchoring 130.' Asymmetric SRV Loads 130.7 Combining SRV and Pool Loads (1) Letter with enclosure, " Request for Additional Information, Mark II Containment Dynamic Forcing Functions Information Report (DFFR) ", to A.W. Wofford, Long Island Lighting Company, from K. Kniel, NRC, dated June 23, 1976.
l Ca> Series 020: Containment Systems Branch Series 130: Structural Engineering Branch (3) Generic responses submitted by letter to J.F. Stolz, NRC,
, from L.J. Sobon, General Electric, MFN-151-79, dated l May 29, 1979 1
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2 of 2 Revision 4 - February 1981
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TABLE D-2 OUESTIONS OF JANUARY 17, 1977EO Number QuestionC2) Keyword Index to Questions 020.27 Inventory Effects on Blowdown 020.28 Wetwell Backpressure 020.29 Vent Thrust Loads 020.30 Vent Lateral Drag Forces 020.31 3D Tests
- 1. Scaling Analysis
- 2. Vent Flow Asymmetry
- 3. AP/AV Spatial Variation 020.32 Incorrect Reference 020.33 Diaphragm Floor Upward Load 020.34 High Vent Flow Loads 020.35 Pool Swell Model l (/)
N- 1. Isentropic Bubble
- 2. Breakthrough
- 3. Test /Model Comparison
- 4. Drywell Pressure
- 5. Compressibility Effects
- 6. Pool Swell Parameter Sensitivity 020.36 Breakthrough Model 020.37 Downcomer Lateral Load 020.38 Multiple Vent Chugging - Statistical 020.39 Boundary Load Parameter Sensitivity 020.40 Multiple vent Chugging - Tests 020.41 50 Percent Design Margin 020.42 Loads on Submerged Structures i
l 020.43 Downcomer Support Loads 1
020.44 Pool Swell Waves and Seismic Slosh O
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l 1 of 3 Revision 4 - February 1981
TABLE D-2 (CONP 'D) i Ntznber Question (a) Keyword Index to Ouestions 020.45 NEDE 13442P-01 Deficiencies
- 1. All Test Data
- 2. Instrumentation 020.46 Run 5101-29 Raw Data 020.47 Ramshead Radial Orientation 020.48 SRV Load Models and Calculations 020.49 Multiple SRV Actuation
- 1. Model
- 2. Pressure-Time Histories 3.. Wall Load Plots
- 4. Test /Model Verification 020.50 SRV Bubble Dynamic Model
- 1. Bubble Formation Efficiency
- 2. Drag Coefficient
- 3. Multiple Bubble Effects
- 4. Pool Boundary Effects 020.51 Bubble Formation LocaM.on 020.52 Ramshead Influence Coefficient
- 1. Parameters
- 2. Linear Superposition
- 3. Nomenclature 020.53 Bubble Frequency
- 1. Equations and Assumptions
- 2. Initiating Transient
. 020.54 Quencher Loads 020.55 SRV Loads on Submerged Structures 020.56 Primary and Secondary Ioads
! 020.57 Additional Pressure Suppression Tests 020.58 Plant Unique Pool Swell Calculations
- 1. Deviations from Model
- 2. Model Inputs
- 3. Parameter Time-Histories
- 4. Comparison with 4T
- O 2 of 3 Revision 4 - February 1981
TABLE D-2 (CONP *D) ~ l Number Question (*) Keyword Index to Questions 020.59 Downcomer Lateral Braces
- 1. Description
- 2. Effects on Pool Swell
- 3. Impact and Drag Loads
- 4. Downcomer Flanges s 020.60 Wetwell Pressure History 020.61 Pool Swell Inside Pedestal 020.62 Pool Temperature Monitor 020.63 Suppression Pool Temperature Limit
- 1. Temperature Time-History
- 2. Instrumentation / Alarm
- 3. Operator Action 130.8 Load Combination i 130.9 Sructural Design 130.10 Basemat Model 130.11 Acceptance criteria 130.12 SRV Structural Response 130.13 Static Equivalent Lateral; Fluid Structure Interaction 130.14 Chugging; Plujd Structure Interaction 130.15 Wall Stiffness Effects
(*) Letter with enclosure, " Request for Additional Information, Mark II Containment Dynamic Forcing Functions Information Report (DFFR) ", to A.W. Wofford, Long Island Lighting Company, from K. Kniel, NRC, dated Jar.uary 17, 1977.
(2> Series 020: Containment Systems Branch Series 130: Structural Engineering Branch (3) Generic responses submitted by letters to K. Kniel, NRC, from R.H. Buckholz, General Electric, MFN-103-80 and MFN-104-80, dated May 29, 1980.
O 3 c* 3 Revision 4 - February 1981 l
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TABLE D-3 OUESTIONS OF MAY 29, 1978(1)
Ntunber Questionca) Keyword Inder to Questions 020.64 Lateral Loads on Dom ;omers 020.65 Vent System Configuration of Vent Exit 020.66 Static Equivalent Load for Downoomers 020.67 Plant Specific Application of Chugging Loads 020.68 Maximum Pool Swell Elevation 020.69 Differential Pressare Uplift Specification 020.70 Drag Coefficient Multiplier 020.71 Code LOCTVS for Drywell Pressure Response 020.72 Impact Load Specification for Small Structures 020.73 Pool Swe'1 Velocity Calcula. tion s 020.74 Chugging Wall Load Specification 020.75 Submerged Structure Chugging Loads (1) Letter with enclosure, " Request for Additional Information, Mark II Containment Pool Dynamic Loads," to A .W . Wofford, Long Island Lighting Company, frcun J.C. Snell, NRC, dated l May 29, 1978.
(2 ) Series 020: Containment Systests Branch (3) Generic responses submitted by letter to J.F. Stolz, NRC, from L.J. Sobon, General Electric, MFN-275-78, dated l
June 30, 1978.
l lO 1 of 1 Revision 4 - February 1981
Ouestion 020.1:
O Q Clarify the statement that containment walls due to condensation no load should be applied to the oscillations. Figure 5-7 indicates that condensation oscillations should be applied to the subnerged wetwell and Section 6.1.9 of NEDo-11314-0 8 identifies the condensation oscillation loading that should be applied to the pool boundary as determined from the PSTP tests.
Response
The suppression pool wall pressure data from Phases 1, II, and IiI of the 4T program has been reduced and is described in detail in 'he report, " Mark II Pressure Suppression Test Program - Phase II and III Tests," NEDE-13468-P, submitted to the NRC on January 4, 1977. Application of this information is described in the memorandum, " Mark II Pressure Suppression Test Program -
Phases I, II, and III, of the 4T Tests," (NEDE-23 678-P) , dated January 1977, and subnitted to the NRC on February 24, 1977.
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020-1 Revision 4 - February 1981
Question 020.2:
., O Discuss the manner by which the mean and maximum horizontal condensation loads should be applied to a single downcomer.
Response
Downcomer loads are discussed in DFFR Revision 2 Subsection 4.3 and Table 5-1. The static equivalent load applied to a single downcomer is 8,800 pounds based on the maximum observed t.est load. No mean value is considered.
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O 020-2 Revision 4 - February 1981
Question 020.3:
Discuss the criteris that are used in the multiple loading of the downcomers due to horizontal condensation loads. Specifically, identify what fractj.on of the downcomera experience a load acting in the same direction and identify and justify t.he load to be used.
Response
Multiple loading of the downcomers is discussed in Subseccion 4.3.2.4 of the DFFR Revision 2. Ir '-f4is section, a methodology is described, based on the application of a probabilistic analysis, for determining what fraction of the downc .rs experience the application of in phase loads. (Also see e .on 020.67a and c.)
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l lO 020-3 Revision 4 - February 1981 l
Ouestion 020.4:-
It is not obvious how the downcomer horizontal condensation loads, loading time interval, and load period were ontained from the test data presented in NEDE-21078P. Provide specific references and a discussion of how the foregoing parameters were obtained, including any statistical analysis techniques that were Used.
Response
The horizontal loads presented in NEDE-21078 were determined from test data collected from the test facility described in Section 3.2 of the report. Prior to the tests, strain gages (SG) and linear displacement transducers (LDT) which were used for defining vent loads were calibrated by applying known static loads to the 24 in. vent. Based on this static calibration of the SG and LTD, test readings from these instrunents could be converted dire-tly into static equivalent loads on the vent.
7tiese loads are sununarized in Section 3 of the report.
As discussed in Subsection 4.3.2.3 of DFFR Revision 2, the maximum load observed during the-whcle test series was specified as the design load for Mark II vents. This maximum observed load was 8,800 lb (static equivalent) and occurred during test number
- 4. No statistical techniques were involved in the definition of this single vent design load specification. The loading time Q
C interval presented in NEDE-21078 is that a significant vent lateral load will occur approximately once every second. This specitication was based on a review of all the recorded test data.
A review of the data summe given in Table 3-6 of NEDE-21078' shows a load frequency sligritly less than one per second; however, the complete data base indicates that a 1 second loading period is more appropriate. The 50 m second loading period specified in the DFFR was based on judgement because no direct reading of loading duration was made during the tests.
The loading specification for multiple vents is based on a statistical technique that uses the observed loading probability distribution. This subject is discussed in Subsection 4.3.2.4 of DFFR Revision 2. The 4T test program has provided significant additional insight on the characteristics of the downcomer loads that occur during steam condensation. The true forcing function ,
appears to consist of a short duration, high magnitude impulsive loading, having static equivalent values significantly less than the 8,800 lb (static equivalent) specified for Mark II design assessment. The 4T results are discussed in NEDO/NEDE-13442P and l NEDO/NEDE-13468P. (Also see Question 020.67a and c.)
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l 020-4 Revision 4 - February 1981 l
i Ouestion 020.5:
The pool swell model discussed in Section 4.4.1 of the DFFR has been used to calculate the water surface velocity associated with the impact pressures presented in Fig. 4.4-24 through 4.4-26.
Discuss the adequacy of the model to conservatively predict the velocity of the pool surface considering the assumptions that the entire mass of water associated with the vent submergence must be accelerated by the bubble-pressure.
Response
The pool swell model described in Subsection 4.4.1 of DFFR Revision 2 will give water surface valocities that are conservative provided no credit is taken for any loss of energy from the dry well air. Several assumptions used in the nradel lend to the credibility of this assertion. These are discussed as follows:
- 1. Following vent clearing, only air flows into the suppression pool rather than a mixture of air and steam.
This maximizes the mass flowrate of the noncondensables and, hence, the resultant pool swell will be maximum.
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- 2. The mass flowrate of air through the vent is calculated based on adiabatic flow with friction. This will tend to maximize the air bubble pressure and, hence, the pool swell velocity.
l 3. The air in the drjwall is isentropically compressed.
This maximizes the - 9. drywell pressure.
- 4. Frictional losses between the water and the confining walls are negligible. This will also lead to higher pool surface velocity.
The net effect of these assumptions is to maximize the water surface velocity calculated by the pool swell model. (Also see Questions 020.71 and 020.73) .
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&aestion 020.6: i O Provide the matrix of tae. 4T tests relative to the- Mark II design.
that will provide data Identify the key pool swell l
l parameters tnat were obtained from the test data. Identify the i range of independent variables that were covered by the test program.
Response
Refer to the generic response.
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O 020-6 Revision 4 - February 1981
4 Ouestion 020.7:
Provide the following additional infonnation related to the "4T"
'd tests:
- 1. Provide a detailed' scaling analysis for those parameters ;
that will not be full scale in the tests. Specify the portions of the pool dynamics transient in which the scaling analysis is applicable.
- 2. Discuss the manner by rhich test data will be applied to specific plant designs. Include in this discussion the influence of vent flow rate (or transient drywell pressure) , vent submergence, and wetwell airspace volume.
- 3. Provide a comprehensive error analysis for the key independent variables measured in the test program.
Discuss how these errors were factored into a determination of conservative dependent variables.
- 4. Discuss the potential influence of the "4T* geometry and configuration on the test results. Specifically, consider the effects of the tank walls on the measurement of the lateral loads and pool surface velocity, and the eff ect on the vent exit (i.e., without
, bolt flange) on the lateral loads and bubble formation.
- 5. Identify and multiple vent tests data that can or will be used to substantiate the unit cell approach used in the *4T" test facility.
Response
Refer to the generic response.
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020-7 Revision 4 - February 1981
Question 020.8:
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Video tapes of tests performed on a vent system similar to the Mark II design exhibited a significant amount of wave formation in the pool following the initial pool swell transient. Discuss the relevance of this phenomenon to the Mark II design, including the origin and anticipated magnitude of loads due to waves.
Response
The wall pressure probes in the 4T test showed periodic variation in pressure, however, at a given location this periodic behavior is a steady state condition. This variation in pressure could ne caused by wave formation at the surface. If the peak to peax amplitude of the measured pressure is attributed solely to surface wave phenomenon, the lateral hydrostatic pressure caused by the waves would be less than 1.0 pai. niis is equivalent to waves that are less than 2.0 feet from crest to trough.
The lateral loading caused by waves of the magnitude mentioned above is adequately covered in the design process by the lateral drag forces due to vent clearing (Section 4.2.1) and chugging (Section 4.2.5.1) .
In addition, seismic sloching of the supt r 'esion pool results in pool surf ace displacements on the order of 4 feet as described in Section 2.3. Therefore, the above loads, wnich are part of the design loads, will bound any ef f ects due to pool surf ace waves.
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020-8 Revision 4 - February 1981 l
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- . Question 020.9
l' i-Discuss the design fea tures of the Mark II ' containment, or I potential design modificacions, which would be used to mitigate pool dynamic loads.
! Pasponse:
heter to the generic response.
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020-9 Revision 4 - February 1981 .
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Question 020.10:
In Subsection 4.4.4 of this report, all of the Mark II plants have been grouped according to key geometric similarities.
Discuss the manner by which the solutions of the pool swell model for each of the test cases are to be applied to the other plants in each class. If the solutions for a test case are to be applied equally to all other plants in a particular class, justify the approach with respect to differences in drywell pressure response and geometry between the test case and other plants in the same class.
Response
The purpose of the grouping of all Mark II plants was to select one typical plant from ec.ch of the three groups and then analyze thesc plants for their pool swell response. The solut2On obtained for each typical plant was not 2ntended to be applied to other plants in the same class. Any specific plant whose drywell pressure response and geometrical paramete::s are different from that of the typical plant shall be analyzed for pool swell by using the analytical model given in Subsection 4.4.1.
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020-10 Revision 4 - February 1981 I
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Question 020.11:
() Subsection 4.4.4 of the report identifies Fig. 4-28 as being the transient suppression pool air space pressure, whereas this figure is apparently the transient bubble pressure. Clarify this discrepancy.
Response
Subsection 4.4.4 of the report has been modified in Revision 2 such that reference to Fig. 4-28 is no longer appropriate.
Reference to Fig. 4-28 is not required for coherence of Section 4.4.4 of DFFR Revision 2.
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020-11 Revision 4 - February 1981
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Question 020.12:
Discuss the manner by which fluid velocity is determined for the computation of drag loads on submerged structures and piping.
Response
The fluid velocities for the computation of the drag loads on submerged structures and piping due to pool swell are determined according to the procedures outlined in the DFFR Revision 2, Section 4.4, and NEDE-21544.P dated December 1976. A 10 percent margin is applied for conservatism.
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O 020-12 Revision 4 - February 1981
Question 020.13:
Subsection 4.4.5.3 of the report indicates that the bubble pressure should be applied as a uniform increase in hydrostatic pressure.
- 1. Justify this approach with respect to potential differential pressures that could be generated across equipment or structures due to bubble propagation through the pool, specifically consider the reactor pedestal and the drywell deck column supports.
- 2. Justify the use of the calculated transient bubble pressure in terms of any relevant test data available from the 4T tests.
Response
4
- 1. Submerged structure loads due to air bubble formation and asymmetric air bubble pressure loads on the pool boundaries are discussed in Sections 4.2.2 and 4.2.3.6, respectively.
These load specifications augment the uniform load case described in DFFR Revision 2 and cover the concerns identified above.
- 2. The use of the calculated bubble pressure is Justified because it compares favorably with the 4T test. This comparison is shown in Fig.6-12 of NEDO/NZDE-21544-P.
O 020-13 Revision 4 - February 1981
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Question 020.14:
Section 4.4.5.4 of the report indicates that fallnack loads are determined assuming tne acceleration under gravity of a two 7hase
. fluid. Discuss the manner by which the density of the two phase mixture is determined. In addition, since the majority of Mark II plant- have an initial wetwell air space height below three times the vent submergence, justify the assumptions of acceleration under gravity with respect to a - ntum exchanges due to frcth impingement on the diaphragm (i.e., rebound veloc-ity).
Response
Refer to DFFR Revision 2.
Fallback loads are discussed in the revised portions of Subsection 4.4.5.4. A density of 1.0 (liquid phase) is conservatively used for fallback loads. Pool swell impact on the diaphragm Iloor does not occur (Subsection 4.4.4) .
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020-14 Revision 4 - February 1981
l Question 020.15:
The report indicates that a 50 percent design margin may be applied to the impact loads determined for a structure. Discuss the criteria to be used in determiniug whethar a derign margin should be applied to a particular load.
Response
The DFFR impact load specification is no longer used in the Snoreham design assessment. Refer to DAR Section 4.2.3.1.
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O 20-15 Revision 4 - February 1'i81
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Question 020.16: l l
Discuss the manner by rnich the material in Appendix 4.4 of the ,
report is to be used. In addition, describe how r.he data points !
used to generate Figures A4-1 through A4-3 were obtained.
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Refer to the generic response. ;
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.020-16 Revision 4 - February 1981 I
guestion 020.t 7 Provide a table which sunnaries each of the loads depicted in Figures 5-1 through 5-16. For each load, specify the experimental data and/or analysis which form the basis for the load. References to the test data should indicate the specific test runs.
Response
Refer to the generic response. l O
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- O 020-17 Revision 4 - February 1981
Ouestion 020.18
( Provide the following clarifications regarding the temporal relationships depicted on the load combination histories:
- 1. How was the 0.7 sec vent-clearing time determined?
- 2. The pool swell event is depicted to occur between 0.7 sec and 0.9 sec. The calculations in Subsection 4.4 indicate that the pool swell event takes approximately 0.6 sec. Clarify this inconsistency.
- 3. How was it determined that condensation load would begin at 4 seconds following a postulated LOCA?
- 4. Discuss the manner by which the loading time is determined for drag and fallback following impact or froth impingement.
Response
- 1. The 0 . 7' second vent-clearing time was determined with '
the model described in NEDM-10320, " General Electric Pressure Suppression Containment Analytical Model,"
March 1971, using the typical plant parameters contained in Table 4-1 of the DFFR.
- 2. Figures 5-1, 2, 6, 7, 11 and 15 have been revised to be O- consistent with the breakthrough time of 0.6 second,
! which is a typical value (Figs . 4-20, 21 and 22). Pool swell begins at 0.7 second, the time of vent clearing, and ends 0.6 second later at 1.3 seconds, the time of breakthrough.
- 3. DFFR Revision 2 does not provide a load specification for condensation oscilliation loads. Refer to DAR Section 4.2.4.2 for the load specification. The exact time of onset for condensation loads is not critical -
it is only important to note that. condensation loads will not combine with loads due to air cleariry or bulk pool swell.
- 4. The loading time for drag and fallback loads is based on free fall height and velocity, as described in Subsection 4.4.5.4 of DFFR Revision 2.
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020-18 Revision 4 - February 1981
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Question 020._19: _
l Provide a multiple regression analysis for the quencher relief '
valve design using the entire data base available.
Response
! Shoreham does not employ the General Electric quen-?.er design covered by the DFFR. Refer to DAR Section 3.2.
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l 020-19 Revision 4 - February 1981 l
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i Question 020.20: ;
I Provide the data base being used for the quencher design evaluation. The data should be in tabular form, listing all sensitive test parameters.
Response
I Shoreham does not employ the General Electric quencher design i covered by the DFFR. Refer to DAR Section 3.2.
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020-20 Revision 4 - February 1981
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Ouestion 020.21:
Provide the design quencher loads to be used and their bases.
l Response:
Snoreham does not enrploy the General Electric quencher design covered by the DFFR. Refer to DAR Section 3.2.
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l' 020-21 Revision 4 - February 1981 i
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Question 020.22:
b The load combinetions to be considered for the design assessment of the Mark II containment are presented in Subsection 5.2 of the report. The load combinations for the large line break do not consider actuation of a single SRV concurrent with a large break.
Consideration of a single active f ailure will result in this load combination. Accordingly, we will require tnat the load combination be considered for the Mark II containment design assessment.
Response
As noted in Section 5.2.4 of Revision 2 of the DFFR, this load combination will be used in the assessment of structures. This load combination and its structural assessment have been included in the load combination table (Tablo 2-2) and the margin tables presented in Section 6.5.2 of the DAR.
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l 020-22 Revision 4 - February 1981
Ouestion 020.23:
m In April 1975, generic questions related to pool swell and SRV loads for Mark II type containments were sent to utilities with Mark II containment. In this letter, we requested that information be supplied to " describe the manner by which potential asymmetric loads were considered in the containment design. Characterize the type and magnitude of possible asymmetric loads and the capabilities of the affected structures to withstand such a loading prorile...".
This information was not supplied in the DFFR. Accordingly, we require that an evaluation be presented of asymmetric load in the Mark II containment. Potential asymmetric loads resulting from SRV actuation and from asymmetries in vent flow should De considered. In addition, provide an evaluation of the capability of the Mark II containment for asymmetric pool dynamic loads.
Response
DFFR Revision 2 includes a load definition for asymmetric SRV discharge (Section 3.2.4) which has been used in the Shorenam assessment. Asymmetric LOCA loads include air clearing and chugging which have been addressed in DAR Sections 4.2.3.6 and 4.2.5.2, respectively.
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020-23 Revision 4 - February 1981
I Question'020.24:
The report provides an analytical evaluation of the pool dynamic loads for Mark II contaimnent. At the April 28, 1976, Mark II meeting dealing with Mark II pool dynamic loads, tne Mark II owners group stated that the 4T tests would provide experimental
{ confirmation of the analytical methods described in the report.
It is the position of the Staff that acceptance of the pool dynamic loads by the NRC Staff is contingent on the NRC review
- and acceptance of the results of the 4T test program and a-comparison of the test daza with the analytical methods described in the report.
Response
vi evaluation of the pool dynamic loads during pool swell for
- v. ark II contaisuments involves the definition of the following parameters: pool swell velocity, peak air bubble pressure, and maximum swell height. The pool surface velocity is used in the
- determination of impact and drag loads while the bubble ipressure is used to determine suppression pool wall loads beneath the pool
, surface. The pool swell model, as described in Section 4.4.1 of i
the DFFR, is used for the prediction of these two parameters.
Tne 4T data is being used as a basis for the maximum pool swell height determination as described in the applications memorandum,
" Phase I, II, and III of the 4T Tests", dated December 1976, and i submitted to the NRC on January 25, 1977. The method used to i define the maximum swell height is described in the response to NRC Question 020.68.
l A comparison of the maximum pool surf ace velocity as measured in the 4T tests with the pool swell model predictions is discussed in the response to NRC Question 020.73. These tests include two
' blowdown orifice sizes (2-1/2 and 3 in. diameter), two vent diameters (20 and 24 in.) , three vent submergences (9, 11, and 13-1/2 ft), and steam blowdowns. Over this entire range of test parameters, the pool swell model conservatively predicted the maximum pool surface velocity. Theretore, the model can be confidently used in the prediction of pool swell dynamic loads l
wnich are dependent on the pool surface velocity. However, for additional conservation, a 10 percent factor has been applied.
i The suppression pool wall pressure loads due to the air bubble as measured in the 4T tests have also been compared to the model predictions. This comparison is described in the response to Question 020.13(2) .
! Further discussions of the pool swell model and a more detailed.
, comparison of the model predictions to the 4T test data are j provided in the report, " Mark II Pressure Suppression Containment-Systems, An Analytical Model of the Pool Swell Phenomenon,"
NEDO/NEDE-21544-P, published in January 1977.
O 020-24 Revision 4 - February 1981
4
, Question 020.25:
' We have not received a detailed description of the test matrix to be conducted ior evaluation of the Mark II pool dynamic loads.
The description of the 4T test program we have received indicates that 4T air tests have not been covered. In the evaluation of pool dynamic loads for the Mark I and Mark 7II contain= ant design, air tests were conducted to provide data for some of the pool dynamic loads. Because of the potential for a high air ,
traction in the vent flow during the early portion of a LOCA, we ,
currently believe that air tests should be conducted as part of l
, the Mark II pool dynamic load test program. :
Response
! Refer to the generic response.
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020-25 Revision 4 - February 1981 l
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Ouestion 020.26:
he report presents a description of a number of LOCA-related hydrodynamic loads without differentiating between primary and secondary loads. Provide this differentiation between the primary and secondary LOCA-related hydrodynamic loads. We recognize that this differentiation may vary from plant to plant.
We would designate as a primary load any load that has or will result in a design modification in any Mark II containment since the pool dynamic concerns were identified in our April 1975 generic letters.
Response
We classification of LOCA-related loads as primary or secondary is summarized in Table 1-1.
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O 020-26 Revision 1 - April 1977
guestion 020.27:
(' The calculated drywell pressure transient typically assumes that the mass flow rate from the recirculation system or steamline is equal to the steady state critical flow rate based on the critical flow area of the jet pump nozzle or steam line orifice.
However, for approximately the first second after the break opening, the rate of mass flow from the break will be greater than the steady state value. It has been estimated that for a Mark I containment this ef f ect results in a temporary increase in the drywell pressurization rate of about 20 percent above the value based solely on the steady state criticel flow rate. The drywell pressure transient used for the LOCA pool dynamic load evaluation, for each Mark II plant, should include this initially higher blowdown rate due to the additional fluid inventory in the recirculation line.
Response
The effects of pipe inventory have been considered as indicated in Section 4.2.3. For a f urther description of the calculation of drywell pressure for input to the pool swell analysis, refer l to the response to Question 020.58, part (2).
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O 020-27 Revision 4 - February 1981
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Question 020.28:
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The importance of the effect of wetwell backpressure on Marx II pool dynamic loads (i.e., pool swell and steam loads) was discussed in the 4T test report NEDE-13442P-01 and in the June 14, 1976 4T test application memorandum. The 4T test matrix including Phases I through III does not include tests that allow separation of pool dynamic effects attributable to vent sutunergence and wetwell backpressure. We require that additional 4T tests, with these parameters uncoupled, be performed for the purpose of developing plant specific pool swell and steam loads.
Response
Refer to the generic response. l O
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O 020-28 Revision 4 - February 1981 l
Ouestion 020.29:
Thrust loads on the vent system of a Mark II containment are O* reaction forces due to vent flow caused by the IDCA pressure transient. These loads would be transmitted to the diaphragm separating the drywell and wetwell volumes through the vent deflectors and the vent deflector supports. Analysis of these thrust loads have not been provided in the DFFR. We require that these thruct loads be investigated. Provide a description of the method of analyses, the magnitude, and duration of this load for each Mark II plant.
Response
Refer to the generic response. Downcomer thrust loads for Shoreham are negligible.
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020-29 Revision 4 - February 1981
Question 020.30:
Significant differences in the pool area / vent area ratio exist frcan location to location within a given Mark II plant. These differences may lead to cross flow and lateral drag forces on the vents during pool swell. Based on the DFFK Section 4.4.7 it would appear that this lateral drag load on the vents would be computed based on the maximum pool surface velocity and the density of water. Confirm this interpretation of the DFFR. In addition, provide the magnitude and duration of this load for each Mark II plant. Alternatively provide justification for not including this load.
Response
Drag loading on downcomer vents during bulk pool swell is described in DAR Section 4.2.3.3. Since the motion of the pool during bulk pool swell is essentially one d1=ensional, and whatever departure there is from one dimensional flow consists largely of water being pushed away from the vents by the nunble rising along the vent axis, no significant lateral loading on the vents during the bulk pool swell phase is expected. During the initial bubble formation phase, however, a lateral component of loading is expected and is included in the Shoreham design assessment as described in DAR Section 4.2.2 and Appendix K.
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020-3G Revision 4 - Fenruary 1981
i Question 020.31' We require that 3D tests be performed to substantiate the pool swell loads. These loads are currently based on a one dimensional pool swell model and single vent 4T tests. The following items should be considered as a part of the 3D test program.
(1) A comprehensive scaling analysis of the test tacility and error analysis of the test data.
(2) A determination of the sensitivity of pool swell loads to assymetries in vent flow loads and the drywell/Wetwell pressure transient.
(3) A determination of the effect of spatial variations or the pool area to vent area ratio within a given plant on the pool swell phenomena.
Response
Refer to the generic response.
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020-31 Revision 4 - February 1981
1 Ouestion 020.32:
O The DFFR includes the statement on Page 4-43 that a typical jet impingement load on the basemat can be computed utilizing the velocity attenuation given in Figure 12.3 of Reference 13.
Clarify this reference since reference 13 does not contain a Figure 12.3.
Response
Refer to the generic response.
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.O 020-32 Revision 4 - February 1981
SNPS-1 FSAR Question 020.33:
The diaphragm pool swell upward load was based on the unheated drywell test Run 33. This test was conducted with a vent submergence of 11 ft. Figure 5-28 in Ref erence NEDE-13442P-01 shows that the diaphragm upward load increases with increasing vent subnergence. The current peak upward design load for the diaphragm does not appear to include sufficient margin for both this etfect and uncertainty in the measured load. Address this concern and provide an error analyses to substantiate the peak upward design load for the diaphragm.
Response
Refer to the generic response. l O
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020-33 Revision 4 - February 1981
Question 020.34:
O The DFFR in Section u.2.2 states that downcomer and pool boundary I loads will not be considered during periods of high steam riow l since the load derived from the 4T tests are lower than .
corresponding low steam vent flow lateral loads. It is our !
position that high steam flow loads should be considered since these loads, in combination with other loads, may be significant.
It was stated in the 4T applications memorandum that no significant downcomer lateral loads were observed at high steam vent flow. However, in NEDO-21078 Figure 3-19 toreign licensee data indicate significant lateral loads at a vent flow of 20.7 lb/fta in tests conducted with an air mixture of 1 percent.
Specification of a high vent flow downcomer load should reflect this data as well as the 4T data. For structures in the pool it is our position that the 14 psi, 4 Hz load derived from PSTF tests should be used. This load should be confirmed by data from the 4T tests.
Response
Refer to the generic response. l m
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a 020-34 Revision 4 - February 1981
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Ouestion 020.35: 1
(' With regard to the pool swell dynamic analytic model described in ;
Section 4.4 of the DFFR, we have a number of concerns. We request modifications and/or clarification of the methodology in response to the concerns listed below:
(1) Assuaption 5 on page 4-16 of the DFFR sets the bubble air temper ature equal to the (isentropic) drywell air temperature. This assumption is unrealistic from a physical standpoint, and whether or not it is conservative is not obvious a priori. It is our position that this assumption should either be replaced by an application of the first law of thermodynamics to the bubble or show that the use of the drywell air temperature results in conservative pool swell calculations.
(2) The point at which breakthrough occurs is crucial in deter-mining the loading conditions experienced by the con +.aLnment structure. It is our position that the evidence presented to date does not provide a rational basis for estimating when this event occurs. We cannot conclude on the current breakthrough model without adequate test confirmation. Thus, we require confirmation of the breakthrough model with test data.
(3) In general, confidence in the pool swell model can only T develop when comparison of theory and experiment shows favorable results. It is our position that, at this time, such demonstration has not been made. We require confirmation of the pool swell model with test data.
(4) Equation (4.12) of the revised version of DFFR differs from its counterpart in the earlier version, Equation (4.4.10) .
The latter is correct if P is interpreted as the instantaneoas total pressure in the drywell. The version presented in Equation (4.12) is correct if P is the static pressure evaluated at inlet conditions. Clarification is requested.
(5) Equation (4.10) does not consistently account for compres-sibility effects between the drywell total conditions and the inlet static conditions. These effects should either be accounted for or show that these effects result in conservative pool swell calculations.
(6) The sensitivity of the pool swell modei predictions to the choice of initial condition (e.g. , initial pool velocity and bubble pressure) and vent friction factor has not been examined. It is our position that a parametric numerical study be undertaken to examine the sensitivity or pool swell calculations to these parameters.
Response
Refer to the generic response. l 020-35 Revision 4 - February 1981
Ouestion 020.36:
O The Mark II containment supporting program as NEDO-21297 identifies in Section II.2.A.1 development of swell described in a pool velocity breakthrough model. Provide a detailed description of this model and an evaluation of this model using the 4T test data. The model should be verified over a range of conditions to reflect the variations in the design between Mark II plants.
Response
Refer to the generic response.
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O 020-36 Revision 4 - February 1981
Question 020.37:
O The DFFR in Section specification during low4.3 states steam that flow the downcomer is 8,800 lb. The lateral load basis for this specification is the foreign licensee data reported in NEDO-21078. It is our position that these data are not directly applicable for Mark II olants. Accordingly, we require a clear demonstration that this design it,ad represents an upper bound when all the loads are derived from the 4T test program.
Response
Refer to the generic response.
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020-37 Revision 4 - February 1981
i Question 020.38:
O Provide a description of the analytical efforts described in the U 4T test applications menorandum Section 6.0 to investigate the statistical nature of multiple vent chugging.
Response
4 Refer to the generic response.
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i 020-38 Revision 4 - February 1981
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Question 020.39:
) The 4T test report NEDE-13442P-01 does not provide suf ficient information on pool boundary loads. In the final 4T test report provide a quantitative evaluation of the effect of the following parameters on pool boundary loads:
(1) pool temperatures (2) vent air admixture; (3) vent mass flux; (4) wetwell air space backpressure; (5) downcomer submergence; (6) vent proximity to pool boundary.
The pool boundary design load should consider load sensitivity to '
the above parameters and dif ferences between the 4T test tacility and specific Mark II plant' designs.
Response
Refer to the generic response.
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1 020-39 Revision 4 - February 1981 1
Ouestion 020.40:
O A preliminary uniform and assymetric chugging wall load distri-bution for the Mark II systems was provided in Section 6.0 of the 4T test applications memorandum. This load was developed from 4T test data. The 4T test represents a unit cell with a single downcomer. We require that the boundary loads be based on steam tests which include both single and multiple downcomers.
Response
Refer to the generic response.
Note that the information identified in the generic response has been submitted and demonstrates a multivent multiplier less than 1.0; that is, multivent boundary loads due to chugging are less than the corresponding single vent boundary loads.
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O 020-40 Revision 4 - February.1981
Question 020.41:
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' In NEDO-21297, the Mark II containment supporting program report,Section III.2.A.4.a, it is stated that the applicability of PSTF data to Mark II geometry and structures is provided in NEDE-134257 and NEDC-20989-2P. This information does not appear to have been provided in these reports. Ne require that you provide this information. In addition, provide the basis for the 50 percent design margin applied to impact loads as described in Section 4.4.6.1 of the DFFR.
Response
Refer to the generic response.
Note that the subject load definition is no longer used on the Shoreham plant. Shoreham complies with the Lead Plant Acceptance Criteria as discussed in Appendix A of the DAR.
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020-41 Revision 4 - Feburary 1981
Question 020.42:
O For water whether simply imp m loading of s tructures, one should consider it is necessary to specify the actual loading history the total impulse. It the loading history is needed, the or DFFR (NEDO-21061-Rev. 2) proposes the use of impact pressure correlations (Figures 4-34, 35, 36) and pulse duration (Figure 4-
- 37) corresponding to PSTF conditions (NEDE-13425P) . Both parameters depend on the length of target and the shape of the approaching pool. Provide the basis that allows one to assume that these conditions are the same in an actual Mark II pool and the PSTF.
For flat targets in the range of 13-20 inches, the total impulse due to water impact, as calculated from the pressure correlations (Figure 4-36) and pulse duration (Figure 4-37) in the DFFR, is not conservative compared to PSTF data. For exangle, for 20 in.
I beams, the Mark II impulse is only 60 percent of the PSTP data (as determined from Figure 6-8 NE.DE-13426P) . This non-conservatism eliminates the 50 percent design margin used by GE to specify the design loads.
Response
Refer to the generic response.
Note that the subject load definition is no longer used on the Shoreham plant. Shoreham complies with the Mad Plant Acceptance O Criteria as discussed in Appendix A of the DAR.
O 020-42 Revision 4 - February 1?81
Question 020.43:
Justify use of the PSTF impact data for cylinders and I beams associated with the downcomer lateral support system. Show that this data which was ontained from tests on simple geometries applies to the structures comprising a typical downcomer support system.
Response
Refer to the generic response.
Note that the subject load definition is no longer used on the Shoreham plant. Shoreham complies with the Lead Plant Acceptance
- Criteria as discussed in Appendix A of the DAR.
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l 020-43 Revision 4 - February 1981
i Question _020.44:
O Table 5-1 and Figure 5-1 through 5-16 in the DFFR provides a listing of the loads and the load combinations to be included in the assessment of specific Mark II plants. This table and these figures do not include loads resulting from pool swell waves following the pool swell process or seismic slosh. We require that an evaluation of these loads be providc6 for the Mark II containment design.
Response
Refer fto the response to Question 020.8 and Section 2.3 for a l discussion of pool swell waves following the pool swell process and seismic slosh, respectively.
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020-44 Revision 4 - February 1981
Question 020.45:
O The 4T test report NEDE-13442P-01 exhibits certain deficiencies which should be corrected in the final version, for example:
(1) Morc extensive presentation of measured results should be included in the final report. As an example, the data given in Figure 5-15 should be provided for all test runs.
(2) More detailed description in terms of contiguration, principle of operation, calibration, orientation and location of instrumentation should be included in the final report.
, Responset:
Refer to the generic response.
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020-45 Revision 4 - February 1981
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Question 020.4ti:
Provide raw data generated during a selected 4T test run. Signal traces of the conductivity probes are of particular interest, but wetwell and drywell pressure histories and pitot-static probe traces should also be provided. Both short term and long term histories should be included. The specific run selected for this purpose is Run 5101-29.
Response
Refer to the generic response.
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020-46 Revision 4 - Feburary 1981
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Queation 020.47:
Figure 3.3 Type 2 shows the ramsheads-oriented radially toward the containment wall. The bubble discharged from the ramshead directed toward the boundary may behave differently from the bubble discharged from the ramshead oriented tangentially or. in parallel with the boundary. Since the experiments for the SRV tests such as Quad City and the Monticello tests have been
- performed for the ramshead oriented in parallel with the boundary, discuss and justify the applicability or the test data for ramshead directed toward the boundary.
Response
Refer to the generic response. l I
l b) 020-47 Revision 4 - February 1981
Question 020.48:
Provide a brief description and the name of the cociputer code
- used for the S/R valve load calculations. Include an analysis based upon the following input cata:
(1) Parameters given in Table .2-4 of the topical report lEDE- '
21062-P.
(2) Bubble formation efficiency = 0.1 ,
!' (3) Locations of the pressure transducers I;o. 1 and No. S~as shown on Figure 2-7 of NEDE-21062-P.
(4) Compare the calculated results to those in NEDE-21062-P and.
justify any differences.
Response
Refer to the generic response. l 1
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020-48 Revision 4 - February 1981
s Question 020.49:
('~T Provide a transient analysis of the vent clearing, pool dynamic, t _)
s and bubble pressure phenomena as a result of SRV multiple actuation. Include the following:
- 1. Descriptions of the analytical model, including all assumptions and equations.
- 2. Graphs showing _the vent clearing time and pool dynamic bubble pressure as a function of sequential a ctuation .
The number of sequential actuations should be large enough to clearly indicate that the bubbie pressure due to multiple' actuation has reached the maximum value.
- 3. Graphs showing the peak wall pressure, positive as well as negative, as a function of the sequential actuation of the relief valves.
- 4. Verification of the analytical results by comparison with experimental data.- It the experiments were conducted in a different configuration and/or in a different geometry of suppression pool, justification of applicability of the experimental data to the SRV system for each plant should be provided.
Response
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1 020-49 Revision 4 - February 1981
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Question 020.50: l
} Provide justification for the assumptions used in the SRV bubble s_/ dynamic model. Include the following: ,
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- 1. A detailed discussion of the development of bubble formation efficiency. It should De noted that the oubble tormation efficiency could be a runction or air and water temperatures, air pressure, pipe size, pool geometry, submergence, and the degree or air and steam mixing. Therefore, this empirical correlation developed from some particular test data may not be universally applicable.
- 2. Justification for using a drag coefficient of 2.5 for computing. nubble depth.
- 3. Justification for assuming that the dynamics ot a bubble are not affected by the presence of other bubbles.
- 4. Justitication for assuming that the pool boundaries do not affect the motion of the bubble and the discharge rate of air during the process of bubble tormation.
Response
Refer to the generic response.
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v 020-50 Revision 4 - February 1981
4 Question 020.51:
i Tne analytical model assumes that the bubble will be tormed at a point 4 ft from the exit ot the ramshead. It is noted tnat this assumec bubble initial position was derived from Quad City test data. Therefore, it should be treated as an empirical i correlation rather than a constant. Discuss and justify the applicability of this empirical correlation for. the- Mark II containment.
Response
Refer to the generic response.
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020-51 Revision 4 - February 1981
N Question 020.52:
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Provide the 'following additional information ~ on using the influence coefficient method for ramshead loads computation:
- 1. Discuss and Justity analytically and experimentally the selection of the influence parameters.
- 2. Discuss and justify analytically and experimentally the use of the linear superposition principle ior computing the ramshead loads.
- 3. The nomenclature for those variables shown on Table 3-4.
Response: '
Refer to the generic response.
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O 020-52 Revision 4 - February 1981
t Ouestion 020.53:
'>rovide 'a detailed description of the computational method of bukU.e frequencies due to multiple valve actuation. Include the fc11owing information:
- 1. All equations and assumptions used:
Response
For part 1, refer to the generic response. The response to part 2 is given in NEDE-23983, November 1978 where the pressure rise rate (PRR) following isolation for BWR-4/5 is given. The most probable PRR is 70-80 psi /sec with a maximum of 130 psi /sec.
The maximum PRR is used in the Shoreham assessment to minimize the time between the sequential actuations.
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- O 020-53 Revision 4 - February 1981
4 Question 020.54:
DFFR Section 3.3 presents the quencher loads based on the statistical method described in GESSAR-238 NI Appendix 3B, Amendnent 43. As a result of our review, however, we tind this statistical method is not applicable for the Mark II containment because some of the key parameters, such as the air volume, exceeds the test envelope. Extensive extrapolation of the test data is thus required. tie believe that the current data base is not sutticient to justify the applicability of the statistical method of predicting quencher loads for the Mark II Containment.
Therefore we require additional test data, such as could be provided by the CAORSO test, to demonstrate that the predicted quencher loads are conservative.
Response
Refer to the generic response.
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! 020-54 Revision 4 - February 1981 1
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Question 020.55:
/ \ The computational method described in DFFR Section 3.4 for.
\s- calculating SRV loads on . submerged structures is not acceptable.
It-is our position that the Mark II containment applicants.should commit to one-of the following two approaches:
- 1. Design the submerged structures for. the full SRV pressure loads acting on.one side of the structures; the pressure attenuation law described in Section 3.4.1 of NEDO-21061 the ramshead and Section A10.3 .1 of NEDO-11314-08 for the quencher can be applied for calculating the pressure loads.
- 2. Follow the resolution of GESSAR-238 N1 on this issue.
The applic.at.for GESSAR-238 NI has proposed a metnod presented in the GE report, " Unsteady Drag on Submerged Structures," which is attached to the letter dated March 24, 1976 from G. L. Gyorey to R. L. Tedesco. This report is actively under review.
Response
The method used by Shoreham to. assess submerged structure ~ loads due to SRV discharge is described in Append: 7 K. Appendix ' K is ba sicall) in agreement with Approach 2.
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020-55 Revision 4 - February 1981
Question 020.56:
\ The response to question 020.26 wnerein we requested a
(~s- / differentiation between primary and secondary loads is unacceptable. The original deslyn assessment reports for individual plants with Mark II containa ents specitied substantial changes in Mark II containment structures to accommodate pc sl dynamic loads. We recognize that a specified pool dynamic lcid may not be a primary load on all Mark II plants because ut differences in the design of Mark II plants. However, 11 it is a primary load on any Marx II plant, it should-be treated as such in the generic Mark II pool dynamic load program.
t Based on our preliminary review of the original design assessment reports, the DFFR and the reports submitted to us dealing with the definition of the Mark II pool' dynamic loads we have concluded that the following loads shtmld be view ed as pra. mary loads for the Mark II containment design
- 1. SRV loads for both the ramshead and quencher ~ designs.
- 2. Steam chugging loads including loads on the oowncomers and the pool boundary.
- 3. Pool swell loads including tmpact and drag loads.
Our generic review of these Mark II pool dynamic loads will s consider them to be primary loads unless it can be shown that a Os given load is secondary in terms of structural capability or load magnitude.
Response
Refer to the generic response. l t
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- 020-56 Revision 4 - February 1981
Question 020.57:
r A number of pressure suppression tests will be conducted within the next few years. Results of many of these tests snould be applicable to the Mark II containment design. For cach of the tests listed below discuss the participation or monitoring activities of the Mark II owners group. -
- 1. Japan Atond c Energy Research Institute multivent small scale and full scale 1/18 sector tests.
- 2. Mark I 1/4 scale air, 2 vent full-scale steam, and multivent steam tests.
- 3. German tests
- 4. Livermore air and steam tests
- 5. EPRI 1 vent 1/13 scale Mark II tests and Mark I- scale tests
- 6. LOFT suppression tests
- 7. Mark III rnicivent steam tests.
Response
Refer to the generic response.
l N
L.
020-57 Revision 4 - February 1981
Question 020.58:
. s- Relating to the pool swell calculations, we require tue . f ollowing s ) information for each Mark Il plant:
(1) Provide a description of and justify all deviations from the DFFR pool swell model. Identify the party responsible for conducting the pool swell calculations (i.e. , GE or the A&E) .
Provide the program input and results of Dench mark calculations to qualify the pool swell computer program.
(2) Provide the pool swell model input including all initial and boundary conditions. Show that the model input. represents conservative values with respect to obtaining maximum pool swell loads. In the case of calculated input (i.e., drywell pressure response, vent clearing time) , the calculational methods snould be described and justified. In addition, the party responsible for the calculation (i.e. , GE or the AGE) snould De identified.
(3) Pool swell calculations should De conducted for each Mark II plant. The following pool swell results should be provided in graphical form tor each plant:
(a) pool surface position versus time; (b)- pool surface velocity versus time; (c) pool surface velocity versus position; and
. (d) pressure of the suppression pool air slug and the wetwell air versus time.
(V)
(4) The calculated drywell pressure response and the enthalpy -
flux in the downcomer vent should be compared to the 4T 2 1/2 in. and 3 in. venturi data.
Response
(1) The pool swell computer program (POSH) incorporating the DFFh pool swell model was prepared by Stone 6 Webster Engineering Corporation (S&W) and is described in Appendix H. In order l to qualify tne computer program, pool surface velocity and elevation transients for the three classes of problems presented in the DFFR have been calculated. The results of these calculations are presented on Figs. H-1 through H-b for l comparison with the corresponding plots in the DFFR.
(2) Pool swell model input for SNPS-1 is presented in Tables 4-3 ana 4-4. The drywell pressure transient is calculated by 56W using the LOCTVS computer code described in Section b.2.1.1.3 of the Shoreham FSAR. The vent clearing model used in 14XMVS for containment analysis is conservative with respect to maximizing the vent clearing time and drywell pressure. The effect of pipe inventory in the broken recirculation suction b
020-58 Revision 4 - February 1981
line is calculated using the approach described in Appendix B f- of NEDO-20533. Containment initial conditions are the same
("g) as those in Part II of Table 6.2.1-1 of the Shoreham FSAR used for containment analysis. The drywell pre.=aure transient is calculated using the maximum value c' vent sutunergence to maximize the drywell pressure, nubble pressure, and swell height. Maximum swell velocity is not greatly sensitive to submergence as long as the maximum value of suppression chamber free air volume (low water level) is used. Minimum submergence velocity transients have been used i to calculate loads on structures located near the initial pool surface where acceleration time prior to impact is important.
(3) The following graphical results have been provided:
(a) pool surface position versus time (Fig. 4-9) ;
(b) pool surface velocity versus time (Fig. 4-9) ; and (c) pressure of the suppression pool air slug and the l suppression chamber air space versus time (Fig . 4-8) .
Pool surface velocity versus position can be obtained l from Fig. 4-9.
(4) Refer to the generic response.
O l
l l
l l
l O 020-58a Revision 4 - FeDruary 1981
Ouestion 020.5,9_:
In the 4T test report NEDE-13442P-01 Section 3.3 the statement is made that for the various Mark Il plants a wide diversity exists in the type and location of lateral bracing between downcomers and that the bracing in the 4T tests was designed to minimize the interference with upward flow. Provide the tollowing information for each Mark II plant:
(1) A description of the downcomer. lateral bracing system. This description should include the bracing dimensions, method of attachment to the downcomers and walls, elevation and location relative to the pool surface. A sketch of the bracing system should be provided.
(2) An assessment of the etfect of the bracing system on the pool swell phenomena and drywell pressure response.,
(3) The basis for calculating the impact or drag load on the bracing system or downcomer flanges. The magnitude and duration of impact or drag forces on the bracing system or downcomer flanges should also be provided.
(4) An assessment of the effect of downcomer flanges on vent lateral loads.
Response
(1) A description of the downcomer lateral bracing system is con-tained in DAR Section 1.6 (Figs. 1-6, 7, and 10) .
(2) Refer to the generic response.
(3) The basis for calculating impact and drag loads on the bracing system is given in DAR Sections 4.2.3.1 and 4.2.3.4, respectively.
l (4) The downcomers have no flanges below the suppression pool l
water surface; therefore there is no effect on vent lateral l loads.
l l
O 020-59 Revision 4 - February 1981 1
l l
Question 020.60:
In the 4T test report NEDL -1344 2P Section 5.4.3.2 the statement is made that an underpressure does occur with respect to the hydrostatic pressure prior to the chug. However, the pressurization of the air space above the pool is such that the overall pressure is still positive at all times during the chug.
We require that each Mark II plant provide suf ficient information regarding the boundary underpressure, the hydrostatic pressure, the air space and the SRV load pressure to confirm this statement or alternatively provide a bounding calculation applicable to all Mark II plants.
Response
Tne minimum pressure at te suppression pool boundary resulting from simultaneous SRV discharge and chugging loads occurs at the vent exit plane (assuming the .:naximum negative SRV discharge load is applied uniformly for conservatism) and is equal to 9.6 psig. l The minimum suppression chamber airspace pressure corresponding to the low drywell air content required for significant chugging loads is 21.2 psig. The corresponding hydrostatic pressure at the vent exit plane for minimum vent submergence is 3.5 psi above the airspace pressure or 24.7 psig. The maximum value for the underpressure is obtained by combining the maximum underpressure due to chugging with the maximum underpressure due to ADS actuation using square root of the sum of the squares (SRSS) .
s Applying SRSS to the maximum negative chugging pressure on the v pool boundary from DAR Section 4.2.5.2 (-14 psi) and the maximum l negative pressure on the pool boundary due to ADS actuation from Table 3-5 (-5.6 psi) , a naximum negative pressure of -15.1 psi is l calcu3ated for the underpressure at the pool boundary. The sum of t!.c minimum hydrostatic pressure at the vent exit plane and the maximum underpressure at the pool boundary is 9.o psig. l i
020-60 Revision 4 - February 1981
Ouestion 020.61:
() Significant variations cxist in the Mark II plants with regard to the design of the wetwell structures in the region enclosed by the reactor pedestal. These variations occur in the areas of (1) concrete backfill of the pedestal, (2) placement of downcomers, (3) wetwell air space volumes; and (4) location of the diaphragm relative to the pool surface. In addition to variation between plants, for a given plant, variations exist in some of these areas within a given plant. As a result, for a given plant, significant differences in the pool swell phenomena can occur in these two regions. We will require that each plant provide a separate evaluation of pool awell phenomena and loads inside of the reactor pedestal.
Response
In the case of Shoreham, the only significant difference in parameters affecting pool swell between the region inside the pedestal and the region outside the pedestal is in the pool area -to-vent area ration. The ratio for the region inside the pedastal is 21.0 while that for the region outside the pedestal is 16.2. The average for the entire pool is 16.4. Because of the greater pool are-to-vent area ratio inside the pedestal, this region is subjected to a less severe pool swell transient than the pool in general. However, the downcomer bracing system, the
,O only secondary structure requiring assessment pedestal, is assessed using the same loading conditions specified within the l for the pool in general. In comparing pool swell inside and
! outside the pedestal, the two regions are considered
- independently which tends to maximize differences in pool swell
- response. In reality, the two regions are coupled by means of openings in the pedestal both above and below the water level.
[ However, even using the decoupled, two-region analysis, the pool swell loads differ little from those calculated for the single-region analysis presented in the DAR. The maximum pool swell v elocity outside the pedestal would increase by 0.2 f t per second (approximately 0.5 percent) and the differential air bubble pressure on the pedestal, a loading condition not previously specified, would be approximately 2.5 psid acting i inward. The effects resulting from applying a uniform 2.5 psid l differential pressure on the thick-walled concrete pedestal, i
acting to load the concrete in compression, are considered small.
f 020-61 Revision 2 - September 1977 l
d Ouestion 020.62:
For the suppression pool temperature monitoring system, provide the following additional information:
(1) Type, number, and location of the temperature instru-mentation that will be installed in the pool.
(2) Discuss and justify the sampling or averaging technique that will be applied to arrive at a definitive pool temperature.
Response
Refer to ' DAR Section 10.4 for a complete description of the Shoreham temperature monitoring system.
O I
l O
l 020-62 Revision 4 - February 1981
Ouestion 020.63:
For- limiting the suppression pool temperature, provide the fol-lowing additional information:
(1) Present the temperature transient of the suppression pool starting from the specified temperature limits for the following transients:
(a) Stuck open relief valve (b) Primary system isolation (c) Initiation of auto depressurization system (2) Des cribe the instrumentation which will alert the operator to take action to prevent the pool temperature limit to be exceeded.
(3) Describe the operator actions and operational sequence for those transients stated in Item 1 above. Provide and justify the assumption of time for initiating each action and the corresponding pool temperature.
Response
i Refer to DAR Section 10.2 for a discussion of the Shoreham pool
! temperature response to plant transients involving SRV discharge.
l l
l l
Q b
020-63 Revision 4 - February 1981
Question 020.64:
, The data base from which chugging loads on downcomers was developed indicates that lateral loads were also observed at vent clearing. These loads were as high as 3.5 kips (See Table 3-3 of NEDE 21078-P). Therefore, it is our position that a design load not less than 3.5 kips be specified for downcomers during vent clearing. This static equivalent load should be used for each plant with a vent natural frequency less than 7 Hz. For a vent natural frequency greater than 7 Hz, a higher vent clearing static equivalent load should be specified and justified.
Response
Refer to the generic response. l O
l l
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l lO l
020-64 Revision 4 - February 1981 l
t
Question 020.65:
The data base (NEDE-21078--P) from which the chugging load specification for downcomers was obtained with a vent configuration unencumbered by flanges or other protuberances !
located in the vicinity of the vent exit. It is our position l that these load specifications are not applicable to any Mark II '
plants with vents which are flanged at the vent exit. Either the vent exit flanges should be removed or additional steam tests should be conducted with a vent exit flange. ~
Response
The SNPS vent system does not employ flanges or other protuberances in the vicinity of the vent exit.
O O 020-65 Revision 3 - November 1978
l Question 020.66:
The static equivalent load for a downcomer depends on the natural frequency of the downcomer. The current load specification of l 8.8 kips was obtained in a test facility with a downcomer natural frequency of about 7 Hz. This load has not been demonstrated to be conservative for downcomers with a higher natural frequency.
For a vent natural frequency greater than 7 Hz, a higher lateral i load should be specified and justified. Additional information is needed to establish a static equivalent load for downcomers with a natural frequency greater than 7 Hz. In addition, we require that each Mark II plant provide an evaluation of the downcomers utilizing the dynamic forcing function in Task A.13 in the Mark II supporting program as confirmation of the static equivalent load evaluation. The static equivalent and the dynamic loads for the downcomers described above .are based on tests with downcomer diameters of 24 in. or less. Additional
- information will have to be provided to establish lateral loads for downcomer with a larger diameter.
- Response
Refer to the generic response.
l Shoreham uses 24 in. downcomers with a maximum vent system natural frequency of 10.7 Hz.
O As noted using the in DAR Section 4.2.5.1, a dynamic single vent analysis generic Task A.13 method has been completed for Shoreham with acceptable results. A multivent dynamic analysis is in progress.
i lO 020-66 Revision 4 - February 1981
l Question 020.67:
Ref erring to Section 4.3.2 of DFFR - (NEDO-21061-P, Rev. 2) .
- a. It is noten that the force magnitude distribution employed for the probabilistic analysis of multiple downcomer loading is taken from Table 3-6 of NEDE-21078-P. These data were obtained during steam blowdown with significant air admixture (tests 5 and 7) .
Thus, they do not correspond to the " worst
- loading case (On air admixture) which yields the 8.8 kip maximum lateral load specification. We require that the multiple downcomer loading be modified so as to be consistent with this worst-case distribution.
- b. Since the direction of the combined loads from multiple downcomers is arbitrary, assumption 2 of the analysis is unjustified. We require that the magnitude of the resultant of all forces be employed to define multiple downcomer loads. The analysis and results (Figure 4-10 and 4-10a) should be modified accordingly.
- c. The results shodn in Figure 10-4-a implies that a single downcomer will experience an infinite loading. This is obviously incorrect and suggests that the referenced figure is in error. Provide a corrected version of this figure.
Response
i Refer to the generic response.
l l
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020-67 Revision 4 - February 1981 i
l w-* - - -wei--w-w--w---,w- +-ge -w-g - - - yy- - - *- m-- - - - - -
Question 020.68:
O Based on our review -of NEDE-13468P) and the dated January, the 4T test report (NEDE-13442P-01, Phase I, II, III Application Memorandum 1977, it is our position that the specification for maximmn pool swell elevation account properly for observed trends with submergence and state of the blowdown fluid. We-require that the maximum pool swell elevation specification consist of the maximum of either 1.5 times submergence or that predicted by the pool swell analytical model using a polytropic exponent of 1.2 for wetwell air compression.
Response
For SNPS, the plant el corresponding to 1.5 vent submergence is 40.5 ft. The maximum initial pool surfhce el is 27-0 with the vent exit p} sne at el 18-0.
The maximum swell height for SNPS has been defined independently of the pool swell model by using the pool curface el corresponding to the maximum drywell floor uplift differential pressure as described in the MK II generic response to this question. The drywell pressure at the end of pool swell for SNPS is 46.5 psia. The maximum uplift differential pressure is 2.5 psid (refer to the response to Question 020.69) . Therefore, tne maximum wetwell air-space compression due to pool swell would be 49.0 psia. Assuming a polytropic ccropression in the wetwell air-space with an exponent of 1.2, the maximum pool surface el
_ corresponding to a wetwell air-space pressure of 49.0 psia is plant el 47-0 ft (2.2 vent submergence) . This is the maximum i swell height to be used for SNPS design assessment. It greatly exceeds the largest pool surface displacement observed at 4T l (a pproximately 1.6 vent submergence for Run 31) and demonstrates-applicability of the generic swell height specification to plants with large break areas and small values of submergence.
To provide additional assurance that the method outlined in the generic response to NRC Question 020.68 for calculation of maximum swell height is indeed conservative, ccmparisons have been made with 4T Phase I Tests 27, 28, 29, 30, 31, and 35 and with Phase II Tests 22, 36, 37, and 58. The 4T geometry is represented on Fig. 020.68-1 (Proprietary) with the appropriate Shoreham elevations provided for comparison. Additional gecanetric and thermodynamic comparisons between Shoreham and 4T are provided in Table 020.68-1 (Proprietary) . Tnese comparisons demonstrate that 4T is representative or Shoreham.
To provide a basis for comparison of measured and calculated l
swell heights for 4T, 4T was modeled using the Stone S Webster computer codes LOCTVS for the drywell pressure transient and POSH (SSW implementation of the PSAM) for pool swell. The geometric modeling is straightforward and needs no elaboration. A reasonable lower cound for vent equivalent length was used O to 020-68 Revision 4 - February 1981
maxe the - drywell pressure prediction less conservative than one would find in an actual design calculation, thereby making t a O comparison of measured and calculated swell heights for 4T somewhat less favorable than that expected for a real plant.
i Mass and energy release rates for vapor blowdowns (Runs 22, 27, 28, 29, 30, 31, 35, and 58) were based on saturated steam Moody critical flow with fL/D = 0.0 - _ standard for containment
- analysis. The source pressures were taken from Fig. 5-1 of NEDE-13442P.
i l For Run 31, a slug of wa.ter was initially present in the blowdown line as described in . tion 5 of NEDE-21544-P, resulting in -a liquid blowdown at 3 beginning of the test. In the LOCTVS simulation of Run 31, L initial mass and energy release rates were cased on saturated liquid rather than saturated steam and i were assumed constant at the initial steam generator pressure
- until the liquid was exhausted. Subsequently, vapor blowdown
rates were calculated as described above.
! Runs 36 and 37 were run entirely with _a liquid blowdown.
Modeling a liquid blowdown is considerably more complex than modeling a vapor blowdown due to the potential change of phase at tne blowdown source during depressurization. 'Ihe IOCTVS reactor -
- coolant system model was used to calculate the steam generator response and blowdown quality for 4T in the same manner as would be done in the analysis of a real plant. Liquid blowdowns-typically result in a more conservative calculation ot drywell iO U pr essure than vapor blowdowns for three reasons: (1) increasing conservatism in the Moody flow model with decreasing quality, (2) a conservative treatment of the phase change in the drywell which does not appear in a vapor blowdown case, and (3) greater '
conservatism in the vent flow model due to entrainment. It should be noted that the pool swell design basis accident (DBA) tor Shoreham is a liquid break and that calculation of swell height based on th.s calculated Shoreham drywell pressure carries with it these additicaal conservatisms.
A comparison of 4T Phase I measured vent clearing times and short-term peak drywell pressures witn those calculated with LOTVS is provided in Table 020.68-2 (Proprietary). All Phase I tests were vapor blowdowns except as noted. These results show LOCTVS to be somewhat conservative as expected, but generally in good agreement with test data.
The pool swell analysis for each of the 6 Phase I tests and 4 Phase II tests was carried out using the " applications assumptions" of NEDE-215 4 4-P , -- Section 6.7. The results are provided on Figs. 020.68-2 (a and b) through 020.68-11 (a and b) .
The "a" denotes'the calculated drywell and wetwell pressure as a
. tunction of time and the "b" denotes the calculated pool surtace el as a function of time for each run. The drywell pressure at the end of pool swell was taxen from these analyses and is presented _ in Table 020.68-3 as the basis for the swell height O
020-68a Revision 4 - February 1981
calculation. Table 020.68-3 also provides the "F-factor" value t
and the APUP calculated from the expression APUP = 8.2 - 44 F as given in NRC Question 020.69. Since this expression is based in part on pool and drywell temperature corrections typical of conditions found in MK II plants, it is unreasonable to apply it to 4T per se, and doing so would produce a more favorable comparison of measured and calculated swell heights than would be warranted.
To correct the calcualted A PUP for initial pool and drywell tz:mperatures typical of 4T, 2.3 psi has been subtracted from each calculated value, 0.6 psi to account for pool temperature and 1.7 psi to account ror drywell temperature. In no case would aPUP be less than 2.5 paid. The corrected aPUP values for 4T are provided as the fourth entry in Table 020.68-3.
Tne final two entries in Table 070.68-3 are the wetwell pressure (summations of drywell pressure at the end of pool swell and the corrected a PUP) and the swell height assuming a polytropic compression of the wetwell airspace to tne given pressure with x = 1.2. These calculated swell heights are compared with 4T conductivity probe data on Figs. 020.68-12 and 13 (Proprietary).
The calculated points plotted are those determined from the wetwell pressure; but recalling that the proposed method also include a minimum of 1.5 vent submergence, it should be noted that in applying the method, all 11 ft submergence runs would be shifted horizontally to the ($) - line and all 13.5 ft submergence O runs would be shifted horizontally to the () -- line if tney initially lie to the left of the appropriate line.
These data show that the proposed method conservatively predicts the actual maximum swell height measured at 4T with the exception of a single data point, Run 35, where the actual swell height exceeded the prediction by approximately 6 in. 4T Run 35 is a 9ft submergence, small steam break, small pool area (Phase I) run. When comparing 4T Run 35 to Shoreham, it should be recognized that: (1) Shoreham's DBA is a large liquid break which would produce significantly greater conservatism in the prediction of swell height than the small steam break ot Run 35, (2) the Shoreham pooi area to vent area ratio is closer to that of 4T Phase II than to that of 4T Phase I, and more conservatism exists in the method for larger pool area to vent area rataos, (3) the 4T tests are single cell tests and the results of the EPRI subscale single cell /multivent test comparisons are expected to show that single cell configurations will overpredict swell height when compared to multivent, (4) wetwell heat transfer during pool swell at 4T would be expected to exceed that of an actual plant because of the greater boundary surface area to wetwell volume ratio and, therefore, a k = 1.2 is less likely for an actual plant tnan for 4T, and (5) a similar 4T Phase I run (Run 2ti) was conservatively predicted by the method. On tne casis of the above, it is concluded that the generic metnod will O
020-68b Revision 4 - February 1981
J, i !
- j. conservatively predict maximusa swell height for all Mk II plants,-
i including Shoreham. t i
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020-68c 1 Revision 4 - February 1981
_ _ _ - . . _ , . . - - . _ . - . _ - _ _ _ _ _ _ _ _ _ . , _ _ _ _ _ - . . . . . _ _ _ . _ , _ _ . _ _ . , . _ - , . . . . . - . - . . . ~ . - _ .
J TABLE 020.68-1
. COMPARISON OF GEOMETRIC AND THERMODYNAMIC CHARACTERISTICS OF 4T AND SHOREHAM l
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i i
l PROPRIETARY _- See Proprietary Supplement to this' Report 4
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! COMPARISON OF MEASURED'AND CALCULATED
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TABLE 020.68-3 RESULTS OF APPLYING THE PROPOSED METHOD OF THE 4T TESTS Run DW Pressureu3 F- APupu3 APup'(*3 WW Pressureu3 Swell (*D h (Psian Factorta) (Psil (Psil (Psian Heioht fft) 22 34.0 0.195 2.5 2.5 36.5 37 27 32.6 0.082 4.6. 2.5 35.1 38 28 28.0 0.063 5.5 3.2 31.2 35 29 29.4 0.059 5.6 3.3 32.7 37 30 31.6 0.053 5.8 3.5 35.1 39 31 31.9 0.086 4.4 2.5 34.4 37 35 27.6 0.063 5.5 3.2 30.8 35 36 39.4 0.163 2.5 2.5 41.9 40 37 48.8 0.235 2.5 2.5 51.3 42 58 34.0 0.195 2.5 1.5 36.5 37 O 3 At end of pool swell - Figures 020.68-2a through 020.68-11a
(* ) Defined by NRC Ques' ion 020.69 u > Calculated as outlined in NRC Question 020.69 6* Corrected for pool and drywell temperature characteristic of 4T u) Sum of DW Pressure and Pup *
(*) based on polytropic compression with K = 1.2 1 of 1 Revision 4 - February 1981
J l
O THIS FIGURE CONTAINS PROPRIETARY INFORMATION 8
d FIGURE O20.68-1 t
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Question 020.69:
() Our review and analysis of the data base (4T tests and EPRI results) for upward aP suggests that the 2.5 psi specification is inadequate for certain Mark II plants. We require that the current specification be replaced by:
APUP = 8.2 - 44 F (psi) 0<F $0.13 A PUP = 2.5 (psi) F >0.13 where F is a plant unique parameter defined by F = AB.AP.VS VD . (AV) 2 with AB = break area AP = net pool area AV = total vent area VS = wetwell air-space volume VD = drywell volume
Response
The 2.5 paid uplift specification from Rev. 2 of NEDO-21061 has
, been used for SNPS design assessment. The F factor for SNPS is greater than 0.13, and therefore, the SNPS design assessment is 7-s also consistent with the regulatory position stated above.
V O 020-69 Revision 3 - November 1978
. _ . = - _ _ _- .- - . _ - . _ - - . ..
I l
Question 020.70:
i The DFFR (NEDO-21061) methodology for estimating steady state drag loads on subnerged structures is unacceptable for those cases where the structures represent significant blockage to the pool water slug motion. We require that the drag coefficients used to compute the loads be modified according to traditional literature references which take account of tne etfect of blockage.
Response
The method used to correct the steady drag coefficient for the eftects of blockage is given in Appendix K (K.2.3.2.e) .
l i
O O 020-70 Revision 4 - February 1981 4
f Cuestion 020.71:
p
(_,) In the response to NRC questions M020.58 (4) Onu: questions dated January 14, 1977) , it is stated that " Calculations of pool swell for Mark II containments using the analytical model utilize the appropriate calculated drywell pressure response (NEDM-10320) as an input." It is our position that the specification of pressure history is an essential element of the DFFR methodology and that the particular choice cited above has not been demonstrated to be appropriate. To justify such use, we require that pool swell calculations be made for selected 4T tests using drywell pressure response computed according to NEDM-10320 in lieu of the measured drywell pressure histories. The selected 4T tests are the two saturated liquid blowdowns made during the Phase II test series (Runs 36 and 37) . The response (pool swell elevation, velocity, bubble pressure) calculated in this manner should be compared with measured values and with similar calculations made using the measured drywell pressure histories.
Response
For SNPS, the drywell pressure response is calculated using the Stone & Webster computer code LOCTVS rather than the General Electric model described in NEDM-10320. The LOCWS cortputer code is described in Section 6.2.1 of the SNPS FSAR. Comparisons of the measured 4T drywell pressure transient for Run 36 with that calculated by LOCTVS and the measured pool swell response for
[\d ] Run 36 with that calculated by POSH using LOCTVS drywell pressure input is provided on Figures 020.71-1, 2, 3, and 4. As shown in the Mark II generic responce to this question, an even greater degree of conservatism can be demonstrated for Run 37. The generic response also provides comparisons between the 4T pool swell test data and predictions made using the measured drywell pressure histories.
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i FIG. 020.71 - 2 LOCTVS/ POSH PREDICTION VS OBSERVED SURFACE POSITION AT 4T SHOREHAM NUCLEAR POWER STATION-UNIT 1 PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS I
REVISION 3 -NOVEMBER 1978 l
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4T TEST DATA POSH PREDICTION USING LOCTVS DRYWELL PRESSURE 4T CASE 36-2.5 ORIFICE LIQUID BLOWDOWN FIG O20.71 -4 LOCTVS/ POSH PREDICTION VS h-) OBSERVED BUBBLE PRESSURE AT4T SHOREHAM NUCLEAR POWER STATION-UNIT I PL ANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 3- NOVEMBER 1978
b Question 020.72:
The DFFR impact load specification for small structures is O- inadequate. The current load specification consists of a peak pressure-velocity correlation developed from the Mark III PSTF tests. The peak pressure is used in conjunction with an
" average" 7 msec duration to completely define a pressure pulse.
The use of the same 7 msec duration for all situations has not been justified; thus, the current specification is incomplete.
We require that the load specification be modified so as to establish a conservative pulse for all Mark II anticipated situations of target geometry, target size, pool flatness, and pool approach velocity.
Response
As described in Section 4.2.3.1, of the SNPS-1 DAR, the load specification for small structure impact used in the Shoreham design assessment accounts for target geanetry, target size, pool flatness, and pool approach velocity. The load specification is that provided by NUREG-0487 except where the target is neither flat nor circular (e.g., wedge-like). For a typical configurations not covered by NUREG-0487, the methods of Chuang, S.L. " Investigations of Impact of Rigid and Elastic Bodies with Water," Naval Ship Research and Development Center, Department of the Navy, Report 3248, February 1970 have been used.
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020-72 Revision 4 - February 1981
Question 020.73:
Based on our review of 14 Test Reports, application memorandums, and the pool swell analytical model report (NEDE-21544-P) , it is our position that the specification of pool swell velocity according to the analytical model prediction does not provide suf ficient margin to cover uncertainties in the measurements. We estimate this uncertainty to be on the order of 1101.
Accordingly, we require the addition of a 10A margin to the
, values predicted by the analyses for pool swell velocity.
Response
An additional 10 percent margin has been applied to the SNPS pool swell velocity calculation.
O l
lO l 020-73 Revision 4 - February 1981 l
l
Question 020.74:
The current chugging load specification consists of an oscillatory pressure load derived from a conservative chug in the 4T facility. This load includes the FSI related " ring out" of the test walls. The actual load is an impulsive load resulting from collapse of steam bubbles at the exit of the vents. To confirm that the direct application of the pressure signal to containment walls is conservative, additional information is needed. Wall pressure measurements during a conservative chug at the plane of the vent exit should be used to construct an impulse load at the vent exit. The impulse load specification should be used in the coupled fluid-structure analytical model of the 4T facility described in NEDE 23710-P to confirm the conservative nature of the current chugging wall load specification.
Response
The Mark II generic response to this. question supports the current chugging wall load specification as bounding for SNPS design assessment.
v -
0 020-74 Revision 3 - November 1978
Ouestion 020.75:
The supporting program report NEDO 21297 includes an LTP erfort to define main vent condensation submerged structure loads. The current DFFR and lead plant progs ao not include a definition for this load. Eitner provide a load for steam condensation -
submerged strm:ture drag for the STP, or justify def':rring this item to the LTP.
Resoonse:
For SNPS-1, the assessment of sutunerged structure capability to withstand chugging loads is based on a bounding approach. Chug source strength is based on the 4T Bounding Load Report asymmetric wall load specification with the mvi mum source amplitude being that resulting in the v20 psi, -14 psi amplitude wall pressure. Source amplitude is reduced when considering the influence of more than one vent by using a source amplitude multiplier versus number of vents function based on the asymmetric wall pressure distribution. The asymmetric wall pressure distribution is normalized by setting the roaximum positive and negative amplitude equal to unity and replacing the azimuth values with a linear distribution of vents.
Fig. 020.75-1 is the source amplitude multiplier as a function of number of vents. Submerged structure loads using this bounding approach are further discussed in Appendix K (K.2.1.b).
O l G 020-75 Revision 4 - February 1981
O O O L
0.500 -
0.0 0.500 -
j, i i i e e i e i i e # i i i i i e s #
2 6 10 14 18 22 26 30 34 38 42 46 50 54 58 62 66 70 74 78 82 86 90 FIG. 020.75 - 1 SOURCE AMPLITUDE MULTIPLER VS NUMBER OF VENTS SHOREHAM NUCLEAR POWER STATION-UNIT 1 PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION -3 NOVEMBER 1978
Question 130.1:
O Provide in Section 5 a description of the pressure loadings on tne containment wall, pedestal wall, basemat, and structural elements in the suppression pool, due to the various other-combinations of SRV discharges, including the time function and profile for each combination. If this information is not generic, each affected utility should submit the information as described above.
- Request 130.31:
Provide in appropriate sections a description of the pressure loadings on the containment wall, pedestal wall, basemat, and other structural elements in and around the suppression pool due to SRV discharges, LOCA events, and various combinations thereof, including the time function and profile for each combination.
Response
Refer to the generic response.
For Shoreham, the requested descriptions of pressure resulting from SRV discharge, LOCA events, and various combinations thereot are included in DAR Sections 3, 4, and 2, respectively.
O O
- Request 130.31 was contained in NRC letter dated August 27, 1976 O which transmitted Round 2 requests review.
related to the SNPS-1 FSAR 130-1 Revision 4 - February 1981
Question 130.2:
In Subsection 5.2 it is stated that the load combination histories are presented in the form of bar charts as shown in Figs. 5-1 through 5-16. It is not indicated how these load combination histories are used. In particular, it is not clear whether only loads represented by concurrent bars will be combined, and it should be noted that depending on the dynamic properties of the structures and the rise time and duration of the loads, a structure may respond to two or more given loads at the same time even though these loads occur at different times.
Also, although condensation oscillations are depicted as bars on the bar charts, the procedure for the analysis of structures due to these loads has not been presented. Accordingly, the description of the method should include consideration of such conditions. Also, for condensation oscillation loads and for SRV oscillatory loads, include low cycle fatigue analysis.
- Request 130.32:
In Section 2.3 it is stated that the load combination histories are presented in the form of bar charts as shown in Figs. 2.3-1 through 2.3-18. It is not indicated how these load combination histories are used. In particular, it is not clear whether only loads represented by concurrent bars will be combined. It should be noted that, depending on the dynamic properties of the structures and the rise time and duration of the loads, a structure may respond to two or more given loads at the same time even thougn these loads occur at different times. Also, although condensation oscillations are depicted as bars on the bar charts, the procedure for the analysis of structures due to these loads has not been presented. Accordingly, the description of the method should include consideration of such conditions. A3so for condensation oscillation loads and for SRV oscillatory loads, describe the mnthods that were used for low cycle fatigue analysis.
Response
I Changes have been made to Figs. 5-1 through 5-16 in the DFFR Rev.
2 to make them consistent with the revised Table 5-1 and other
, appropriate report paragraphs in the DFFR Rev. 2.
l l
The description of, and the method of the analysis for, the condensation oscillation load is presented in Sections 4.2.4.2 and 5.2.4.3 of the DAR, respectively.
Discussion of the fatique analysis, due to the SRV actuations and
- condensation oscillations, is presented in Section 7 of this report.
L
- Request 130.32 was contained in NRC letter dated August 27, 1976 which transmitted Round 2 requests related to the SNPS-1 FSAR review.
130-2 Revision 4 - February 1981
Question 130.3:
In discussing the load factors used for loads in various load O^ combination, the probabilistic approach given on page includes comparisons of various load ccanbination probabilities.
Explain how the load factors and load combinations are established on such a probabilistic approach and how the various orders of magnitude as indicated on page are obtained and provide the load factors and load combinations thus established.
- Request 130.33:
In discussing the load factors used for loads in various load combinations, the probabilistic approach given on Page 5.24 of Reference 1 includes comparison of various load combination probabilities. Explain how the load factors and load combinations are established on such a probabilistic approach and aow the various orders of magnitude as indicated on Page 5.24 of Reference 1 are obtained and provide the load factors and load combinations thus established.
Response
The load combination equations and the associated load factor.3 to be used for the assessment of the containment and its internal l structures are given in Table 5-2 of the DFFR, Revision 2, September 1976. The load factors were established to provide g safety margins equivalent to applicable codes on concrete gj containments and other concrete structuren. In particular, ASME BSPV Code Section III Division 2, ACI-349 and the Standard Review Plan 3.8 were used for guidance in aveloping these load factors.
\
- Request 130.33 was contained in NRC letter dated August 27, 1976 which transmitted Round 2 requests related to the SNPS-1 PSAR review.
130-3 Revision 4 - February 1981
Question 130.4:
O Through the use of figures, describe in detail the soll mode?.ing as indicated in Subsection 5.4.3 and describe the solid fiaite elements which you intend to use for the soil.
- Request 130.34:
In Section 5.1.2, it is stated that solid axisymmetric elements are to represent-the supporting soil. Since soil properties are known to be appreciably different from the static values for relatively large dynamic strain, describe how the variations in soil properties are considered in your finite element analysis.
Also, describe the criteria used to determine the boundary conditions and the overall finite element mesh size.
Response
Description of the model and method of analysis is presented in DAR Section 2.4.2.
O l .
- Request 130.34 was contained in NRC letter dated Augusi. 27, 1976 which transmitted Round 2 requests related to the SNPS-1 FSAR O review.
130-4 Revision 4 - February 1981
Question 130.5:
O Describe the mathematical model which you will use for the liner and the anchorage system in the analysis as described in Subsection 5.6.3.
i tReauest 130.35:
Provide the matnematical model which was used for the analysis of the liner and the anchorage sysv.em as described in Section 7.2.3.2.
Response
The mathematical model for the liner and anchorage system is e presented in DAR Section 7.2.4.
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l
- Request 130.35 was contained in NRC letter dated August 27, 197t.
O' which transmitted Round 2 request related to the SNPS-1 FSAR review.
i 10-5 Revision 4 - February 1981
Question 130.6:
In Subsection 5.1.1.1 it was stated that the Snv discharges could cause axisymmetric or asymmetric loads on the containment. In Subsection 5.4.1, an axisymmetric finite element computer program is recommended for dynamic analysis of structures due to SRV loads, and no mention is made of the analysis for asynenetric loads. Describe the structural analysis procedure used to consider asymmetric pool dynamic loads on structures and through the use of '.igures, describe in more detail the structural model which you tatend to use.
- Request 130.36:
In Section 5.1.5.1 it is indicated that containment response to SRV loads was determined for the case of asymme':ric loads on the containment. In Section 5.1.2 it is stated tha.t an axi-symmetric finite element computer program was used for dynamic analysis of structures due to SRV loads, and no mention is made of the analysis procedure used to consider asymmetric pool dynamic loads on structures. Through the use of figur>3s, describe in more detail the structural model which you intend to use.
Response
The finite element computer program used for analysis of axicpnmetric and asymmetric pool dynamic loads is described in detall in DAR Section 2.4.2.
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- Request 130.36 was contained in NRC letter dated August 27, 1976 O which transmitted Round 2 requests related to the SNPS-1 PSAR review.
130-6 Revision 4 - February 1981
Question 130.7: l O In Table 5-1, load couabinations 4a, Sa, and 7a are not acceptable to the NRC Sta f f. Discharge of a single safety / relief valve nnst be combined with the remaining loads of these tvvahinations. A i load factor of 1.0 on the SRV loads in these combinations is l
acceptable to the NRC Staff.
- Request 130.37:
In Table 2.2-1, load cr==hi na tions 4 a, Sa and 7a are not completely acceptable. Discharge of a single safety / relief valve i must be combined with the remaining loads of these combinations.
A load factor of 1.0 on the SRV loads in these combinations is acceptable.
Response
See the response to Question 020.22.
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- Request 130.37 was contained in NRC letter dated August 27, 1976 which transmitted Round 2 requests related to the SNPS-1 FSAR review.
130-7 Revision 1 - April 1977
.- ._ -= - -
Question 130.8:
Responses to previous SEB questions 130.5 and 13'>.2 are insuffi-V cient. DFFR Tables 2-1 and 5-1 have not prorided any load profiles and time histories. DFFR Figs. 5-1 through 5-16 have no indications of how the load time histories are ccumbined. Provide the infonnation requested.
Response
Refer to the generic response.
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i 130-8 Revision 4 - February 1981
Question 130.9:
Clarify the last sentence; on Page 5-20 of the DFFR. Mill structures be designed using load cosabinations 4a, Sa, and 7a of .
Table 5-2? j Res.wnse:
Refer to the generic response. l
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._-.__.___...____._._______._.._.._.._.__,___..u_.__.____......,____._..._. _ . . _ . . . . , _ _ _ . . _ . _ _ _ . . . _ _ _
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Ouestion 130.10:
It is questionable that the base mat or drywell floor may be modeled as a thin shell as described in DFFR Section 5 . 84 . 2 .
Support this assertion or modify the section to eliminate the thin shell modeling option.
Response
Refer to the generic response. l O
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130-10 Revision 84 - February 1981
Question 130.11:
The reference in DFFR Section 5.5 to use of the strength allow-able of ACI-218-71 is not considered appropriate. The specific strength acceptable criteria should be specified. An acceptable set of such allowable are those incorporated into US NRC SRP 3.8.
Response
Refer to the generic response. l l
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O 130-11 Revision 4 - February 1981
Ouestion 130.12: l Reference is made in DFFR Section 5.4.3 to studies of structural response to SRV load. Provide citations for this reference and !
where such studies are not readily available, copies are i requested.
Response
Refer to the generic response.
Studies have demonstrated that building response for Shoreham is relatively insensitive to the use of dynamic strain-dependent soil properties.
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130-12 Revision 4 - February 1981
Question 130.13:
The 4T test ; applications as=nrandum states that high magnitude short duration dynamic lateral loads were observed. Provide a description of the method used to convert from a dynamic lateral load to an equivalent static lateral load. In addition, provide a description of the methods used to assess the effect of load structure interaction in the 4T tests and in the various Mark II vent designs.
Response
Refer to the generic response.
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O 130-13 Revision 4 - February 1981
Question 110.14:
O The 4T test applications memorandum states that pool boundary loads resulting from chugging are based on 4T test data conjunction with engineering application techniques to account in for difforences between the 4T facility and the full scale systems. Provide a description of these techniques. In addition, diceuss load / structure interaction considerations given j to pool boundary loads for each Mark II plant.
Response
Refer to the generic response.
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13,0-14 Revision 4 - February 1981
Question 130.15:
The 4T test report NEDE-13442P-01 does not provide sufficient information related to pool boundary loads. The final 4T test report should provide a quantitative evaluation of the etfect of stiffness of the wetwell wall on pool boundary loads.
Response
Refer to the generic response.
Question 130.15:
The 4T test report NEDE-13442P-01 does not provide sufficient information related to pool bc m.% ry loads. The final 4T test report should provide a q - .2 a cive evaluation of the effect of stiffness of the wetwell wall on pool boundary loads.
Response
Refer to the generic response.
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! 130-15 Revision 4 - February 1981 1
_. . . _ _ _ . _ _ . . _ . . - - . ._ _ .- _ o
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i APPENDIX E FUNCTIONAL CAPABILITY CRITERIA FOR MARK II PIPING O
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Revision 3 - November 1978
i TABLE OF CONTENTS O
\~2 Section Title Pace E1.0
SUMMARY
E-1 E2.0 NOMENCLATURE E-2 E
3.0 INTRODUCTION
E-4 E4.0 FUNCTIONAL CAPABILITY CRITERIA E-5 E4.1 Class 1 Piping E-5 E4.2 Class 2 of 3 Piping E-5 E5.0 EASIS FOR THE CRITERIA E-7 E5.1 NUREG/CR-0261 Report E-7 E5.2 Straight Pipe E-7 E5.3 Curved Pipe or Eutt-welding Elbows E-8 E5.3.1 Pressure Term Index, Bs E-8 E5.3.2 Moment Term Inde:.c and (0.75i) = Br E-8 ES.4 Eranch Connections and Tees E-10 E5.4.1 Branch Connections, Class 1 Piping E-11 E5.4.2 Reinforced and Unreinforced Fabricated Tees, .
Class 2/3 Piping E-11' E5.4.3 Eutt-welding Tees, Class 1 Piping E-11 E5.4.4 Butt-welding Tees, Ciass 2/3 Piping E-12 E5.5 Other Products / Joints E-12 O E5.6 E5.6.1 Limits for Do/t > 50 Temperature Effect on Material Properties E-13 E-13 E5.6.2 Internal Pressure Effect E-14; E5.6.3 Products / Joints Other Than Straight Pipe E-14 E5.7 Dynamic Ef fects E-14 E5.8 Summary E-16 l E
6.0 REFERENCES
E-17
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LIST OF TABLES Table Title E4-1 B-Indices for Functional Capability Evaluation.
Do/t 550, for Use in Equation (9) of NB-3652, Class 1 Piping E4-2 Values to be Used in place of (0.751) for Functional Capability Evaluation D /t 150, for use in Equation (9) of NC-3652.2, Class 2 or 3 Piping l ES-1 Sutmaary of Limit Moments on Branch Connections from Table 9 of NUREG/CR-0261 l ES-2 Test Data on Butt-Welding Tees, From NUREG/CR-0261 O
E-il Revision 4 - February 1981
LIST OF FIGURES O Fiqure Title ES-1 Test Data on Butt Welding Elbows with Arc Angle,a, l of 908 or 1808, from Table 4 of NUREG/CR-0261 C5-2 Load - Displacement Plots, Elbows Identified as l (13)10 and (13)11 in NUREG/CR-0261 ES-3 Test Data on Straight Pipe With D/t >50, from Table 2 l-of NUREG/CR-0261 ES-4 Typical Feedwater Fluid Transient Forces ES-5 Typical Main Steam Trancient Steam Hammer Forcing Function ES-6 Typical Piping Elastic Response to OBE ES-7 Piping System Subjected to Pressure Transients, From Ref. (4)
ES-8 PortionofPipingSystemwithHighestCalculatedstresses,l From Ref. (4)
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E1.0
SUMMARY
/~r This Appendix provides " Functional Capability Criteria" for evaluation of essential piping in Mark II nuclear power plants in order to expedite licensing by addressing a theoretical NRC concern not experienced in operating power plants. The criteria were esta blished so as to be conservative and to assist in assuming maximum reliability of the piping considering all aspects of design, fabrication, in-service inspection, and operation.
The criteria are contained in pages E-S and 6. The criteria are structured to make maximum use of the equations and detinitions contained in the Code (1). However, the functional capability criteria are not intended to substitute for or supersede any requirement of the Code.
The basis for the criteria is described in pages E-7 through 16.
The criteria are based, in large part, on the conservative approach contained in NUREG/CR-0261ca); i.e. on the single-hinge, l limit moment concept with little or no consideration of strain hardening or dynamic ef fects. Recommendations or concepts given in NUREG/CR-0261 for B indices are used. For elbows with a, <g oo , excess conservatism has been avoided by using recommendations given in Reference 3. The criteria uniformly use a right-hand-side Ibnit of 1.5S or 2.0Sy rather than the less applicable factors on S, or as used in the Code for A, B, C, Q(N or D limits.
For Do/t >50, the allowable moments are decreased by increasing the B, indices and equivalents of (0.751) . This is based on test data on straight pipe at room temperature with, for ferritic materials, a temperature factor based on ratios of allowable longitudinal compressive stresses from Reference 1.
Dynamic effects may make the criteria very conservative wnen used for conditions where the loadings are dynamic in nature.
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E2.0 NOMENCLATURE B: = pressure loading index B2 = moment loading index B
2b
=
moment loading index, moments applied to branch pipe B2 r = moment 1 ading index, moments traversing run pipe cba = moment loading index, moments applied to branch pipe, see Code Table NB-3682.2-1 Car =
moment loading index, moments traversing run pipe, see Code Table NB-3682.2-1 D = mean diameter of pipe, of run pipe for branch connections and tees Do = outside diameter of pipe, of run pipe for branch connections and tees d = mean diameter of branch pipe h = 4tR/D2, elbow parameter i =
stress intensification f actor, see Code Fig. NC-3673.2 (b)-1 Kd = Sm/Sy KS = Sh /Sy M = moment M* = M /Me ., Me. calculated by Equation (4) of NUREG/CR-0261
=
Me test-determined limit moment at =2 e
=
Md moment applied to branch, see footnote (5) to Code Table NB-3682.2-1 M en =
Code allowable moment, Class 1, Eq. (9) of NB-3652 with right-hand-side limit of Sy M e2 =
Code allowable moment, Class 2, Eq. (9) of NC-3652.2 with right-hand-side limit of Sy Mrs =
functional capability criteria allowable moment, Class 1, with right-hand-side limit of S y Mp2 =
functional capability criteria allowable moment, Class 2/3, with right-hand-side limit of S Mi = resultant moment (see Code defibitions for details)
ML = test determined limit moment M max = maximum load applied during a test Mx, My, M z = set of moments applied on elbow, see Fig. 9a of NUREG/CR-0261 M x3, M M y3, 23 = set of moments applied to branch pipe, see Table 13-3682.2-1 P = internal pressure, see pages 2 and 4 for criteria definition R = bend radius of an elbow Rm = mean radius of run pipe, see Fig. NB-3686.1-1 Sn = allowable stress, tabulated in Code Table I-7.0 Sm = allowable stress intensity, tabulated in Code Table J-1.0 Sy = yield strength of material. See pages 5 and 6 for criteria definition. On test evaluations, Sy is the actual yield strength of material used in the test specimens.
Ty = temperature (deg. F) in the criteria T = wall thickness of run pipe, see Fig. NB-3686.1-1 t = nominal wall thickness of pipe (branch pipe for branch connections and tees), see NB-3683.1 O
E-2 Revision 4 - February 1981
Z = section modules of pipes j t = lesser of T or (i) (t) 8 a = multiplier 6f S, to define right-hand-side of Eq. (9) in O- NB-3652 4
ae = arc angle (degress) of an elbow 4 = multiplier of Sh to define right-hand-side of Eq. (9) in NC-3652.2 6 = displacement in test
) 3 , = extrapolated elastic displacement in test
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E
3.0 INTRODUCTION
Functional capability of piping is defined as its fluid-flow capability. The f unctional capability may be bmpaired by the extremely unlikely event of large reductions in the cross-sectional flow area and the criteria given herein were selected to assure that significant reductions in cross-sectional area do not occur.
The criteria given here were selected so as to be conservative when used with an elastic analysis of piping systems. It may be possible to show that lunctional capability is assured by using more sophisticated analysis techniques, such as determining loads to produce a collapse mechanism or by conducting an elastic plastic analysis. The criteria contained herein are not to be construed as prohibiting more sophisticated analysis methods.
The term " conservative" is used herein to mean more-than-adequate for the specific aspect of functional capability. However, excessive conservatism in any one aspect of designing a piping system (e.g. in postulation of loads, combination or loads, combinations of stress, etc) does not necessarily mean that the piping system will be of optimum reliability. txcess conser-vatism may lead to unnecessary snubbers or supports, which must be attached to the pipe; with possible problems at the attachment po ints , reduced inspectability and possibly signiticant adaltional loads due to malfunction of snubbers or to the mass of the snubbers.
The criteria aza contained in the four pages under the heading
" Functional Capability Criteria" . The basis for the cr16eria is contained in the remainder of this document.
The criteria are structured so as to make maximum use of the equations and definitions contained in the Code. However, the tunctional capability criteria given herein are not intended to substitute f or or supersede any requirement of the Code or tuture changes, thereof. The criteria herein are for functional capability and do not depend upon Service Levels (i.e. A, B, C, or D) used in the Code.
The term, " Code", as used in Appendix E, is the ASME Boller and Pressure Vessel Code,Section III, Division 1,
- Nuclear Power Plant Components, 1977 Edition with Addenda up to and including Summer 1978(*). References to portions of the Code are indicated l by identifications used therein; e.g. NB-3652.
O E-4 Revision 4 - February 1991
E4.0 MJNCTIONAL CAPABILITY CRITERIA
() E4.1 Class 1 Piping Equation (9) of NB-3652 shall be satisfied with the following requirements.
(1) The right-hand-side of Equation (9) shall be 1.5S y, for products other than branch connections or tees, and shall be 2.0Sy for branch connections or tees.
(2) The B-indices shown in Table E4-1 shall be used. l (3) For D,/t >50, Be , Bab , and Be r shall be divlided by:
(1.3 -
0.006 Do /t) (1.033 - 0.00033T) for ferritic materials, and by (1.3 -
0.006 Do/t; for other materials.
The indices are not applicable for D o /t >100.
(4) The detinitions given in NB-3652 are applicable except that:
(a) P is the pressure coincident with the moments, and (b) Sy is the yield strength of the product material at the metal temperature, T (deg. F) , coincident with O the occurrence of the loads; trom Table I-2( 1 ) .
(c) o. = arc angle (degrees) of an elbow E4.2 Class 2 or 3 Piping Except as permitted in (5) below, Equation (9) or NC-3652.2 shall be satisfied with the following requirements.
(1) The right-hand-side of Equation (9) shall be 1.5S .
7 (2) The values to be used in place of (0.751) are shown in Table E4-2, except for " branch connection" for which l Class 1 criteria shall be used.
(3) For D o/t >50, the values to be used in place of (0.75i) shall be divided by:
(1.3 -
0.006 D o/t) (1.033 - 0.00033T) for ferritic naterials, and by (1.3 -
0.006 Do /t) for other materials.
The values to be used in place of (0.75i) are not applicable for D o/t >100.
O (4) The definitions given in NC-3652 are applicable, except that:
E-5 Revision 4 - February 1981
(a) g, is the pressure coincident with the moments, and (b) S is the yield strength of the product material at tde metal temperature, T, (deg. F) coincident with the occurrence of the loads; from Table I-2(*) or, )
if the material is not included in Table I-2, trom other authoritative sources, adjusted to minimum extiected yield strength like Table 1-2 data.
(c) a,= arc angle (degrees) of an elbow (S) Piping constructed in accordance with the Code rules for Class 2 or 3 may be evaluated for functional capability using the criteria given herein for Class 1 piping.
When using this alternative, S shall be established as in 4 (b) .
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E5.0 BASIS FOR THE CRITERIA O
V E5.1 NUREG/CR-0261 Report NUREG/CR-0261, " Evaluation of the Plastic Characteristics of Piping Products in Relation to ASME Code Criteria *(2), addresses the Code rules which are for pressure boundary integrity and not necessarily for fur tional capability. However, the report points out that if allowable moments are restricted to limit moments, as defined by the criteria that the limit moment is that moment of which 6 = 26 , then restrictions in flow area will be small (less than 5 perce$t) and functional capability wili be assured. Test data from Reference (22) of hDREG/CR-0261 indicates that reduction in cross-sectional area of elbows at the limit moment Ma never exceed 2 percent. Accordingly, the data and recommendations given in NUREG/CR-0261 are pertinent to and 4
are used as a main basis for the functional capability criteria. ,
NUREG/CR-0261 addresses the adequacy of the Code criteria as judged by the existence of limit moment conditions at some point in the piping system that cause a single " hinge". The report points out that, for gross plastic deformation to occur, a
" collapse mechanism" must be developed by occurrence of more than one " hinge", and gives a simple example where the collapsa-mechanism load is 33 percent higher than the load creating tne 4 first hinge. The report uses theory (for straight pipe) which ignores strengthening by strain hardening. Dynamic effects are r- not taken into account in NUREG/CR-0261; however, as discussed at the end of this chapter, dynamic ef fects should make the criteria very conservative when used for conditions where loadings are dynamic in nature. Accordingly, NUREG/CR-0261 recommendations are based on " conservative" evaluations and the functional capability criteria share that conservatism. As additional data becomes available, the functional capability criteria should be reviewed and modified as appropriate to remove excess conservatism.
E5.2 Straight Pipe The criteria for Class 1 and Class 2 or 3 piping are identical.
Both can be expressed as:
0.5 jul + 1.o lt s 1.5 S y 2t g (1)
This criteria is deemed appropriate by NUREG/CR-02b1; see Recommendation (4) therein. However, for functional capability, we simply use the more significant limit or 1.54, not restricted to some constant times S m or S n. These are hre restrictive criteria than Code C-Limits for SA312-TP304 above 1000F (Class 1) or above 5000F (Class 2 and 3) .
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E-7 Revision 4 - February 1981
E5.3 Curved Pipe or Butt-Welding Elbass The criteria tor Class 1 and Class 2 or 3 piping are identical except for the coefficient of PD /2t, which varies from 0.0 to 0.5 for Class 1 and (because there is no equivalent n' Bt in Equation (9) of NC-3652.2) is 0.5 for Class 2 or 3 piping.
E5.3.1 Fressure Term index, g1 As oointed out in Recommendation (5) in NUREG/CR-0261, iaternal pressure tends to increase the limit moment. These criteria implement the concept that B can be small when the elbow behavior is predominant (which occurs when h is small) but it should be 0.5, the same as for straight pipe, when the elbow behavior is predcrainantly like straight pipe. The elbow behavior is characterized by the parameter h = 4tR/D2 where t is the hall thickness, R is the bend radius and D is the cross section mean diameter. The criteria for B is such that B = 0.0 for h 50.25 (at h = 0.25, B, = 3.28 for a, =900) and B = 0.5 for h = 1.5 (at h 21.5, B, = 1.00, the same as for straight pipe) . Accordingly, the criteria for B recognizes that pressure does not reduce functional capability when h is small and also recognizes that, when the einow behaves like straight pipe, the limit moment may be reduced by pressure.
E5.3.2 Moment Term Index and (0.75i) =B, The criteria for Class 1 and Class 2 or 3 piping are identical h for the coef ficient of Mi/Z. For a, = 900, Recommendation (6) of NUREG/CR-0261 has been used; i.e. Br = 0.67Cr rather than B, =
0.75Cr. However, recognizing,that a, (the elbow arc angle) may oe less than 900 and to remove excess conservatism from this particular aspect of piping evaluation, the criteria makes use of recommendations given in ORNL/Suo-2913/7, "End Effects on Elbows Subjected to Moment Loadings"(3). This report gives the results of an extensive parametric study of calculated clastic stresses and shows that as a decreases below 900, the maximum calculated stress for an "in plane" moment also decreases. ORNL/Sub-2913/7 recommends that the C, index, for in-plane moment [ identified therein as Crr, see P. 23 of Reference 3] be given by:
Cr =
1.95/h 2/3 for a, 2 9 0 0 Cr = 1.7 5/h 0. 5 6 for a,= 450 C, = 1.0 for a,= 0 Linear interpolation with a, shall be used but Cr shall not be less than interpolated for a o = 300 and not less than 1.0 f or any a, .
The criteria uses this recommendation with B, = 0.67C,, but not less than 1.0, except that the provision that Ce shall not be less than extrapolated f or a =300 1s not used. This is because E-8 Revision 4 - February 1981
the. bound at =300 was motivated by concern over poss1 Die
- interaction between two closely spaced welds from a fatigue (s) standpoint and hence is not relevant to functional capability.
It is significant to note that the criteria is based on "in-plane" moment loading; this moment produces higher elastic stresses and lower limit moments than the other two moments making up M i; i.e. out-of plane and torsion. Accordingly, the criteria is conservative if all of My is in plane moment and may be excessively conservative if Mi is mostly out-of-plane or torsional moments. Unfortunately, this possible excess conservatism can not be removed in the simplified approach used in this criteria but could be used in more sophisticated analysis metnods which, as remarked in the introduction, are not prohibited by this criteria.
Figure ES-1 shows the test data on elbows from Table 4 of l NUREG/CR-0261. The criteria, for a, 2900, h 50.25, is:
O f1 B, 51.5S (r/4) Dat I.
or, in terms of the parameters used on Fig. ES-1, recalling that l Br = 1.3/h 2/3,
_s s _ M ,, =
( r/4 x 1.5) ha/3 = 0.906h2/3 (2)
D2t3 = 1.3 Equation (2) is plotted on Fig. ES-1, giving a visual comparison of test data with the tunctional capability criteria for elbows.
It can be seen on Fig. ES-1 that the test data vary with h by roughly h 2/3 as used in the criteria. Recalling that the theoretical limit moment for straight pipe is Mr = DetS y, the fact that the test data M/ (Dat values are less than 1.0 indicate that elbows with h < U. 4 )4 are not as strong as straight pipe. However, the important aspect is that the criteria conservatively evaluates this " elbow-effecta for all but two points which we briefly discuss in the following.
One or the two points slightly (~10%) Delow the criteria is from Reference (23) of NUREG/CR-0261. Reference (23) of hDREG/CR-0261 tests were different in that +M2 was applied first to.a magnitude ot about 1.3 x 106 in-lb, then -M2 was applied to about the same ma gnitude . This cycle was repeated several (~10) times; each time the maximum moment was increased. On the final cycle, the l magnitude of +M2 2 was 3.8 x 106 in-lb (with load still increasing with increasing neflection) and the magnitude of -M was about 3.3 x 106 In-J b (with load not increasing wits increasing deflection, indicating that the maximum % load had been
,_ reached). The maximum moment was therefore (3.3/1.5) times M on the first half cycle. The maximum moment is indicated on l (d)
Fig. E5-1 by the leader to "Mmax"- !
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The other of the two points slightly (~14%) below the criteria, is the test identified as (13) 11 in Table 9 of NOREG/Ch-0261.
This test is anomalous in that pressure decreased rle limit moment, as compared to a presumably identical loading with P = 0; that being the test identified as (13)10. Figure ES-2 shows the complete load vs. deflection plot of these two tests.
Untortunately, the yield strengths of the materials in the elbows tested are not known; stainless steel elbows were assumed to have a yield strength of 30,000 psi. Accordingly, the seeming anomaly of pressure reducing M2, as shown on Fig. ES-2, may be due l entirely to a higher yield strength in the elbow tested with P = 0. In any case, the substantial stiffening ettect of internal pressure at high loads is apparent on Fig. ES-2 and the maximum applied moment is indicated on Fig. ES-1 by the leader to l "M mai -
The criteria for Class 1 with P = O, a = 900 are more restrictive than the Code C-limits for SA-312 TP304 at temperatures above 1000F.
The criteria for Class 2/3 with a = 900 are almost always significantly more restrictive than the Code C-limits. This is due to the fact that Be is larger than (0.751) ; i.e.,
B, =
1.3/h2/3 0.75i = 0.75 x 0.9/ha/3 = 0.675/h2/3 B2 /(0.751) = 1.3/0.675 = 1.93 This aspect can be seen on Fig. ES-1 by the line identified as l
" Code of limit Class 2/3 with(9)
Equation 0Kg o=f 1.SS " . OKg is the right-hand-side NC-3652.2 and, for SA312-TP304 at
~5000F, is equal to 1.5S7 This is one of the aspects where the Code rules are not supported on the basis of the conservative single-hinge, limit load concept used in NUREG/CR-0261. This is not to say that the Code rules are necessarily unconservative because a more sophisticated approach considering collapse mechanisms, strain hardening, dynamic effects, etc might show the Code rules to be appropriate.
E5.4 Branch Connections and Tees For Class 1 rules, the equation to be satisfied is:
Bt PDn +B6 Mh + Bar My 5 K (3) 2t 4b Zr Definitions of Bb, 2 Mb, Z , Bar, Mn and Zr are given in footnotes (5) , (7) , and (9) to Table NB-3682.2-1, and a Ka represents the right-hand-side limit of Equation (9) of NB-3652.
The equation to be satisfied for Class 2/3 piping is:
E-10 Revision 4 - February 1981
PDn + (0.75i) MJ50Kg (4) 4t Z
(.,)
V For Class 2/3 piping, each end of the branch connection or tee is checked separately in accordance with NC-3652.4 (c) or (d). For checking the run ends, Z is simply the section modulus of the run pipe, Z= (w/4) . For checking the branch end, however, Z 1s defined as (w/4) DaT[3, da where t 3 is the lesser of T ror (1) (t) .
E5.4.1 Branch Connections, Class 1 Piping TaDie 9 of NUREG/CR-0261 summarizes limit moment test data or branch connections. Table E5-1 herein is Table 9 of l NUREG/CF-0261 with one column added; the column headed M L/ Men, where M p is the moment permitted by the criteria given herein for Class 1 piping. The ratio ML /M p is the ratio of test limit moment (essentially Me) to the moment allowed by the criteria with the right-hand-side limit of Equation (9) of NB-3652 equal to S . The criteria set the right-hand-side limit as 2.0S .
Accordingly, to justify the criteria limit ot 2.0Sy, all values in the column of Table ES-1 headed Mart shou 1F be 2.0 or l higher. It can be seen in Table ES-1 that this is the case, except for the test identified as (34) 6, where M /ML rs = 1.95. In most tests, the criteria are very conservative (q,/M pa >> 2) Dut, as can be seen by comparing M t /Mf with ML /M p not as excessively conservative as the present Code with a right-hand-side limit of the same value.
O)
(, E5.4.2 Reinforced Piping and Unreinforced Fabricated Tees, Class 2/3 The column in Table ES-1 headed M L /Me , is the ratio of the test l limit moment (essentially M 2) to the moment permitted by the criteria for Class 2/3 piping with the right-hand-side limit of Equation (9) of NC-3652.2 equal to . The criteria set the right-hand-side limit as 1.54 Accordingly, to justify the criteria limit of 1.54, all valbes in the column of Table h5-1 headed ML/M p should be 1.5 or higher. It can be seen in Table ES-1 that this is the case, except for the test identified as l (35)2, M g , where M I/Mc2 = 1.4. In most tests, tne criteria are very conservative (ML /Me r >> 1. 5) , although not as conservative as the criteria for Class 1 piping.
E5.4.3 Butt-Welding Tees, Class 1 Piping Table E5-2 shows data from the insert table on page 29 of l NUREG/CR-0261. The column headed M L/M p is the ratio of the test limit moment to the moment permitted by the criteria for Class 1 piping with the right-hand-side limit of Equation (9) of NB-3652 equal to 4 The criteria set the right-hand-side limit as 2.0S AccoYdingly, to justify the criteria limit of 2.0SI, all valuds. of Table ES-2 in the colunn headed ML/Mgt should be 2.0 or
,_ higner. It can be seen in Table ES-2 that this is essentially the case.
J E-11 Revision 4 - February 1981
E5.4.4 Butt-Welding Tees, Class 2/3 Piping The column headed M /M t g2 in Table ES-2 is the ratio of the test l limit moment to the moment permitted by the criteria for Class 2/3 piping with the right-hand-side 1rmit of Equation (9) of NC-3652.2 equal to S . Tne criteria set the right-hand-side limit as 1.SS . Accdrdingly, to justif y the criteria limit of 1.5S , all vald's in Table ES-2 in the column headed g/M 2 shou 1d be 1.5 or higher. It can be seen in Table ES-2 that th[s is essentially the case.
E5.5 Otner Products / Joints The coverage of products / joints in the criteria is intended to be the same as the Code.
Class 1 Piping Footnote (a) to Table E4-1 uses the concept given in l NUREG/CR-0261, page 52, " Piping Products With B = 0.5, B, = 1.0". For reducers, the criteria uses the B and B, indices given in the Code. However, the 1.5S limit in tne criteria provides additional conservatism as c6mpared to Code rules where the limit exceeds 1.5S 7; e.g. for Level C-Limits , SA312 TP304 above 1000F.
For girth fillet welds, the Code gives B = 0.75, B, = 1.50. In piping, tillet welds are used to join the pipe to socket welding fittings, socket welding valves, slip-on flanges or socket welding tlanges. From a functional capability standpoint, the fitting, valve, or flange reinforces the pipe at tne weld so that the pipe cross section will not deform to the extent that straight pipe, remote f rom such reinforcing, would deform under the same loads. Accordingly, the criteria conservatively uses the same indices for fillet welds as for straight pipe; i.e.
B = 0.5, B 3 = 1.0. As discussed in Reference 4, pages 56-57, t the Code La and B, indices are higher than for straight pipe to l encompass the possibility (albeit remote) that a f ull fillet weld (a weld with legs not less than 1.4 trmes the pipe wall l thickness) may not be achieved with ANSI B16.11 socket welding 11ttings . However, this is an aspect related to pressure boundary integrity, not functional capability. It should be l recalled that the functional capability criteria given herein are l not intended to supersede any requirement of the Code.
Class 2/3 Piping Footnote (b) to Taule E4-2 uses the same concept as Footnote (a)
! to Table E4-1, although necessarily in dif f erent words.
Footnote (a) to Table E4-2, for sake of completeness, recognizes i that there are some products covered by Fig. NC-3673.2 (b) 1 which i are seldom used in Mark II piping systems. It also recognizes that, for " branch connections", the equation given in Fig. NC-3673.2 (b) 1 is not appropriate and points out that Class 1 criteria may be used.
l E-12 Revision 4 - February 1981 l
I l
l For concentric reducers, the criteria uses (0.751) from i Fig. NC-3673.2 (b) -1. However, the 1.5S limit in the criteria
( provides additional conservatism as codpared to Code rules where
\_s the limit exceeds 1.5S y; e.g., tor Level C-Limits, SA 312 TP304 above 5000F.
For girth tillet welds and brazed joints, the criteria applies the same value of (0.751) used for straight pipe; (0.751) =1.0.
The basis for this is the same as discussed under Class 1 piping; the presence of a fitting at a fillet welded or brazed joint icreases the functional capability load capacity as compared to straight pipe remote from reinforcement, hence use of (0.751) =1.0 is conservative for functional capability.
ES 6 Limits for D /t >50 NUREG/CR-0261 points out that for D/t >50, it is not prudent to assume that the limit load theory is an adequate assessment of the moment capacity ot straight pipe.
Table 2 of NUREG/CR-0261 includes considerable test data on the moment capacity of straight pipe with D/t >S0; these data are plotted on Fig. ES-3. An indication of relative moment capacity l of pipe as a function of D/t can also be obtained from Paragraph UG-23 (b) of the ASME Boiler and Pressure Vessel Code,Section VIII, Division 1(1). This paragraph provides rules for establishing maximum allowable longitudinal compressive stresses on cylindrical shells. This is applicable for a uniform (around O the circumference) longitudinal stress and hence is conservative for evaluating moment loadings, where the maximum compressive stress occurs at only one point around the circumference.
< Nevertheless, the relative values of allowable longitudinal stress as a function of D/t should be applicable to bending of straight pipe. The relative valuea, for cerbon steel with specified minimum yield strength or 30 to 38 kai at temperature up to 3000F, are shown on Fig. ES-3. Based on this data, a criteria factor for D /t >50 har been established as indicated on Fig. ES-3.
It- may be noted that a few test points lie beneath the criteria line. Such data, of course, is expected to erhiDit some scatter due to variations in test measurements, drmensions of test specimens and determination ot material yield strength. It is significant to note that S, as used in the criteria, is the minimum specified yield strengfh at 1000F and is the minimum expected yield strength at higher temperature. Most piping materials will have yield strength well above the specified or expected minimums, hence there is a statistical margin-of-saf ety in the criteria.
E5.6.1 Temperature Effect on Material Properties All of the data on Fig. ES-3 represent data from room temperature l
(}
\_/
tests. At elevated temperatures, the allowable moment decreases because Sy decreases. To assess whether this is adequate to E-13 Revision 4 - February 1981
compensate for buckling, the data given in Paragraph UG-23 (b) ot 9cterence 1 were used. For austenitic stainless steel, the decrease in allowable longitudinal stress with increasing temperature is essentially proportional to the decrease in yield strength with temperature. However, for territic steels like SA106 Grade B, the decrease in allowable longitudinal stress with increasing temperature from 100 to 7000F is about 25 percent more than the decrease in yield strength with increasing temperature from 100 to 7000F. Accordingly, a f actor of (1.033-0.00033T) :ss been used i n the criteria for ferritic steels to account for the temperature eff ect on material properties.
E5.6.2 Internal Pressure Ef fect All of the data on Fig. ES-3 represent tests with zero internal l pressure. As discussed in NUREG/CR-0261 (p. 18-19, on " Buckling With Internal Pressure") , some test data indicate that internal pressure will increase the moment capacity of pipe with D o/t >50.
However, these criteria take the conservative approach that internal pressure does not increase moment capacity: 1.e. the B Indices are not a f unction of D o/t.
E5.6.3 Products / Joints Other Than Straight Pipe Available test data are on straight pipe; it is deemed to also be directly applicable to girth-butt welds and transition joints.
For elbows with small h, the criteria are probably very conservative because the B, index for elbows reflects the tendency for the cross section to become out-of-round. However, in the absence of supporting data, these criteria apply th3 corrections for Do/t >50 to all products and joints.
E5.7 Dynamic Ettects In many nuclear power plant piping systems, the large and therefore design-controlling loads are caused by dynamic conditions such as steam hammer, reliet valve transients, postulated earthquakes or pipe breaks. These have the characteristic of rapidly oscillating loads with relatively snort total time duration. Some examples of the load characteristic are shown on Figs. ES-4 through 6 herein. l The test data cited in NUREG/CR-0261, with the exception of one set of data, represent tests in which loads were slowly applied.
The one exception is shown on Fig. ES-1; the five data points l encircled and identified as "Teidoguchi Dynamic Tests". These were tested under dynamic loadings (on a shake-table) at a treguency of about 3 Hertz. It can be seen on Fig. ES-1 that l these test data are about a f actor-of-two higher (stronger) than would be expected from the other static loading test data. The elbow did not loue functional capability during the tests, although fatigue failure was to be expected and did occur. This set of tests suggests tnat, under dynamic loading, the criteria given herein may be excessively conservative but the single set E-14 Revision 4 - February 1981
of data is not sufficient to provide a more realistic basis over the parameter ranges covered by the criteria.
(3
\~ I hater hammer occurs in industrial piping systems and can, for piping mad e of brittle material such as cast iron, cause failure of the piping. However, for piping made of ductile materiai, the ability of piping to absorb these dynamic loads without failure is impressive. Reference 5 gives a quantification or a service occurrence which is relevant to dynamic effects and is discussed in the following.
Due to malf unction of a regulation valve, pressure transients occurred in the piping system shown on Figs. ES-7 and 8. The location of the portion of piping shown on Fig. E5-8 with respect to the entire piping system can be seen by noting correspondence at Points 58 and 94 on the two figures. During a shutdown alter the incident, the piping system was examined for signs of damage to the piping; none were found. In addition, estimates were made of pipe movements as indicated by the physical evidence. These movements were then used as input data for static analysis of the piping system.
Several combinations of movements were evaluated; the combination of interest herein is identified in Ref erence 5 as Case I. A transient bending stress of 153,328 psi was calculated at point "X" as shown on Fig. ES-8. While most of the piping in the system was 24 inch, Schedule 120 (1.812 inch wall) , the U-shaped bypass shown on Fig. ES-8 was 6 inch, Schedule 760 (0.718 inch (j~3 s nominal wall).
To evaluate this incident in terms of the criteria, we note that (a) P = operating pressure = 1250 psi (b) T = operating temperature = 3400F (c) Material was SA-106 Grade B, Sy=30.5 kal (at T = 3400F)
(d) Point "X" on Fig. ES-8 is " straight pipe" (e) The bending stress due to weight at 'oint "X" was 651 psi Accordingly, the criteria gives:
B3 PD_ +B 2 M s 30,500 2t" Z l x 1250 x b.625 + 1.0 x (53,328 + 651) s 45,750 2 2 x .718 2883 1 53,979 5 45,750 I \ 56,862 5 45,750 O
E-15 Revision 4 - February 1981
The ratio of 56,8b2 to 45,750 is 1.24, hence the stress in this incident exceeded the criteria limit by 24 percent. However, there was no evidence of loss of functional capability; or indeed of any damage to the pipe.
This bit of quantitized service experience adds confidence to the adequacy of the criteria. Additional evidence of this type may be developed and should aid in removing excess conservatism in the criteria.
E5.8 Summary For 43 /t >50, the allovable moments are decrea.wi by increasing the aj) indices and equivalents of (0.75i) . This is based on tests data on straight pipe at room temperature with, for ferritic materials, a temperature factor based on ratios of allowable longitudinal compressive stresses from Reference 1.
Dynamic effects may make the criteria excessively conservative when used for conditions where the loadings are dynamic in nature.
The functional capability criteria Ior the essential piping in the Shoreham plant are in conformance with the GE topical report NEDO-21985(6) and the regulatory evaluations in Reference 7.
O l
l l
l l
l O
E-16 Revision 4 - February 1981
E
6.0 REFERENCES
- 1. ASME Boiler and Pressure Vessel Code,Section VIII, Division 1, " Pressure vessels", 1977 Edition with Addenda up to and including Susuner 1978.
- 2. Rodabaugh, E. F- and Moore, S. E., " Evaluation of the Plastic Characteristic- of Piping Products in Relation to ASME Code Criteria", NURSG/CR-0261, ORNL/Sub-2913/8, July 1978.
- 3. Rodabaugh, E. C., Iskander, S. K., and Moore, S. E., "End Effects on Elbows subjected to Moment Loadings",
, ORNL/Sub-2913/7, March 1978.
- 4. Rodabaugh, E. C. and Moore, S. E., " Stress Indices for Girth Welded Joints, Including Radial Neld Shrinkage, M; smatch, and Tapered-Wall Transitions", NUFF.G/CR-0371, ORNL/ Sub-2913/9, September 1978.
S. " Analysis of Reactor Feedwater (Piping System) Under Pressure Transients", Sarger* S Lundy Report, Dresden-3, Project No. 4989, October 19'r4.
- 6. General Electric Topical Report NEDO-21985 " Functional Capability Criteria for Essential Mark II Piping", September 1978.
- 7. Knight, J. P. (USNRC) memorandum to R. L. Tedesco (NRC)
"Evaulation of Topical Report - Piping Functional Capability Criteria", July 17, 1980.
O E-17 Revision 4 - February 1981
__ _. _.._ __ __ . ~ - . - . . _ _ _ _ _ _ , . ___ _ _ -. - - - -
O O
, TABLE El.-1 B-INDICES FOR FUNCTIONAL CAPABILITY EVALUATION, Dn/t 550 4 FOR USE IN EOUATION 19) OF NB-3652, CLASS 1 PIPING Piping Products or Joints B. B.
Straight Pipett) 0.5 1.0 Curved Pipe or (-0.1 + 0.4h), a o 2908, 1.3/h afs { but not buttwelding el- but not less ao= 45*, 1.17/h *.S* f less than bows per ANSI than 0.0 or ao= 0* ,_1.0- 1.0 B16.9, ANSI greater than i B16.28, MSS 0.5. For Interpolate linearily for SP-48 or MSS B, = 1.0, .other values of oo.
SP-87 B, = 0.5.
Branch Connections 0.5 except if Brb = 0.50 Cab , but not less
- per NB-3643 either Bab or ~ than 4/3 l- Bar is 4/3, B = 2/3 B,, = 0.75 C,, , but not less
., than 4/3 Butt-welding- 0.5 except Bab = 0.4 m m /Tr) */8, but not less tees per ANSI B16.9, if either than 4/3 Bab or Bar SP-48 or MSS is 4/3, B: = Bar = 0.50 (Rm/T,) afa, but not less SP-87 2/3 than 4/3 i
Butt-welding 1.0 1.0 g
reducers per ANSI B16.9, MSS
, SP-48 or MSS SP-87 Girth fillet weld 0.50 1.0 4
to socket weld fittings or valves slip-on flanges, or socket welding flanges t
4 d
(8) Applicable to all piping products / joints which have B. = 0.5, B, = 1.0 in Table NS-3682.2-1.
1 of 1 Revision 3 - November 1978
l TABLE E4-2 VALUES TO BE USED IN PLACE OF (0.75i) POR FUNCTIONAL CAPABILITY EVALUATION, Dn/T <50, FOR USE IN EQUA. TION (9) OF NC-3652.2, CLASS 2 OR 3 PIPING Description (*) (0.751)
Straight Pipe (2) 1.0 Welding elbow, pipe bend, ao2 908, 1.3/h afa inat not less or miter bend ao = 45*, 1.17/h *.56 than 1.0
= 08, 1.0 ao 1 Interpolate linearily for other values of ao.
Reinforced fabricated tee 0.751, but not less than 1.0. Use i as defined on Figure NC-3673.2 (b)-1 Unreinforced fabricated tee 0.751, but not less than 1.0. Use i as defined on Figure NC-3673.2 (b) -1 Welding tee per ANSI B16.9 0.901, but not less than 1.0. Use O' i as defined on Figure NC-3673.2 (b) -1 Fillet welded joint, socket 1.0 welded flange, or single welded slip on flange Brazed joint 1.0 Concentric reducer (ANSI B16.9 0.751, but not less than 1.0. Use i or MSS SP48) i as defined on Figure NC-3673.2 (b) -1 l I
l (13 Piping products included on Figure NC-3673.2 (b)-1 but not l covered by this criteria are: " Closely spaced miter benda; "Widely spaced miter bend"; " branch conneetion", " threaded pipe joint or threaded flange" and " corrugated straight pipe or corrugated or creased bend". Functional capability of these piping products shall be demonstrated by appropriate methods. For " branch connection", the Class 1 criteria are applicable.
(a) Also applicable to " butt weld, tn >3/16 and 6/ g 50.1 in. ;
" butt weld, t n 53/16 or s/tn >0.1 in. ; " full fillet weld"
- O1 and "30 degree tapered transition (ANSI B16.25) ".
1 of 1 Revision 4 - February 1981
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as
- M. =
"V"C* , where Me . la calculated by Equation (4) of NuxtG-0461.
Mg is trae experimental lamat moment.
V.aues tollowed by "um ma e deemeus unstable by authoza or cited reterer.ces.
saa 8, a stress inadex calculated by Cooe rules.
(** M,3 = allowaole un-nt by Ct><te Equation (9) with a a a = 1.0, Class 1.
Values tollcwed by
- are, tests with a/u larger than covert.d b) be .
468 n,e = allowar,1t mosnent by tooe Equatros. (9) witt.Br.g = 1.0, ca.as 4.
(*8 ng, = allowable naoment by zur.ctional capablaity criterra with sight-bano-sace lis.it or b 4
1 2 ot 2 hevsslun * - Fet.ruly 19e1
TABLE ES-2 l O- TEST DATA ON BUTT-WELDING TEESCa), FROM NUREG/CA-0261 M Lta) s7 (a) ML Mt., ML ML Spec.
No. Ms(*)
e Me r(*) Mf (*) M f ,(*)
in-lb psi (c)
PT-1A 421,200 44,600 2.70 1.35 2.15 1.62 PT-1B 442,800 44,600 2.84 1.43 2.27 1.71 PT-1E 422,100 44,600 2.71 1.36 2.16 1.63 PT-2A 396,900 42,250 2.69 1.35 2.14 1.60 PT-2B 411,300 42,250 2.79 1.40 2.22 1.67 PT-26 362,000 42,250 2.46 1.23 1.95 1.47 (1) 6 x 6, Sch. 40 (.280 nom. wall) , D/Tp = 22.7, d/D = 1.0; t/Tr = 1.0.
(2)
ML = limit moment. In first three tests, the branch pipe yielded without any visible plastic deformation of the tee. In the second three tests, the tees deformed appreciably while the branch pipe remained straight.
(3) Sy = Yield strength of tee material
(*) Me n = allowable moment by Code Class 1 with 1.0Sy limit M e, = allowable moment by Code Class 2 with 1.0Sy limit Mf = allowable moment by Functional Capability Criteria Class 1 with 1.0S, limit M f, = allowable moment by Functional Capability Criteria Class 2/3 with 1.0S y limit l
O 1 of 1 , Revision 4 - February 1981
2.0 , , , , , , , , ,
'a a
TEIDOGUCHI -CODE CLASS 1.0 -
D}yjpyc' 2/3 WITH
- 0. 9 -
Mmax SKS =l.5 Sy A
O.8 - ' -
O a O.7 -
o , A -
O.6 -
8 -
0.5 -
Mmn sa -
e 33 FUNCTIONAL O. 4 - CAPABILITY _
M e CRITERIA coL> 90' D 2 tS y e 0,3 - & -
(23),- M z 0.1 ' ' ' ' ' ' ' ' '
O.I O.2 0.3 0.4 C5 0.6 0.70.80.91.0 >
h = 4t R/ D 2 i NO WITH l PRESSURE PRESSURE REF (22) OR (23) e o REF. (13) OR (24) A a FIG. E 5-I TEST DATA O'l BUTT WELDING ELBOWS WITH ARC ANGLE, aO,OF 90* OR 180*,
FROM TABLE 4 OF NUREG/CR-0261 O SHOREHAM NUCLEAR POWER STATION -UNIT I PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS l
l l REVISION 4-FEBRUARY 1981
O 6 , i , i <
l 8e _
~
8e , M max > 338,000 in -lb -
PD/2t =l8.6 ksi 5 - -
(13) ll 4 - -
~
a f ,,M m ax E l99.OOO in -lb
$3 -
3)10 0
3 M2 = 160,000 in -lb FOR P= 0, (13) 10 2 - -
M2 = 123,000 in-Ib FOR PD/2t =18.6 ksi,(13) 11 I -
O' ' ' ' ' '
O 10 0 200 300 400 500 600 END DISPLACEMENT (mm)
MOMENT (kg-mm) IS 707 X LOAD FIG, E5-2 LOAD-DISPLACEMENT PLOTS, 4
ELBOWS IDENTIFIED AS (13) 10 AND (13) ll lN NUREG/CR-0261 SHOREHAM NUCLEAR POWER STATION-UNIT 1 PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 4-FEBRUt.RY 1981
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FIG. E 5 -6 T TYPICAL PIPING ELASTIC RESPONSE TO OBE s SHOREH AM NUCLEAR POWER STATION-UNIT 1 PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 4-FEBRUARY 1981
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V SHOREHAM NUCLEAR POWER STATION - UNIT I PLANT DESIGN ASSESSMENTFOR SRV AND LOCA LOADS REVISION 4-FEBRUARY 1981
1 I
i APPENDIX F IS INTENTIONALLY LEFT BLANK l
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6 Revision 4 - February 1981
. . - - _ - . _ . . - . - . . . _ . . . _ . , . _ . _ _ _ _ . _ . _ _ _ _ . - - . - . _ - . _ . _ _ _ . . _ _ _ . . . _ _ . _ . - _ . _ - . . - . - _ ~
APPENDIX G JUSTIFICATION OF MARK-II LEAD PLANT SRV DE, SIGN BASIS W D DEFINITION i
a l
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i Revision 4 - February 1981
- - ... ...- - .-. - - .. -,. - -.-_ - - - ..... -. - - - .. - ..-~ - - - -. - . - , - - - .- . - --..- -. -..
~ - __
APPENDIX G TABLE OF CONTENTS Section Title Page
SUMMARY
G-iv G
1.0 INTRODUCTION
G-1 G2.0 CONSIDERATION OF THE POOL BOUNDARY SRV LOADS G-1 G2.1 Major Elements of a Pool Boundary Pressure Specification G-2 G2.2 Assessment of the Ramshead Load Cases G-3 G2.3 Shoreham Interim T-quencher Load Specification G-b G3.0 i COMPARISON OF CONTAINMENT STRUCTURE RESPONSES G-8 G3.1 Introduction G-8 G3.2 Method of Analysis G-9 G3.3 Results G-10 G3.4 Conclusions G-13 I
G4.0 COMPARISON OF PLANT COMPONENT RESPONSES G-13 G4.1 Piping Subsystems G-13 G4.2 Equipment G-17 s G5.0
SUMMARY
AND CONCLUSIONS G-20 REFERENCES G-22
- O G-i Revision 4 - February 1981
LIST OF TABLES Table Title G-1 Piping Characteristics G-2 Typical Results at Selected Locations - Part 1 Support Load Comparison - Part 2 Primary Stress Intensity Comparison G-3 Typical Results at Selected Locations G-4 FW 301 Mode-By-Mode Contribution Resultant Moments G-5 CRD 1265 Mode-By-Mode Contribution Resultant Moments G-6 CLCW 031 Mode-By-Mode Contribution Resultant Moments G-7 RWCU 013 Mode-By-Mode Contribution Resultant Moments G-8 MS 2500 Mode-By-Mode Contribution Resultant Moments G-9 Overall Average Ratios of (SRVg) (SRVTQ I O
O G-li Revision 4 - February 1981
1 l
LIST OF FIGURES l N 2 Fiqure Title G-1 General Arrangement of Reactor Building G-2 Structural Model Nodal Point Numbering Scheme G-3 Loads Applied to Model G-4 Idealized ARS Comparison: Ramshead All SRV Discharge G-5 Pedestal ARS Comparison: Ramshead All SRV Discharge (Cases 1, 2, S 4)
G-6 Primary Containment ARS Comparison All SRV Discharge (Cases 1, 2', S 4)
G-7 Secondary Containment ARS Comparisons: Ramshead All SRV Discharge (Cases 1, 2, S 4)
G-8 Pedestal ARS Comparison: Ramshead All SRV Discharge (Cases 1, 2, S 4)
G-9 Pedestal ARS Comparison: Ramshead All SRV Discharge (Cases 1, 3, S 5)
G-10 Secondary Containment ARS Comparison: Ramshead All SRV Discharge (Cases 1, 3, S 5) l G-11 Idealized ARS Comparison: Ramshead and T-Quencher All SRV Discharge G-12 Amplified Response Spectra of Vertical Acceleration Top of Reactor Support Pedestal T-Quencher All SRV Discharge - Pressure Trace No. 1 G-13 Amplified Response Spectra of Vertical Acceleration Top
(
of Reactor Support Pedestal T-Quencher All SRV Discharge -
Pressure Trace No. 2 G-14 Amplified Response Spectra of Vertical Acceleration i Top of Reactor Support Pedestal T-Quencher All SRV Discharge - Pressure Trace No. 3 G-15 Pedestal ARS Comparison: Ramshead and Interim T-Quencher All SRV Discharge G-16 Primary Containment ARS Comparison: Ramshead and Interim T-Quencher All SRV Discharge
, G-17 Secondary Containment ARS Comparison: Ramshead and l Interim T-Quencher All SRV Discharge l
G-18 Static Coef ficients - Primary Containment Upset Condition l N+OBE+SRV
! G-19 Static Coefficients - Primary Containment Faulted Condition N+SSE+SRV+LOCA G-20 Static Coefficients-Secondary Containment Ppset Condition N+OBE+SRV G-21 Static Coefficients-Secondary Containment Faulted Condition N+SSE+SRV+LOCA G-22 Static Coefficients-Primary Containment Upset Condition Design Curves G-23 Static Coefficients-Primary _ Containment Faulted Condition Design Curves G-24 Multif requency Qualification Testing-Horizontal G-25 Multifrequency Qualification Testing-Vertical O
G-iil Rev2sion 4 - February 1981 l
l
SUMMARY
This appendix demonstrates that the safety / relief valve (SRV) air clearing loads presented in the Design Assesment Report (D AR) ( 1 )
Section 3.2.1.1 and based on the DFFR(2) ramshead (IW) methodology result in conservative pool boundary loads. Tne concerns expressed by the NRC regarding SRV bubble phasing and trequency characteristics have been addressed by consideration of the maximum credible range ot input conditions with respect to resulting pool boundary loads. Furthermore, comparisons of the effects on the plant are made with those resulting from a conservatively constructed load definition based on the KWU T-quencher. This quencher is the actual discharge device installed in Shoreham and other lead plants. Quantitative results in terms of building response, pipe stress, pipe support reactions, and equipment loads are presented to provide a basic means of demonstrating the conservatism cf the Mark II lead plant load definition (3) based on the RH device. It is, therefore, concluded that the load definition methodology utilized provides an acceptable and conservative basis for the design of the plant.
O i
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l l
O l
G-iv Revision 4 - February 1981 l
G
1.0 INTRODUCTION
O)
(d The purpose or this Appendix is to demonstrate that the design basis load definition described in DAR(1) Section . 3.2.1.1 for the safety / relief valve (SRV) air clearing pool boundary loads is conservative.
The load tinition described in DAR Section 3.2.1.1 is for a RH discharge s vice. The actual discharge device installed in Shoreham is the KWU T-Quencher (1N) , however, af ter reviewing the Marx II lead plant (e.g. Shoreham) design basis load definitions, the NRC concluded that the RH ramshead load predicting methodologyta) is generally conservative for lead plants employing the KWU TQ device (3).
Due to the lack of a complete quencher data base and to introduce additional conservatism, the NRC defined a simultaneous entry, in phare oscillation load case, which was referred to as Load Case 5 in Reference 3. This was to be considered together with a wide frequency range and absolute sum combination of the etrects of each SRV bubble, in the plant assessment, In response to NRC concerns and to demonstrate that the excessive conservatism of Load Case 5 is not required, several postulated load cases with various time phasing and frequency ranges are considered in this Appendix with the use of RH methodology.
s To further justify exempting the RH load definition trom Load l (( ) Case 5, Shoreham has established an interim TQ load definition based on the in plant test results from KWU BrunsDuttel plant.
It is shown that the plant design assessment basis (RH without simultaneous, in phase oscillation) bounds the TQ load definition even with application of Load Case 5 to the TQ load. The interim TQ load is subject to confirmation by the Lead Plant Generic Load Definition currently under NRC review (*,5).
Section G2.0 provides a brief discussion on the consideration or various pool boundary load detinitions. Section G3.0 discusses the containment structure responses due to the various pool boundary loads. Section G4.0 provides a description of typical piping subsystem and equipment responses. Section G5.0 provides the concluding remarks on the analysis presented in this report.
G2.0 CO!1 SIDERATION CF THE POOL BOUNDARY SRV LOADS In accordance with FFR Rev. 2(a), four pool boundary load cases should te conside. for the application of RH load specification. Th four cases are: All SRV sequential actuation, ADS SRV discharge, asymmetric SRV di scharge , and l single SRV discharge. The extremely conservative all valve
- simultaneous entry and in phase oscillation " Load Case 5" is not l creditable for Mark II plants. The assumption that the buobles g from all SRV discharge lines enter the pool simultaneously and
(~s G-1 Revision 4 - February 1981
oscillate in phase with the same frequency is unrealistic.
Consideration of the actuation mechanism, line length, and friction factors would lead to the conclusion that such a situation would not occur.
To address the adequacy of the Shoreham design assessment basis load detinition, four hypothetical all valve RH loaa cases are established. These hypothetical all valve load cases include various degrees of conservatism on phasing and frequency considerations. The intent is to demonstrate that the sequential all valve load case bounds both the maximum credible hypothetical RH and TQ loads and hence can be used as the design basis load definition.
G2.1 Maior Elements of a Pool Boundary Pressure Specification The actuation of an SRV will produce a transient pressure time history on the pool boun dary . The characteristics ot the pressure time history include:
- 1. peak pressure magnitude,
- 2. dominant oscillation frequency, and
- 3. pressure attenuation rate.
- u. pressure oscillation duration.
The above characteristics, together with the pool boundary pressure disrlbution, constitute the major considerations of a pressure time history specification.
Futnermore, if there is more than one SRV actuated, the time phasing between the pressure time histories will also have a major effect on plant response.
The DFFR Rev. 2(2) RH methodology is based on an analytical model which predicts the pressure magnitude, frequency, attenuation rate, and the duration of an expected RH air clearing pressure time history to be used with the specified load cases.
It is well known that a RH device produces a higher pool boundary pressure than a TQ device under otherwise identical plant conditions. In recognizing the conservative nature of the SRV load definition by the RH methodology, Mark II lead plants conclude that the load cases specified in DFFR-2(2), including tne use of the most probable bubble frequency and the sequential actuation of all valves provides a conservative SRV load de21nition f or plant assessment.
G-2 Revision 4 - February 1981 O
G2.2 Assessment of the Ramshead Load Cases
[}
N/
Recognizing that the time phasing and frequency content of the SRV air bubble oscillation are important parameters in the load specification, the Mark II lead plants have postulated four hypothetical load cases for investigation. Since the sequential actuation case defined in DFFR Rev. 2(2) is generally the most severe load case of the design basis load, it is used as the basis tor comparison and is included for discussion below. These five all-valve cases (base case and hypothetical cases) are:
- 1. Sequential Actuation Case (base case) ,
- 2. Simultaneous-Discharge In-Phase-Oscillation Case,
- 3. Resonant Sequential Actuation Case,
- 4. Simultaneous Actuation Case, and
- 5. Resonant-Actuation Shifted-Frequency Case.
The individual air bubble characteristics, e.g., bubble frequency and pressure amplitude, are calculated according to the RH methodology with consideration of individual SRV line length, air volume, and valve set pressure. The pool boundary loads are then calculated by the method of images as described in DFFR Rev. 2(2).
("h
(') Detailed descriptions of these five all-valve cases are provided in the following subsections.
G2.2.1 Sequential Actuation case (base case)
The SRV's are actuated according to their set pressures. The valves with the lowest set pressure are actuated 11rst. The valves with higher set pressues are actuated sequentially according to the selected reactor pressure rise rate. In this case, the highest possible reactor pressure rise rate is selected to define the time intervals between the actuations of the SRV's in different SRV set pressure groups. The intent of this time phase case is to have all the bubbles entering the pool in the shortest time duration within a realistic range of plant transient conditions. The highest 'ossible pressure rise rate, based on plant transient analyses, for a Mark Il plant is about 130 ps1/sec. The air bubbles from SRV lines of the smne set pressure group will not enter the pool simultaneously due to differences in line lengths and the resultant difterences in vent clearing times. The vent clearing time for each SRV line is predicted by the vent clearing model of the DFFR RH methodology.
The pressure loads on the submerged boundaries are calculated by the method or images and based on algebraic sum.
O G-3 Revision 4 - February 1981
G2.2.2 Simultaneous-Discharge In-Phase-Oscillation Case For Case 1 and Cases 3 through 5 the air bubbles from different SRV lines do not enter the pool simultaneously. This results trom differences in vent clearing times due to differences in SRV line lengths and the differences in SRV set pressures. The bubbles from different SRV lines will also have dif f erent predominant oscillating frequencies due to dirferences in SRV line characteristics, e.g., air volume. The SRV line vent clearing time and the bubble frequency are predicted by DFFR RH me thodology.
In this case, it is assumed that the bubbles from all SRV lines enter the pool simultaneously and oscillate with the same treguercy. In the present analysis, the highest predicted frequency of all bubbles is used because previous analyses showed that this represents a more severe case than that using the lowest predicted frequency. The submerged boundary loads from different bubbles are added by the square root of the sum of the squares (SRSS) method. For all the sequential cases, the algebraic sum method is used to calculate the submerged boundary loads from contributions of the bubbles trom dif f erent SRV lines .
The SRSS method of adding the loading components for this case provides a more reasonable consideration of the randomness in SRV bubole oscillations as discussed in Reference 2.
G2.2.3 Resonant Sequential Actuation Case The purpose of this case is to provide a maximized in phase bubble oscillation case with realistic assumptions. The time intervals between the actuations of SRV's of different set pressures are chosen such that the bubbles rrom ShV lines of ditferent sqt pressures will tend to oscillate in phase. This is achieved by selecting a pressure rise rate, within the realistic limits (for Mark II plants, it is approximately from 40 to 130 psi /sec), according to the predicted bubble oscillation period.
In general, the first pressure peak of the bubbles from the second set pressure group valves tends to coincide with the second pressure peak ot the f irst (lowest) set pressure group valves. Once the reactor pressure rise rate is determined to satisfy the anove resonant condition, the DFFR RH methodology is employed to calculate the loads on submerged boundaries.
G2.2.4 Simultaneous Actuation Case For this load case, it is assumed that all SRV's are actuated simultaneously regardless of their set pressures. This is not a credible load case but provides the shortest time duration for all the bubbles entering the pool. However, the bubbles from ditierent SRV lines will not enter the pool at the same time due G-4 Revision 4 - February 1981 O
! 'to differences in line lengths and thus the vent clearing times 1
are different.
O
\w / G2.2.5 Resonant-Actuation Shifted-Frequency Case By the use of the DFFR RH methodology, predicted treguencies for bubbles from different SRV lines are not the same. This. is due to the differences in SRV line characteristics. The bubble frequency also changes from cycle to cycle as the bubble rises to the pool surface. These considerations predict a range of frequency for the SRV load cases for the lead plants. However, the frequency range provided by the DFFR methodology (for lead plants, it is about 7-10 Hz) is less than that required by the NRC specification. In order to provide some information about how a wider frequency range would affect the plant design assessment, the_ resonant-actuation shifted-frequency case is constructed and evaluated.. Based on NRC specification, it was stipulated that the frequency- range be increased to include the frequency predicted by the RH methodology together with a l 150 percent margin. To obtain a wider frequency range, a frequency multiplier ranging from 0.5 to 1.5 is applied to the predicted frequency. The bubble with the highest predicted frequency is increased by a multiplier of 1.5 and the bubble with the lowest predicted frequency is decreased by a multiplier nt 0.5 The frequencies of the rest of the bubbles are adjusted by multipliers between 0.5 and 1.5 at equal intervals. In the wall.
load calculation, the resonant sequential time phase method as described in Section G2.2.3 is used. It is noted that y_ frequencies determined by this method are not a credible case for i the TQ plant. From the Karlstein test results, the bubble
, frequency for a Mark II TQ plant would be in tne range of
! approximately 3 to 9 Hz. The frequency range' covered by this j method is about 3 to 13 Hz. The purpose of this SRV case is, therefore, to assess the effect of a wider frequency range on the containment structural responses. The DFFR RH methodology usually predicts a somewhat narrower frequency range for the all-valve case. Through this SRV case, it is intended to demonstrate that a wider frequency range will not increase the structural responses and, therefore, the effects on plant components and supports. It is also noted that the treguency range provided by the ramshead methodology adequately covers the higher treguency range of the test results.
G2.2.6 Conclusions The broad-based investigation of all the SRV load cases described in this section was performed to meet the intent of the NRC specification. The Mark II lead plants have recognized and addressed the two major NRC concerns, i.e., time phasing and frequency range, by evaluating several time phasing and frequency cases to define a realistic but bounding load case for plant design assessment. The entire phenomena are conservatively modeled using RH load definition methodology even though much
,V G-5 Revision 4 - February 1981
lower loads are anticipated from the TQ device installed in the plant. Theretore, the cases investigated will bound the actual discharge cases expected during plant operation with substantial margin.
G2.3 Shoreham Interim T-Ouencher Load Specification The current Mark II lead plant SRV load definition is based on the DFFR RH methodology. In order to make a preliminary evaluation about the expected plant structural responses resulting trom the SRV actuation with the actual quencher device, a TQ interim load definition is presented in this section. This load detinition was previously submitted to the NRC in the Shoreham DAR, Rev. 3.
The quencher device installec in the lead plants is the TQ designed by KWU for Mark II plants. Based on the knowledge gained from the quencher development program ano a review of results from about 200 in plant tests, KWU has defined a conservative load definition for the Mark II TQ plants.
The quencher wall load definition consists of three pressure traces from the KWU Brunsbuttel in plant tests. Tnis selection is the result of a close examination of all KWU in plant test results at different KWU power plants. These pressure traces were selected not only for their f requency variations but also for their relatively large pressure amplitudes of 0.5 to 0.8 bar (7.3 to 11.6 psid) . The 0.8 bar pressure trace contains the highest pressure amplitude ever measured for in plant tests of BWR plants equipped with KWU quenchers. In order to provide a conservative load definition, a pressure amplitude multiplier (larger than 1) is applied to the pressure traces and a frequency range is obtained by modifying the time scale of the pressure traces. It is also noted that the selected pressure traces are from subsequent SRV actuation (multiple pop of the same line) cases for additional conservatism.
For plant structural analysis, the pressure distribution on the submerged boundaries is required. This pressure distribution is conservatively defined by assuming that air bubbles f rom all SRV lines oscillate in phase with the same intensity (highest possicle). This results in a f ull bubble pressure (pressure trace with the amplitude multiplier) applied unirormly over tne basemat and 7.25 feet above the basemat on containment and pedestal walls (the quencher is located at about 4.75 teet above the basemat for Shoreham plant) . In the region above 7.25 feet, the pressure amplitude linearly decreases to zero at the water surface. Figures showing the three pressure traces and tne normalized pressure distribution on the submerged boundaries are illustrated in the proprietary supplement to Section 3 of the Shoreham DAR and will not be presented here.
G-6 Revision 4 - February 1981 O
To assure a conservative load definition, a pressure amplitude
,_ multiplier of 1.1 and a frequency range are used with the three t pressure traces. The frequency range is specified for the three C) pressure traces by expandirg the time scale of the pressure-traces into a longer duration by a factor of 1.8 and into a smaller time duration (contraction) by a factor of 0.9. The three pressure traces have measured pressure amplitudes from 7.3 to 11.6 psi and predominant frequencies from 6 to 8 Hz. By the application of the amplitude and treguency multipliers, the interim T quencher load definition develops a maximum pressure of 12.8 psid and a frequency range of 3 to 9 Hz. In the structural analysis, this load definition consists of a number of subcases along the following lines. The maximum pressure (with multiplier) is applied to the pressure trace witn a selected frequency and a structural analysis is performed. For each pressure trace, a number of frequencies are selected to sweep the frequency range. This procedure is done for all three pressure traces. For the present analysis, five frequencies are selected for each pressure trace. The load definition, therefore, consists of 15 simultaneous-discharge in phase cases. This is equivalent to assuming that all the bubbles enter the pool simultaneously and oscillate with the same frequency, say 3 Hz, for one case, and oscillate with another frequency, say 9 Hz, for another case. From a phenomenological point of view, this simply cannot happen. However, it does provide a very conservative load cefinition.
rx It should be noted that the amplitude multiplier of 1.5 was recommended by KWU as a bounding multiplier for the design ot the (V) TQ prior to full scale confirmation test (Karlstein) for Marx 11 plants. However, for the interim TQ load detinition, an amplitude multiplier of 1.1 is used for the followng reasons:
- 1. The highest pressure amplitude pressure traces ever measured in in plant tests ar e used for the load definition. They are subsequent SRV actuation data and are generally substantially higher than pressures measured for first actuation.
- 2. It is considered very conservative to assume that bubbles from all SRV lines have this un.i.que high pressure amplitude regardless of the differeness in SRV line cnaracteristics. It is also noted tnot test data exhibit some statistical deviations with _egard to pressure amplitude and Irequency. Measured pressure amplitude and frequency are different for different SRV actuation tests under similar conditions. This is particularly true for subsequent SRV actuations.
- 3. The assumption is made that all the bubbles oscillate in pnas e. Because of the differences in SRV set pressure and line characteris tics, this simultaneous-discharge in phase-oscillation condition will not occur for any t
V G-7 Revision 4 - February 1981
real plant condition and is, therefcie, an unrealistically conservative assumption. Unoer real plant operating conditions, bubbles from all SRV lines will enter the pool at different times and oscillate with different treguencies. A load def3nition based on time phasing methodology is most appropriate for the all SRV actuation case.
- 4. The tnree pressure traces used in the load definition are from in plant test results of KWU plants. The TQ was designed based on considerations of differences between non Mark II (KWU) plants and Mark II plants such that a direct load transposition is possiole. The 1.1 multiplier is used to add further conservatism by introducing an additional 10 percent design margin on the measured pressure amplitudes.
The conservatism of the interim quencher load definition will be demonstrated when the Shoreham plant specific 'IQ load detinition is available. It taxes into consideration the Shoreham plant parameters and a statistical evaluation of measured pressure data which include the Karlstein test results. The Shoreham plant specific TQ load definition is based on the Lead Plant generic TV methodology and is described in Reference 4. The Lead Plant generic TQ methodology is currently under NRC staff review.
G3.0 COMPARISUN OF CONTAINMENT STRUCTURE RESPONSES G3.1 Introduction l The objective of this section is to provide an extensive l description of the dynamic response of the containment structures l to each of the safety / relief valve (SRV) discharge loads described in Section G2.0. These include 5 cases baed on
! ramsnead methodology encompassing a wide range of variation in l both bubble f requency and phasing, as well as conservatively l applied loads based on TQ discharge device test results from both the Brunsbuttel and Karlstein facilities.
The presentation of reactor building response to each of these loadings provides a substantial basis for addressing two l tundamental issues. The first is the NRC's concern over SRV l nubble f requency and phasing characteristics. The second, and
! ultimately more important issue, is the verification that loads based on the ramshead methodology provide a conservative design basis relative to the TQ device which are being intalled in the lead plants.
I The nature of SRV loads is such that they result in small i building displacements (and, therefore, small structural internal loads) but relatively high accelerations. The magnitudes or these accelerations are generally comparable to those associated with a seismic event but occur at somewhat higher frequencies.
G-8 O
Revision 4 - February 1981
Lead plants have snown in their Design Assessment Reports that n their containment structures have more than sufficient margins of safety to sustain the ef fects of SRV loads. The area of concern
(\ -) is in regard to the eff ects of the resultant building vibraticns (acceleratione; on piping components and mechanical systems. The traditional means of representing building accelerations is by use of the amplified response spectra (ARS). An ARS is a plot of the maximum response of a single degree of freedom system-to an acceleration time history as a function of the system natural frequency and damping ratio. It is a direct representation of the acceleration ' load' on a piping or mechanical system and will be used througnout this section to quantify the effects of the various SRV loadings considered herein.
All _results presented in this section have been generated specifically for the Shoreham plant, but are believed to be representative of results for the lead plants. The general arrangement' of reactor building structures is shown on Fig. G-1.
G3.2 Method of Analysis The method of analysis used to determine the dynamic response of the containment structures, when subjected to SRV discharge loads, has been described in detail in the DARC1). A su:mnary of the basic features is presented here.
The three-dimensional axisymmetric continuum is represented
~T either as an ax1 symmetric thin ahell, as a solid of revolution, g j or as a combination of both. The axisymmetri shell is discretized as a series of frustums of cones and the foundation is represented as a solid of revolution by triangular or quadrillateral torold elements connected at their nodal point circles. A finite element formulation is utilized and the equations of motion are solved numerically by direct integration.
ARS are developed from the resultant structural acceleration time-histories. The ARS's are generated using the exact analytical solution to the governing difterential equations of motion for single degree of freedom systems under successive linear segments of excitation.
Figure G-2 depicts the structural model used to lepresent the complete reactor building and supporting soil. As indicated there, solid axisymmetric elements are used to represent the soil to a radius and depth of approximately 1.5 mat diameters with ax1 symmetric thin shell elements representing the structures.
The external dimensions of the soil are selected to preserve the free-field motions. An examintion of Fig. G-2 indicates a closer spacing of elements in the area of the suppression pool with increasing element size in areas sufficiently rar from the pool.
This spacing has been used because the SRV discharge loads arc applied to the pool boundaries within this area and a precise
-s definition of internal loads is required.
v G-9 Revision 4 - February 1981
In general, the pressure loads on the suppression pool boundaries due to an SRV discharge event vary both circumf erentially and meridionally with time. At any point in time this pressure tield can be represented meriodionally by a descretization into zones and circumferentially by a Fourier series at each of the meridional zones. Therefore, the spatial and time-wise variation of pressure can be represented by pressure time histories at each zone for each Fourier series term.
Figure G-3 provides a graphical representation of the manner in wnlen the SRV pressure profiles are represented and applied to the structural model in terms of equivalent line forces and moments at the respective nodal circles.
G3.3 Results G3.3.1 Ramshead Load Cases Five SRV actuation cases based on RH methodology and encompassing a wide range of bubble frequency and phasing characteristics have been described in detail in Section G2.2. The following is a summary of the primary features of each of these load cases.
- 1. Sequential Actuation (base case) - Highest possible pressure rise rate (136 psi /sec) , with consideration of set pressures and vent clearing times.
- 2. Simultaneous Discharge - Bubbles enter the pool simultaneously and in phase. Bubble loads combined by SRSS since this case is most unrealistic.
- 3. Resonant Sequential Actuation - Bubbles are grouped according to set pressures and groups enter the pool sequentially with the basic feature that the first peaks of the second and third groups correspond to the second peaks of the first and second groups, respectively.
- 4 Simultaneous Actuation - Bubbles enter the pool out of phase due to different line lengths.
- 5. Resonant Actuation, Shifted Frequencies - BubDle with highest calculated frequency has frequency increased by l 50 percent. Lowest calculated frequency is decreased by
! 50 percent. Other bubbles have frequencies shifted at j equal intervals between these extremes. This results in I characteristic bubble frequencies between 3 and 13 Hz.
Resonant sequential time phasing is used. *This is even l beyond the range specified in Section 3.2.2.1 (heterence
- 3) resulting in an even wider range ot frequencies.
A complete dynamic analysis of the reactor building complex has been performed ror each of these loaa cases. As described in Section G3.1, ARS will be used as the neans of CGmparing the G-10 Revision 4 - February 1981 O
resulting building responses. A number of general trends occur throughout the results and are characterized in an idealized ARS as shown on Fig. G-4 and discussed below.
Response 'can be categorized into the three following frequency ranges:
- 1. Low frequency - Peaks in 5 to 10 Hz range (0.2 to 0.1 sec. period)
- 2. High frequency - Peaks in 15 to 20 Hz range (0.07 to 0.05 sec. period)
- 3. Very high frequency - Various peaks above 25 Hz (below 0.04 sec. period)
There is a consistent inverse relation between magnitudes of response in the high and low frequency ranges, i.e.. it response is high in one range it is low in the other. There is no clear evidence of a relation between bubble frequency and phasing characteristics and relative magnitude in the high and low frequency ranges. It is noted, however, that the Sequential Actuation case which has the least degree of bubble phasing consistently results in the largest amplitude in the high frequency range and the lowest amplitude in the low frequency range. The effect of this feature of the response on typical plant components will be discussed in detail in Section G4. No consistent pattern of relative magnitudes. is observed in the very high f requency range.
() Representative ARS actually computed for the five load cases considered here are presented on Figs. G-5 through 10.
Figures G-5 through 7 present ARS at the top of the pedestal, the ,
primary containment at the elevation of the stabilizer truss, and a secondary containment floor slab for the load cases of Sequential Actuation, Simultaneous Actuation, and Simultaneous Discharge. Figures G-8 through 10 present ARS at the same locations for the cases of Sequential Actuation (repeated as a casis of comparison) , Resonant Sequential Actuation, and Resonant-Shifted Frequencies.- Inspection of these actual ARS will demonstrate all of the generalizations discussed above.
G3.3.2 Interim T-Ouencher Load Specification The dvnamic response of containment structures has been determined f or the interim TQ load definition described in Section G2.3. The structural response was determined for the three TQ pressure traces with five time scale factor applied to each trace to encompass a broad range of predominant frequencies.
The time scale factors utilized were 0.8 (providing the highest f requency) , 1.0, 1.2, 1.4, and 1.8 (providing the lowest Irequency) . The individual ARS resulting from the three pressure curves each with five time scale factors are all enveloped and increased by the specified multiplier. ot 1.1 to produce G-11 Revision 4 - February 1981 4
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- - - . - - - - - , - . . , - - - - - - - - , , , , .- - - ~ - - - - - - - - - - - - - - - - - - - - - - - - - - - - -
conservative ARS for the assessment of piping and mechanical systems.
Figure G-11 depicts the general relation of ARS resulting from the interim TQ load definition to the ARS resulting from the lead plant load definition based on RH methodology. The TQ results generally follow the inverse relation between magnitudes in the high and low treguency ranges as observed in the RH cases. T: e amplitude of TQ results tend to be relatively low in the high frequency range and higher in the low frequency range. As mentioned in Section G3.3.1, the effect of this behavior on plant component analysis is addressed in detail in Section G4.0.
In order to view the dif ferences in response due to the three pressure traces, and also to see the effect or altering the treguency content by time scaling the curves, a set or comparison plots of ARS have been developed. The vertical response at the top of the pedestal due to an all valve discharge is used as representative of the behavior. Figures G-12 through 14 contain ARS plots developed for the three pressure traces, with the plot ror each curve identifying the general effect or time scaling (altering the f requency) . Two conclusions can be drawn from these results. First, in the frequency range above approximately 10 Hz, the response tends to be most severe for the higher frequency cases of each curve (tor the smal]er time scale tactors) . Also, it is noted that the response to curve No. 2 nearly envelopes the response to curves No. 1 and 3 at all treguencies. Curve No. 2 can be characterized as having the most regular shape of the three, generally resembling a decaying sinusoidal curve. Note that the amplitude multiplier or 1.1 has not been included in the ARS of Figs. G-12 through 14 which are intended only to depict the effects of time scaling (frequency shifting) and the dif f erences between the three curves.
Structural response has been determined for all three curves each with five time scale factors and the 1.1 multiplier. Enveloped ARS have been developed from the 15 individual ARS and used for plant component assessment in Section G4.0. Representative envelope ARS are presented on Figs. G-15 through 17 and compared to ARS from the sequential actuation and resonant sequential actuation RH cases as a basis of comparison.
G3.3.3 Plant Specific T-quencher Load Preliminary dynamic structural analysis had been performed for wall loads based on the Karlstien test data. Response spectra resulting from this preliminary analysis are comparable to that from the interim TQ load speciiication. It is expected that the Shoreham plant specific TQ load definition described in Section G2.3 will result in generally lower building response.
4 G-12 Revision 4 - February 1981 l
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G3.4 Conclusions This section has provided quantitative results of dynamic
-N_s
[} analyses of a typical Mark II reactor building complex tor an extensive number of SRV loading cases. The loadings considered have been selected to provide a substantial data base from which to address two major issues. These are the NRC's concern over the frequency and phasing of SRV bubbles and the relative effects of RH and TQ loads on building response.
Basic similarities are found in building response to all SRV loadings considered. A fundamental feature is that the response can be characterized by high and low frequency ranges with generally an inverse relation between magnitudes of response in the respective ranges. Results from the RH sequential actuation load definition tend to be the highest in the high frequency range and lowest in the low frequency range. At the other extreme is the TQ response with comparable magnitudes of response in both ranges, both being substantially less than peaks in the high frequency range from RH sequential actuation load. Results trom all other SRV loadings considered essentially are bounded between these two cases.
The ultimate ef f ect that these dif ferences in building responses have on plant components is discussed in detail in Section G4.0.
G4.0 COMPARISON OF PLANT COMPONENT RESPONSES
()
v To demonstrate that the lead plant RH load definition is conservative, this section compares the RH sequential and interim T' loads on plant components, and reveals the quantitative design rargins for components designed for RH loads when they are subjected to the TQ loads . Section G4.1 describes the results for piping subsystems and Section G4.2 describes the results for equipments.
G4.1 Piping Subsystems G4.1.1 Selection of Piping Subsystems for Data Base Nine piping subsystems were selected to cover a range of piping l
characteristics as follows:
- 1. Five out of the nine subsystems were selected trom those areas of severe SRV loadings.
- 2. Four additional subsystems were selected because of their low fundamental frequencies. From an inspection of the SRV amplified response spectra (ARS), it was expected that the low frequency piping subsystems would provide a lower bound on load reduction from use of the interim TQ.
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\ _ _ --
Table G-1 presents the nine piping subsystems considered in terms of building locations, piping sizes, and fundamental frequencies.
As can be seen in Table G-1, the piping fundamental frequencies range from 2.9 to 8.1 Hz , and they are located at various elevations within the primary and secondary containments.
G4.1.2 Analysis or Amplified Response Spectra As discussed in Section G-3, ARS have been generated for five RH load cases as well as for the interim TQ load case. An idealized ARS comparison for these load cases is shown on Fig. G-11. The actual ARS at structure support locations were used for the piping analysis. They may deviate from the idealized curves locally, however, a general trend was observed in the similarities and relative importance of these spectra curves as exemplified on Fig. G-11. To investigate their significance to piping components, the responses can be categorized into three frequency ranges:
- 1. Low frequency - Peaks around 5 to 10 Hz
- 2. High frequency - Peaks around 15 to 20 Hz
- 3. Very high frequency - Various peaks above 25 Hz The RH sequential load definition consistently produces the highest response in the nigh frequency range and generally the lowest response in the low frequency range. The interim TQ load derinition produces the lowest response in the high frequency range and a higher response in the low frequency range.
These two limiting cases were used to compare the effects on piping system designs. The very high treguency range is considered to have no decisive effects in all cases.
G4.1.3 Results and Discussions The ramshead-sequential (SRVRH ) and interim TQ (SRVTQ) are the two limiting cases to determine the ef fect of the piping analysis in this report.
Tables G-2 and 3 present the support load and pipe stress comparisons at three selected locations for three severe SRV subsystems and four low frequency subsystems identified in Table G-1. Since it is impractical to present detailed information for all piping locations, the following selection criteria, in order of priority, were used:
- 1. Locations of high stress.
- 2. Locations representing all types of piping components.
- 3. Locations evenly distributed along piping subsystem.
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TaDles G-2 AND 3 also show the ratios of SRVRH to SRVTQ. The ;
intent is to show the effect of using the RB or TQ load in one of 1 h
N/
the controlling load combinations (N + SRV + OBE) and also to identity the significance of TQ reductions.
An inspection of TaDie G-2 concludes that the ratios of SRVrh to SRVtg are sufficiently large. The ratios of pipe stresses trom the load combination N + SRV & OBE also show significant positive design margins, although the ratios are substantially reduced.
This is due to the presence of predominant internal pressure stress in the load combination. However, the ratios on pipe support loads remain significantly large as evidenced by the tabulated values.
The results trom the low frequency subsystems as shown in Table G-3 are not significantly different from Table G-2 tor other subsystems, even though the fundamental frequencies are in region where TQ load significantly exceeds RH load. In order to demonstrate the results are indeed accurate and correct, an extensive investigation on the dynamic response of piping subsystems to SRV load is made. Particular consideration is given to systems which have fundamental frequencies low enough to have one or more modes in the frequency range in which the TQ ARS may be greater than the RH ARS.
The modal response of a multi-degree of freedom system subject to dynamic support motion (such as a piping system subject to SRV actuation building vibrations) depends on two basic parameters.
("")
(,j The first is the modal participation factor which is a function of physical characteristics, e.l., geometry and mass distribu(tion. The second parameter is related to the amplitude (G 's) of the support acceleration at the modal frequency, i.e. ,
the ARS value. For a typical piping system with complex geometry (daree dimensional pipe routing, multiple bends, numerous interior pipe supports unevenly spaced, branch lines, etc. ) and many concentrated masses (valves, reducers, elbows, tees, equipment, etc.), the amplitudes of the modal participation
! factors are quite varied with significant f actors associated with l many modes f rom the very low to the very high frequencies. Large modal responses will occur at modes with both significant modal participation factors and significant ARS amplitudes. The total response is contributed to by many modes or comparable significance over a wide range of frequencies. For the complex piping systems in the Shoreham Nuclear Power Station, the contribution from the fundamental (lowest frequency) mode is of no special significance. It may be among the most or least important depending on the parameters discussed above.
Tables G-4 through 8 present the mode-by-mode contributions to resultant bending moments (acting on the pipe cross section) in l typical piping systems when subject to SRV building vibrations.
l Tnese systems include one ' higher
- frequency system (Table G-4)
~ tor comparison of behavior with the four ' low
- frequency systems l l
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(Tables G-S through 8) . The data presented on these tables are based upon the actual building ARS for both the RH and TQ loads.
These building ARS are applied simultaneously in the plant north-south, east-west, and vertical directions and vary from elevation to elevation within the reactor building. Figure G-11 is identified as an " Idealized ARS" to show the major f eatures of response. It should not be used to make quantitative comparisons for specific piping systems even though it does reflect overall trends. Inspection or the tabulated results (Tables G-4 through
- 8) reveals the following important points:
- 1. Many modes contribute significantly to the total response regardless of the fundamental frequency, the location of piping, or the discharge device.
- 2. The contribution from the fundamental mode is of no special consequence.
- 3. For the system with all f requencies above 7 Hz (Table G-
- 4) , all modal responses from the TQ load are less than the modal responses from the RH load.
- 4. For systems with lower f requencies, all modal responses below 7 Hz from the TQ load are greater than those from the RH load, while all modal responses above about 15 Hz are higher from the RH load. From 7 to 15 Hz results are comparable.
S. Systems with the lowest fundamental frequencies have only a small percent or their significant modes below 7 Hz.
- 6. Because many modes contribute to the total response, the RH load provides conservative results even when a few modes occur in the frequency range where TQ response is the greater.
The detailed analysis results presented above conclude that the RH load provides conservative results for the low frequency suosystems comparable to those for the other subsystems. The overall avcrage ratios of SRVTQ to SRVRH for the nine piping subsystems presented in Tabla G-9 further substantiate the above rindings. Tnis is the result of comparing load components (stres s , force, or moment) for the RH and TQ loads for each piping subsystem. The RH to TQ ratios in each subsystem cover a range or values that are both larger and smaller than the averaged values presented in the table.
The results from this piping subsystem study indicate:
- 1. Generally, RH is expeced to produce an overall average of 40 percent larger loads than the TQ on typical G-16 O
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i i
i piping. However, on low frequency piping systems, the percentage is reduced.
O
\- / 2. RH is expected to produce larger loads on most piping components. On those few remaining components where TQ load is larger, pipe stresses are generally not expected to be critical. This is because if a component gets larger response from TV load, it must have higher modal participation at low frequency. Since the vibration modes are more sparsely spaced in the low frequency range, and the amplitude of TQ excitation is low (see Fig. G-11), these result in a relatively low stress value for such components.
Based on the results of this analysis, it is concluded that the piping systems and components designed for RB sequential actuation load are more than adequate to accommodate the TQ load and, therefore, the design assesment basis SRV load definition is clearly conservative.
G4.2 Equipments G4.2.1 Introduction Since the previous sections demonstrate that the only potential tor the TQ results to exceed RH results is for components with low (less than about 7 Hz) fundamental frequencies, this will be the area concentrated on herein. Equipment qualified both by (j analysis and by test is addressed in the following sections.
G4.2.2 Floor-Mounted Equipment For Mark II load evdluations, the controlling UPSET and FAULTED condition load combinations used are N + SRSS (SRV + OBb) and N +
SRSS (SRV & LOCA + SSE) , respectively, where N refers to normal operating loads and SRV is the envelope of all SRV discharge cases, as stated previously. LOCA is also the envelope of all i LOCA subevents-vent clearing, condensation oscillations, and chugging. Regulatory Guide 1.61 damping values of 2.0 percent t
and 4.0 percent, respectively, are used for all loads in each combination.
Tne original seismic qualification was based on equipment damping values lower (more conservative) than that required by NRC Regulatory Guide 1.61. Damping values of 0.5 percent and 1.0 percent were used for UPSET and FAULTED conditions, respectively. This has resulted in a margin of safety in the low l frequency range which is sufficient to cover the combined seismic j and hydrodynamic results independent of the SRV discharge device.
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G4.2.2.1 Qualification by Analysis Mechanical equipment is requalified for Mark II loads by static analysis to equivalent dynamic loads h (static coefficients) obtained from enveloping the dynamic responses of several representative equipment models. A static coefficient curve resulting from the dynamic analyses related the equivalent static load on a component to the fundamental frequency of the component. The effects of the higher modes are thereby simply accounted for in subsequent evaluations. This method as applied to seismic analysis has been described in the Shoreham FSAR (Section 3.7.3.5.1 A) and as extended for Mark II loads in DAR(1)
(Section 9.1.2.3. ) .
Static coefficient curves for UPSET and EAULTED conditions have been redeveloped using the TQ load definition and are presented on Figs. G-18 and 19 for the primary containment and on Pig. G-20 and 21 for the secondary containment. The TQ based static coefficients exceed the rmashead based coefficients only for primary containment equipment with a fundamental frequency below about 3 Hz.
A review of Shoreham reactor building (primary and secondary containment) equipment indicates that there are no components which fall into this special situation where TQ results could exceed those from the RH. There is no floor-mounted equipment witn a fundamental frequency less than 3 Hz located in the primary ~ontainment and qualified by analysis. Equipment in the primary 7tainment includes cable tray, conduit, and duct systems _ aell as excess flow check valves. All of these components have f undamental f requencies greater than 10 Hz. For floor-mounted equipment in the secondary containment, the RH results bound TV results for all f requencies.
Furthermore, as stated earlier, any reactor building equipment in this low frequency range would be designed to higher static coetticients corresponding to lower seismic damping values. The level of the original design curves in this frequency range in shown on Figs. G-22 and 23 f or the primary containment. The same relationship is true in the secondary containment. As can be seen on Figs. G-22 and 23, there is a suf ficient margin of satety l in the low frequency range of the original design curve to cover l the effects of either the RH or TQ discharge device.
l G4.2.2.2 Oualification by Test l
Class 1E tloor-mounted electrical equipment in the secondary containment is qualified by test at the assembly level (rack, I panel, etc.), it possible, or ny a combination or device (relay, switch, etc.) testing with analysis to obtain the loads acting at l the device mounting base. The very small amount of 1E floor-
! mounted electrical equipment in the primary containment is qualified by single frequency testing.
1 i
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Input loading for testing is quantified fram the required response spectrum (RRS) corresponding to the location of the O equipment. For single frequency testing, the peak input acceleration must be equal to or greater than the acceleration at-the high frequency end of the spectra, that is, the zero period acceleration (ZPA). Inspection shows the ZPA of the RH spectra always to be larger than that of the TQ spectra. For multifrequency testing, the input acceleration time history must I
be shaped so chat the test response spectra (TRS) envelopes the RRS between th's equipment fluidamental frequency and the ZPA~
esymptote. ?;he means of developing the TRS is such that its magnitude over the entire frequency range is governed by the low -
- frequency paaks. This is demonstrated on Figs. G-24 and 25'for the lowest frequency (4 Hz) component in the secondary containment by multifrequency tersting. As can be seen, the TRS developed to encompass the original seismic peaks provides ample conservatism to cover all Mark II loads above 4 Hz, independent of the SRV discharge device.
It is thus shown that the- RH spectra is bounding for the qualification by test of Class 1E floor-mounted electric equipment in the reactor building.
G4.2.3 Pipe-Mounted Equipment The response of pipe-mounted equipment to seismic and hydrodynamic loads is detiermined by the dynamic response of the piping system (which includes the equipment component) . This nas been addressed in detail in Section G4.1. Specifically, since the piping system response is contributed to by a large number of modes, the higher TQ response at a few low frequency modes is not sufficient to cause the total stress from the TQ loads to exceed the total stress from the RH londs.
G4.2.4 Conclusion This section has specifically addressed the relative effects of RH versus TQ loads on reactor building equipment. The same trends that were found in the piping evaluation are also found for equipment. The predominant effect on both is an increasing relative importance of TQ loads with decreasing fundamental treguency of the component. For piping, it was shown that for low frequency systems, even though the TQ became relatively more important, the RH was still bounding. For equipment, it has been shown herein that for the special situation of floor-mounted equipment, with a fundamental frequency lev than 3 Hz located in the primary containment and qualified by analysis, the TQ results could exceed RH results. However, Shoreham has no equipment components in this situation and furthermoe, low frequency equipment in general has been designed to conservative seismic loads which bound both RH and TQ results.
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_.__,m___ _ _ _ _ . _ _ - - . . _ _ . _ _ . _ _ _ _ _ _ _ _ . . _ _ _ _ _ - - _ _ _
On this basis, it is considered to be evident that all reactor building equipment requalified to Mark II loads including the SRV RH load definition will be able to sustain the ef fects resulting from SRV discharge with a TQ device. This has been demonstrated herein using several conservative assumptions on the nature of the TQ load.
GS.O 'UMMARY
- AND CONCLUSIONS The fundamental conclusion of this report is that the Mark II lead plant RH load definition for SRV air clearing loads on the suppression pool boundaries is conservative and provides an dCCeptable basis for the design assesment of the lead plants.
All lead plants have chosen to install the TQ discharge Gevice which has been shown by testing to produce reduced air clearing loads. However, in order to expedite the licensing process, certain lead plants have proceeded, with NRC concurrence, to use the load prediction methodology based on the RH discharge device.
This report has provided quantitative evidence in support of the lead plant position. Using the RH load prediction methodology, pool boundary loads were developed for five load cases with widely varying assumptions on SRV bubble phasing and frequency characteristics. The assunptions employed encompass the maximum credible range of input condititas with respect to resulting pool boundary loadings. A detailed analysis of a representative lead plant subjected to these five SRV loadings has been performed and ARS of building accelerations were generated. A comparison of the resulting spectra reveals that even with these variations assumed in bubble phasing and frequency content, there are no ma jor diff erences in the resulting building responses.
Building response was also computed for a conservative TQ load definition. Included in this definition are the assumptions that dll Dubbles enter the pool simultaneously and oscillate in phase with a pressure amplitude 10 percent greater than any measured wall pressures in the Brunsbuttel in plant tests. In ad dition ,
the predominant frequency of oscillation is swept through a wide frequency range. The resulting enveloped ARS are significantly less than the ramshead response at hi;h frequencies and slightly higher at low treguencies. Preliminary confirmation of this TQ l load definition was performed by comparing structural response to it with that obtained from a conservatively constructed load definition based on the Karlstein test results.
In order to assess the impact of these results, several Shoreham piping systems were evaluated for the TQ respons e and results were compared with those resulting from the ramshead-sequential actuation case. Response spectra resulting trom all other SRV load cases considered are bounded by these two limiting cases.
Tne pipe stress analysis of several representative piping systems results in an overall reduction in both pipe stress and pipe support reactions from the use of the conservative interim TQ G-20 Revision 4 - February 1981 i
, loading as compared to RH loadings. The same conclusion is also found for Shoreham plant equipments.
The final plant specific TQ load definition, which addressed specif 4. 313y t-he Shoreham plant parameters, is expected to result in the sa:- enclusion that RH sequential load is Conservative.
This presen.: tion of quantitative results for actual plant i components subject to SRV loads in a typical Mark II containment is considered to be the most conclusive evidence in demonstrating the conservatism of the lead plant RH load definition used as the basis for the Shoreham design assesment.
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REFERENCES
- 1. Plant Design Assessment for SRV and LOCA Inads, Shoreham Nuclear Power Plant Station Unit 1, Revision 4, February 1981.
- 2. Mark II Containment Dynamic Forcing Functions Information Report (DFFR-Rev. 2) , NEDO-21061, September 1976.
- 3. Mark II Containment Lead Plant Program Load Evaluation and Acceptance Criteria, NUREG-0487, U.S. Nuclear Regulatory Consnission, October 1978.
- 4. Application of the SSES-Test Measurement Results to the Overall Loafing or the Suppression Chamber of the Shoreham Plant by Depressurization Processes - Revision 1, Kraftwerk Union Working Report R141/141/79/E, October 1979.
- 5. Supplement No. 1 to the Mark II Containment Lead Plant Program Load Evaluation and Acceptance Criteria, Supplement 1, NUREG-0487, U.S. Nuclear Regulatory Commission, Dratt, May 1980.
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l llh G-22 Revision 4 - February 1981 l
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,. - . _ -_ .-. -- ~ _ . . . - . . ~ . . _ - . . . . - . .--
s TABLE C-1 PIPING CHARACTERISTICS J Locatiar Pipe Size Fundamental Piping Subsystem (teet) finches) Frequency (Hz) Catment hedwa u 300 Primary El. 87 4 - 20 l 8.1 Severe SRV Secondary El. 56
, Peedwater 301 RPV El. 132 12 '98 7.1 bevere SRV !
Primary El. 87
@re Spray 100 RPV El. 132 10 7.6 Severe SW Primary El. 104 Secondary El. 60 min Steam 240 RPV El. 145 10 - 24 Primary El. 66 7.0 Severe SRV j Reactor Water Cleanup 012S Secondary El. 120 3-6 7.5 Severe SRV (bntrol Rod Drive 1265 Primary El. 100 3/g -1 Secondary El. 87 2.9 Low Frequency Closed Loop Cooling dater 031 Secondary El.12 6-8 Secondary El. 47 3.6 Iow Frequency Reactor Water Cleanup 013 Secondary El. 89 2-4 Secondary El. 136 4.9 Iow n equency Main Steam 2500 Primary El. 82 24 5.0 i
Secondary El. 82 Iow Frequency t
L
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TABLE G-2 TYPICAL RESULTS AT SELECTED LOCATIONS PART 1 SUPPORT LOAD COMPARISON Ramshead T-Ouencher Piping N+SRMUI+OBE N+SSRVTQ +OBE SRMU{
Subsystem Component (lb or ft-lb) (lb or ft-lb) SRVrq Feedwater Anchor 3 48,982 38,331 1.44 301 Resultant Moment Snubber 98 21,432 13,224 1.62 Axial Force Anchor 125 244,800 191,600 1.48 Resultant Moment Core Srpay Anchor 1 23,867 21,651 1.17 100 Resultant Moment Snubber 23 4,847 4,070 1.22 Axial Force Anchor 118 28,603 19,676 1.58 Resultant Moment Reactor water Anchor 30 8,672 6,901 1.85 Cleanup Resultant Moment 012S Snubber 42 4,024 2,485 1.65 Axial Force Anchor 180 1,953 O Resultant Moment 1,812 1.21 PART 2 PRIMARY STRESS INTENSITY COMPARISON Piping N+SRV}ul+OBE N+SRWq +OBE SRVlui Subsystem Component (psi) (psi) SRVTQ 2eedwater Reducer 65 16,482 14,795 1.60 301 Tee 67 16,205 14,333 1.29 Elbow 84 23,373 20,625 1.36 Core Spray Elbow 9 20,091 19,378 1.12 100 Valve 86 17,507 13,574 1.65 Elbow 92 23,822 16,568 1.78 i
Reactor Water Tee 63 14,973 13,412 1.35 Cleanup Elbow 75 18,458 16,370 1.24 012S Tee 132 21,135 17,886 1.26 i
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l 1 of 1 Revision 4 - February 1981 l
TABLE G-3 TYPICAL RESULTS AT SELECTED LOCATIONS Part 1 Support Load Comparison Piping N+SR N+SRVpq+0BE SRVnu Subsystem Component _(lb o]g r f+0BE t-lb) (lb or it-lb) SRV""
Control Rod Restraint 65 42 35 1.44 Drive 1265 Resultant Force Restraint 85 16 13 1.40 Resultant Force Restraint 110 34 33 1.09 Resultant Force Closed Loop Restraint 95 1,815 1,801 1.08 Cooling Water Resultant Force 031 Restraint 175 2,041 2,013 1.18 Resultant Force Anchor 5 3,131 3,052 1.17 Resultant Moment Reactor Water Anchor 184 649 424 1.91 Cleanup Resultant Moment 013 Snubber 186 567 417 1.40 Axial Force Oi Snubber 623 Axial Force 338 279 1.44 Main Steam Anchor 340 147,000 135,000 1.11 2500 Resultant Moment Restraint 365 33,000 30,000 1.12 Resultant Force Snubber 400 13,000 10,300 1.31 Axial Force 3
i 1 of 2 Revision 4 - February 1981
,e. -rw y , - . . < ,w -. -- -- +--.w -
- w. vy
TABLE G-3 (CONT'D)
() Part 2 Primary Stress Intensity Comparison Piping N+SRMut+OBE SRVuu Subsystem Component (psi) H+SRVrq(psi +) 0BESRVyf Control Rod Valve 1 8,076 7,631 1.08 Drive 1265 Elbow 50 4,353 4,176 1.07 Run 185 6,658 6,355 1.17 Closed Loop Elbow 10 4,720 4,689 1.07 Cooling Water Tee 205 3,078 2,932 1.16
- 031 Run 275 5,706 4,645 1.38 Reactor Water Valve 151 8,206 7,319 1.50 Cleanup Run 919 7,795 7,280 1.39 013 Reducer 339 10,207 9,807 1.22 Main Steam Valve 345 8,793 8,409 1.23 2500 Elbow 415 8,893 8,783 1.27 Run 440 13,161 12,978 1.22 O
l i
1 l
1 i
O 2 of 2 Revision 4 - February 1981
TABLE G-4 O FW 301 MODE-BY-MODE CONTRIBUTION RESULTANT MOMENTS REDUCER 65 TEE 67 ELBOW 84 FREQUENCY RH TQ RH TQ RH TQ MODE (HZ) (f t-kip) (f t-kip) (ft-kip) (ft-kip) (f t-kip) (f t-kip) 1 7.1 1.5 0.9 0.6 0.4 0.9 0.5 2 3.0 8.2 8.1 16.6 16.4 12.2 12.0 3 9.4 ca) 1.2 1.2 4.4 4.3 2.9 2.9 4 11.9 10.3 7.4 4.5 3.2 6.3 4.5 5 13.7 4.4 3.4 4.0 3.1 3.2 2.9 6 16.4 19.2 5.9 9.4 2.9 4.7 1.5 7 19.6 6.9 4.7 3.7 2.5 5.1 3.4 8 22.7 10.7 7.3 7.5 5.2 2.0 1.4 9 24.4 3.6 3.2 1.9 1.7 2.3 2.0 10 26.3 4.7 2.6 4.6 2.6 0.9 0.5 11 26.5 0.8 0.5 0.7 0.4 0.1 0.1 12 28.6 2.5 1.6 8.7 5.5 4.5 2.8 13 30.3 5.9 3.3 6.4 3.6 10.4 5.9 14 32.3 (3) 2.0 1.4 1.2 0.8 3.2 2.2
- 15 33.3 7.9 5.1 13.2 8.6 8.8 5.7 16 34.5 6.7 3.8 0.6 0.3 9.7 5.4 17 37.0 2.9 2.3 1.8 1.4 0.3 0.2 18 38.5 2.8 2.2 2.9 2.3 4.6 3.6 s 19 40.0 4.1 3.0 4.5 3.3 5.7 4.1
\
20 40.0 1.8 1.2 0.6 0.4 0.9 0.6 21 43.5 3.8 2.9 1.0 0.6 1.3 0.9 24 58.8 1.4 0.9 1.8 1.2 0.8 0.6 25 58.8 0.6 0.4 0.3 0.2 0.2 0.2 SRSS ( * ) 31.1 18.7 29.1 22.3 25.2 18.6 Above Combine (5) 35.6 22.3 36.7 28.0 32.2 23.6 All Modes i
(1) Low frequency modes below Hz, TQ>RH.
(2) Intermediate frequency modes between 7 to 15 Hz, T@eRH.
> (3) High frequency modes above 15 Hz, RH>TQ.
(*) For reference only, contributions from other modes are less significant.
(5)
Basis for stress calculation, moments are combined by Reg.
Guide 1.92 Grouping Method.
O 1 of 1 Revision 4 - February 1981
TABLE G-5 CRD 1265 MODE-BY-MODE CONTRIBUTION RESULTANP POME?rrS VALVE 1 ELBOW 50 RUN 185 FREQUENCY RH 'IQ RH TQ RH TQ MODE CHZ) (f t-lb) (ft-lb) (f t-lb) (f t-lb) (f t-lbl (f t-lb) 1 2.9 0.0 1.1 0.0 0.0 0.0 4.5 2 5.1 (*) 1.0 9.0 0.0 2.4 0.0 4.5 3 6.6 2.2 9.0 1.0 2.2 2.0 9.9 4 8.0 2.2 2.4 0.0 0.0 3.6 5.0 5 10.0 7.3 15.3 2.0 3.5 2.2 5.9 6 11.8 (a) 7.1 19.0 2.0 4.4 1.4 4.1 7 12.5 12.2 23.1 3.2 4.5 2.0 2.2 8 15.2 36.9 16.5 7.3 3.5 16.5 8.0 9 16.7 2.0 1.5 1.0 1.1 1.0 1.1 10 19.2 1.4 1.5 0.0 0.0 1.0 1.1 11 25.0 0.0 0.0 0.0 0.0 3.0 1.9 12 27.0 11.2 5.6 1.0 1.1 3.2 2.2 13 27.8 3.6 1.5 0.0 0.0 7.0 2.2 14 32.3 (3) 11.9 6.1 0.0 0.0 1.0 1.1 15 37.0 11.2 8.9 1.0 1.1 0.0 0.0 16 38.5 12.5 9.4 1.0 1.1 1.4 0.0 17 47.6 1.0 0.0 0.0 0.0 7.0 5.5 N 18 50.0 1.0 1.1 0.0 0.0 3.0 2.2 19 50.0 2.2 1.1 1.0 1.1 2.0 2.2 20 58.8 3.3 2.4 1.0 1.1 1.0 0.0 21 62.5 20.3 16.8 9.1 7.8 1.0 1.1 22 66.7 4.2 2.2 1.0 1.? 6.0 0.0 SRSS(*) 51.4 46.0 12.7 12.0 20.9 18.5 Above Combine (5) 53.8 50.8 13.6 12.9 22.8 19.6 All Modes NOTES:
See Table G-4 for Footnotes.
O 1 of 1 Revision 4 - February 1981
TABLE G-6 O CLCW 031 MODE-BY-MODE CONTRIB1TPION RESULTANT MOMENTS ELBOW 10 TEE 205 RUN 275 FREQUENCY RH TQ RH TQ RH TQ MODE (HZ) (f t-lb) (f t-lb) (ft-lb) (f t-lb) (f t-lb) (f t-lb) 1 3.6 10. 61. 4. 19. 4. 19.
2 4.3 12. 85. 5. 32. 7. 45.
3 4.9 (1) 11. 113. 11. 122. 17. 183.
4 5.5 21. 288. 6. 83. 5. 76.
5 6.1 7. 45. 105. 710. 9. 55.
6 8.1 72. 108. 31. 46. 7. 11.
7 8.9 153. 286. 64. 120. 21. 39.
8 9.8 36, 37. 44. 46. 46. 47.
9 10.0 65. 70.. 19. 20. 28. 29.
10 11.2(a) 112. 101. 24. 28. 12. 13.
11 11.8 283. 350. 16. 20. 22. 28.
12 12.7 54. 63. 51. 58. 21. 25.
13 14.1 301. 274. 138. 125. 24. 22.
14 14.8 210. 141. 102. 69. 35. 23.
15 17.5 465. 256. 9. 6. 3. 2.
16 18.2 245. 180. 2. 2. 5. 3.
17 19.2 6. 4. 1041. 711. 1280. 867.
O 18 22 24 24.0(3) 30.3 32.7 2.
4.
1.
1.
3.
1.
740.
608.
17.
507.
446.
13 .
425.
2698.
60.
292.
1977.
48.
SRSS(*) 736. 734. 1433. 1238. 3018. 2190.
Above Combine (5) 1004. 935. 1468. 1262. 3055, 2220.
All Modes l
l NOTE:
See Table G-4 for Fbotnotes.
l O
1 of 1 Revision 4 - February 1981
O 'M (j (% C'1 w/
l.
JALLE G-7 Rwcu 013 MODE-isY-MODP. CON'1%IBUTION RESULTANT MOMENTS VALVE 151 RUN 919 hEDUCER 339 Frequency RH W Rif W Eli W
.; Ptxies (Hz) (f t-lb) (f t-lb) (ft-lb) ift-Ib1 tit-lb) tit-lb)
- 1 4.9 0. 2. 6. 83. 6. 83.
- 2 6.5 0. O. O. O. O.
4 O.
3 6.6 (*) O. 1. 10. 43. 10. 44
] 4 6.8 0. O. O. O. O. O.
5 7.7 1. 1. 63. 58.- 17. 17.
I 6 7.9 6. 11. O. O. O.
! 7 O.
8.0 0. 9. O. O. O. O.
8 9.0 4 4. 44. 53. 28. h.
9 11.1 1. 2. O. O. 0.. O.
10 11.4 (a) 3. 3. O. O. O. O.
11 11.8 76. 147. O. 1. O. O.
12 12.5 14. 17. O. O. O. 0.
13 12.8 19. 28. O. 'O. O. U..
4 15 14. 1 8. 8. 22. 21. 13. 11.
e 17 15.t> 127. 50. 35. 13. 13. 4.
18 16.1 138. 61. 79. 34. 24 11.
20 16.8 8. 3. 66. 26. 30. 12.
I 19.4 23 138. 74. 15, 8. 3. 1.
25 21.6 381. 196. 23. 11. - 2. 1.
1 27 22.9 71. 44 108. 66. 4. 2.
28 23.0 ?60. 111. 73. 51. 20. h.
29 24.1 235. 169. 254. 184. 15. 11.
} 30 25.0 37. 25. 29. 20. 108.
2 73.
31 25.0 (*) 61. 42. 24. 15. 63. 43.
32 26.6 110. 84. 188. 143. 8. 7.
35 28.6 291. 198. 18. 12. 2. 1.
37 31.3 6. 3. 42. 28. 11. -7.
39 32.3 10. 4. 22. 14. 12.
- 7.
42 37.0 78. 57. 45. 33. 21. 15.
44 38.5 19. 14. 6. 4. 29, 24 47 43.5 68. 60. 5. 4. 2. 1 48 43.5 174. 125. 21. 15. 5. 3.
51 47.6 78. 65.
57
- 3. 2. 2. 1.
55.6 2. 1. 7. 4. 55. 37.
SRSS(*)
Above 664 437. 378. 286. 155. 143.
Combine " ) 938. 615. 466. 334. 213. 173.
All Modes
- tDTBS
See Table G-4 for Footnotes.
1 of 1 hevision 4 - February 1981
i l
l TABLE G-8
) MS 2500 MODE-BY-MODE CONTRIBUTION RESULTANT MOMENTS Valve 345 Elbow 415 Run 440 Frequency RH TQ RH TQ RH TQ Mode (Hz) (f t-kip) (f t-kip) (f t-kip) (ft-kip) _(ft-kip) (f t-kip) 1 5.1 (1) 0.2 0.8 1.1 4.1 2.8 10.9 2 13.5 4.6 5.3 3.8 4.3 3.3 3.5 3 14.5 (2) 3.5 6.0 0.6 1.0 0.3 0.5 4 15.2 12.4 20.0 0.9 1.5 0.6 0.9 5 16.7 4.5 3.3 18.4 13.4 14.9 10.9 6 17.9 39.3 23.7 0.3 0.2 0.2 0.1 7 20.8 2.1 1.2 9.2 5.2 9.3 5.3 8 23.8 8.2 5.2 9.0 5.7 10.2 6.5 9 26.6 36.5 17.7 2.7 1.3 3.6 1.7 10 29.4 22.2 31.4 1.3 1.8 0.4 0.6 11 3C.3 (3) 1.9 2.1 1.2 1.4 1.6 1.8 12 35.7 3.7 2 .7 22.0 16.2 15.3 11.3 13 37.0 6.8 5.4 20.3 16.0 14.3 11.3 14 43.4 0.9 0.6 2.9 2.3 27.7 21.4 15 43.5 0.1 0.1 1.3 1.1 44.2 37.7 16 43.5 0.8 0 .6 3.1 2.5 27.3 22.3 17 45.5 0.8 0.6 6.4 5.5 2.6 2.2 0' 18 19 52.6 55.5 2.5 2.3 2.2 2 .1 15.0 15.1 12.8 12.3 2.1 3.8 1.8 3.1
(*)
SRSS Above 61.1 49.2 44.0 34.1 66.2 54.6 ss)
Combine All 63.5 51.8 48.4 38.0 105.7 86.7 Modes NOTES:
See Table G-4 for Footnotes.
w 1 of 1 Revision 4 - February 1981
TABLE G-9 O OVERALL AVERAGE RATIOS OF (SRVRH ) (SR M PART 1 SEVERE SRV SYSTEMS Piping OVERALL AVERAGE RH/TO RATIO Subsystem Pipe Support Loads Pipe Stresses Feedwater 300 1.42 1.31 Feedwater 301 1.49 1.49 Core Spray 100 1.34 1.41 r Main Steam 240 1.54 1.48 Reactor Water Claanup 012S 1.26 1.40 TOTAL 1.41 1.42
? ART 2 LOW FREOUENCY SYSTEMS Control Rod Driver U 65 1.24 1.22 Closed Loop Cooling Water 031 1.05 1.17 :
Reactor Water Cleanup 013 1.48 1.41 Main Steam 2500 1.16 1.24 1.23 1.26 TOTAL NOTES:
RH = Ramshead - Sequential TQ = Interim T-Quencher O
1 of 1 Revision 4 - February 1981
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FIGURE G-2 STRUCTURAL MODEL NODAL POINT NUMBERING SCHEME SHORE NUCLEAR POWER STATION-UNIT I PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 4-FEBRUARY 1981 [
SUPPORT PRIMARY
'h PEDESTAL CO N TAIN M ENT
(' )
v ,:. e l .
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PRESSURE
( LB./ IN.2 I SUPPRESSION POOL N
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(^3 MUMENT(FT .K/FT. )
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".'.w*' k IDEALIZED NODAL FORCES AND MOMENTS FIG. G -3
/^'s LOADS APPLIED TO MODEL (s) SHOREHAM NUCLEAR POWER STATION - UNIT I PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 4-FEBRUARY 1981
O O O n
i f
i z i 9
y 2 3
x
, d 4 5
l d 5 0 4 4
3 2
1 i
i i e i , , _ ,
0.00 0.o2 o.04 o.os o.08 o to 0.12 0.14 0.16 0.18 0.2o I PERIOD (SEC)
LEGEND D ALIZED ARS COMPARISON:
- 2. SfuulTA oUS D S7HARGE
- 3. RESONANT SEQUENTIAL ACTUATION RAMSHEAD ALL SRV DISCHARGE
' 4. SIMULTANEOUS ACTU ATION SHOREHAM NUCLEAR POWER STATION - UNIT 1 i
REVISION 4-FEBRUARY 1981 1
~
(
U) 3 00 2
2.50 -
SEQUENTIAL ACTUATION j ESONANT SEQUENTIAL ACTUATION 2.00 4
z -
jT 9 j\ RESONANT CHIFTED FREQUENCIES
$ I e w i.30 -
l g /\\
, w I /
8 I I t.OO -
f T,U /
. / s n \- /
0.50
.1
^ / /. w.m y ,_-
j
/ % %. -
~-
3, . .
0.00 O O2 O 04 0.06 0.08 0.10 0 .12 0.14 0.16 0.18 0.20 0.22 0.24 PERIOD IN SECONDS
.1 FIGURE G-5 PEDESTAL ARS COMPARISON.RAMSHEAD NODE 161-PEDESTAL ATEL.90 FT. ALL SRV DISCHARGE (CASES 1,2 S4) t% OSCILLATOR DAMPtNG SHOREHAM NUCLE AR POWER STATION-UNIT 1 VERTICA L- ARS BY TIME HISTORY
O O O 3.00 1 4 2.50 -
I ll SEQUENTIAL ar7':ATION II \ l\
\\ I\
l SIMULTANEOUS DISCHARGE 2.00 -
I I z I.jil\\ l
{
g O
P If 5' I SIMULTANEOUS ACTUATION
< gll l
, I i
ll \\ !I r1 b '\
)I 1 00 -
1 f ll {ylf t-0.50 -
) %
GQ '
000 ------ -----
o.00 002 0.04 oos oos o lo 0.12 o 14 o.ls o 18 o.2o 022 o.24 PERIOD IN SECONDS NOTE. FIGURE G-6 NODE 16 5-PRIM ARY CONTAINMENT AT ell 37' 1% OSCILLATOR DAMPlNG ALL SRV DISCHARGE (CASES 1,2 & 4)
VERTICAL- ARS BY TIME HISTORY SHOREHAM NUCLEAR POWER STATION-UNIT 1 PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISloN 4-FEBRUARY 1981
O O O i.20 1.00 -
SEQUENTIAL ACTUATION 8 f p I O.60 -
I 1 SIMULTANEOUS DISCHARGE I
y \ l 4
0.40 -
g i SIMULTANEOUS ACTUATION r- ys ; i
'v%. V u* .r. ,
0.20 s j /
v .\.
y.s~_____ --
~. . - __.--_ ____
, , , . . . . . . . - . . . . . . _ _ =
000 0.02 O 04 0 06 O 08 0.10 0.12 0.14 OJ6 0.18 020 0.22 PERIOD IN SECONDS NOTE- GE G 4 NODE 149-SECONDARY CONTAINMENT AT EL.73 FT.
1% OSCILLATOR DAMPING '
~
VERTICAL-ARS BY TIME HISTORY PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 4-FEBRUARY 1981
O O O 3.00 SEQUENTI AL ACTUATION 2.50 -
A RESONANT SHIFTED FREQUENCIES 2 00 -
/
, z l 9
. n g.
$ i 30 _ I I j RESONANT SEQUENTIAL ACTUATION
$ b 8
a l! /
\
sj
\
l00 -
I li i s
s
^g t i N
/ ' 'i .m
\
\ 3 s N-^j\ s s j / N .A s's 0 50 -
(.
b y- \ %
) i
' Q ' ~ ~~ ~ ~ ~, _ - - -
i t i
i t' 0 00 t t t t t 0 00 002 0.04 0 06 008 Q 10 0 12 0 14 0 16 0 18 0.20 0 22 0.24 PERIOD IN SECONDS 2
N TE: FIGURE G-8 PEDESTAL ARS COMPARISON:RAMSHEAD
' ALL SRV DISCHARGE (CASES 1,2 & 4) kOSci AT R DAMPIN SHOREHAM NUCLEAR POWER STATION-UNIT 1 VERTICAL-ARS BY TIME HISTORY PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS
- REVISION 4-FEBRUARY 1981
O O O 6 00 i
5 00 -
4 00 - RESONANT SHIFTED FREQUENCIES z
9 RESONANT SEQUENTIAL ACTUATION
! <x 3 00 -
SEQUENTIAL ACTUATION d
8 *4 j.
2.00 -
t I\
, 100 -
\
If\ \
l \s %-
000 t i i Q2 WL.
t i
'~-~~ ~ .
0 00 0.02 0.04 0.06 0 08 0.10 0 .12 0 .14 0.16 0.18 0 20 0.22 0.24 PERIOD IN SECONDS l
FIGURE G-9
, NOTE: PEDESTAL ARS COMPARISON: RAMSHEAD NODE 165-PRIMARY CONTAINMENT AT EL.137 FT ALL SRV DISCHARGE (CASES 1,3 a 5) 1% OSc!LL ATOR DAMPING SHOREH AM NUCLE AR POWER STATION-UNIT I VERTICAL. ARS BY TIME HISTORY PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 4-FEBRUARY 1981
O O O I
f.2O RESONANT SHIFTED FREQUENCIES i 00 -
,f-RESONENT SEQUENTIAL ACTUATION z O.80 -
.! SEQUENTI AL ACTUATION e l r
& I -
I x -
l\
w 0.60 a
l \. Ig w N \ li y j f I.
O40 - -
I \ l \
l f~
0 20
&. ,I D- _.
-- =- - ~ .- __ __ _ _ _ _ _
__ , ,. - - ~.
0 00 _______
O 00 O O2 0.04 0.06 0.08 0.10 0.12 0.I4 0 16 0.18 0.20 0.22 0 24 PERIOD IN SECONDS NOTE.
FIGURE G-lO SECONDARY CONTAINMENT ARS COMPARISON:
NODE 149-SECONDARY CONTAINMENT EL.73 FT RAMSHEAD ALL SRV DISCHARGE (CASES 1,3 SS) 3% OSCILL ATOR DAMPING SHOREHAM NUCLEAR POWER STATION-UNIT 1 i VERTICAL- ARS BY TIME HISTORY PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS i
REVISION 4-FEBRUARY 1981
O O O 5 2
- p
! 4 3
@ 4 5
d o
o 5 6
< 4 3
2 1
i 0.00 i
0.02 i
0.04 e
0.06 i
0.08 i
0.10 K'D 0.12 0.14 0.16 0.18 0.20 PERIOD (SEC)
LEGEND
- 1. SEQUEt.TIAL ACTUATION
- 2. SIMULTANEOUS DISCHARGE FIG. G-il
- 3. RESONANT SEQUENTI AL ACTUATION IDEALIZED ARS COMPARISON:
- 4. SIMULTA NEOUS ACTU ATION RAMSHEAD AND T-QUENCHER ALL SRV DISCHARGE
- 5. RESON ANT SHIFTED FREQUENCIES SHOREHAM NUCLEAR POWER STATION - UNIT 1
wJ J ~.J 3.00 1.0 PERCENT OSCILL ATOR DAMPING 2.50 -
8 200 -
1.0: TIME SCALE FACTOR 1.2 = TIME SCALE FACTOR b 1.4 : TIME SC ALE FACTOR
{e t . 50 -
1.8 = TIME SCALE FACTOR U
u o
4 1.00 -
T
- l. \
j
[
j vw \
oso -
p W& \
- ~s ~.~.m o.00 I I I I I I I I I I l 0.00 002 0.04 0.06 0.08 0.10 0.12 o 14 o.is o.18 0.20 022 0.24 PERIOD (SECOND )
tJOTE'. THE PRESSURE MULTIPLIER l .ED HERE IS EQUAL To 1.o FIG. G- 12 AMPLIFIED RESPONSE SPECTRA OFVERTICAL ACCELERATION TOP OF REACTOR SUPPORT PEDESTAL T-QUENCHER ALL SRV DISCHARGE-PRESSURE TRACE NO.1 SHOREH AM NUCLEAR POWER STATION-UNIT 1 PL ANT DESIGN ASSESSMENT FOR SRV AND LOC A LO ADS REVISION 4-FEBRUARY 1981
O O O 3.00 1.0 PERCENT OSCILL ATOR D AMPlNG 2.50 -
S 2.00 - 1.0= TIME SCALE FACTOR g 1.2 = TIME SCALE FACTOR
- 1.4r TIME SC ALE FACTOR 4 1.8= TIME SCALE FACTOR
$ 1.50 -
d 8
1.00 -
. ^-
/ N 4
$f ./h - --. '\.
s\ ~
~~~~;
o 00 I I I I I I I I I I I o.00 002 0.04 0.06 0.08 0.10 0.12 0.14 0.16 0.18 0.20 0 22 024 PERIOD (SECOND)
NOTE: THE PRESSURE MULTIPLIER USED HERE IS EQUAL To 1.0 FIG G-13 AMPLIFIED RESPONSE SPECTRA OF VERTICAL ACCELERATION TOP OF REACTOR SUPPORT PEDESTAL T-QUENCHER ALL SRV DISCHARGE-PRESSURE TRACE NO.2 SHOREH AM NUCLEAR POWER STATION-UNIT 1 PL ANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 4-FEBRUARY 1981
O O O l
3.00 1.0 PERCENT OSCILLATOR DAMPING 2.50 -
E 2.00 1.0= TIME SCALE FACTOR 1.2= TIME SCALE FACTOR
$ 1.4 = TIME SC ALE FACTOR j p 1.8 = TIME SCALE FACTOR i $ 1.50 -
i w
! O l < 1.00 -
1 f* .
i
/ N N O.50 -
/ .
s'X -
\.
N s N s U== P I I I I ! I I I I l 0.00 I I 0.00 0.02 0.04 0.06 0.08 0.10 0.12 0.14 0.16 0.18 0.20 0.22 0.24 PERIOD (SECOND) i NOTE: THE PRESSURE MULTIPLIER USED HERE IS EQUAL TO 1.0 FIG. G- 14 AMPLIFIED RESPONSE SPECTRA OF VERTICAL ACCELERATION
, TOP OF REACTOR SUPPORT PEDESTAL T-QUENCHER ALL SRV DISCHARGE-PRESSURE TRACE NO.3 SHOREH AM NUCLEAR POWER STATION-UNIT 1 PL ANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 4-FEBRUARY 1981 i
i O O O 3.00 250 - SEQUENTI AL ACTUATION RESONANT SEQUENTI AL ACTU ATION g INTERIM T-QUENCHER LOAD DEFINITION
- 200 -
l 9 I Q l E
i.s o -
l d I O I W l '
A. '
too -
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,/ \
e I (.s. - h~._.
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9 050 ,/ . s \
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oco I I I I I I I I 'I~+--*--
o.co o.o2 o.o4 o.os o.os o.io o.52 oi4 o.is o.is o.2o o.22 o.24 PERIOD IN SECS.
FIG. G-15 PEDESTAL ARS COMPARISON:
NOTE' R AMSHEAD AND INTERIM T-0UENCHER NODE 16 4- PEDESTAL AT EL.9o FT ALL SRV DISCHARGE 1% osclLLATOR DAMPING SHOREHAM NUCLEAR POWER STATION-UNIT 1 VERTICAL- ARS BY TIME HISTORY PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 4-FEBRUARY 1981
~ '
i 3.00 E50 -
SEQUENTIAL ACTUATION f1 RESON ANT SEQUENTIAL ACTUATION
\
G I g INTERIM T-QUENCHER LOAD DEFINITION g 2.00 -
I I
o E l N 1.50 -
\ p l 1
l ef '
f, \
S l \ b \
1.00 -
\
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ooo i I I I I 4%x%- %y I I t-
~ ~ L - - l_ _ _l _ _ _
o.oo o.02 o.04 0.06 0.08 0.10 0.12 o.14 o.is o.t e o.20 0.22 0.24 l
PERIOD IN SECS.
! FIG. G- 16 NOTE:
PRIMARY CONTAINMENT ARS COMPARISON:
R AMSHEAD AND INTERIM T-QUENCHER NODE is5 - PRIMARY CONTAINMENT EL.137 FT.
ALL SRV DISCHARGE
- 1% OSCILLATOR DAMPING SHOREHAM NUCLEAR POWER STATION-UNIT 1 i
VERTICAL- ARS BY TIME HISTORY PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS 4
4 REVISION 4-FEBRUARY 1981 4
4
O O O i 1.20 1.00 -
SEQUENTI AL ACTU ATION
$ 0 80 -
RESON ANT SEQUENTI AL ACTUATION z
o g f INTERIM T-QUENCHER LOAD DEFINITION Q l1
$ o60 -
I' d ,
1 N II l o.4o -
/g p\ l f~J %
/ '
0'20 I f
Y, /. %
-&/ N
~
k k.
[' .h.
~/
o.oo I I I I _T ' H - -- - - -
O.oo 0.02 0.04 o.06 0.08 0.10 o.12 o.14 0.16 o.18 o.20 o.22 0.24 PERIOD IN SECS.
FIG. G- 17 SECONDARY CONTAINMENT ARS COMPARISON:
NOTE' RAMSHEAD AND INTERIM T-QUENCHER NODE 149 -SECONDARY CONTAINMENT EL.73 FT. ALL SRV DISCHARGE 1% oSclLL ATOR DAMPING SHOREHAM NUCLE AR POWER STATION-UNIT 1 VERTICAL - ARS BY TIME HISTORY PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LO ADS REVIStoN 4- NOVEMBER 1980
O O O ll 5 -
i HORIZONTAL- RH i
4 -
o
[ HORIZONTAL-TQ ,
3 -
52
'i y u
VERTICAL-RH U
2 s I \
i i
~ ~
V ERTICAL -T-O
/ \ / N i _/ \ /
'J i '
O I I I I 2 5 10 20 30 FREQUENCY ~ HZ FIG. G-18 STATIC COEFFICIENTS-PRIMARY CONTAINMENT UPSET CONDITION N+OBE+SRV SHOREHAM NUCLEAR POWER. STATION-UNIT I PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 4-FEBRUARY 1981
.a.
- O O O i
1 5 -
1 5
4 -
p HORIZONTAL-R H o
2 z3 -
y HORIZON TAL - T- O E N D g F ,
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o 2 -
N ~l VERTICAL-RH P \
@ \ c- -
K
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i -
( VERTICAL - T-Q o i I I J 2 5 10 20 30 FREQUENCY ~ HZ FIG. G-19 STATIC COEFFICIENTS-PRIMARY CONTAINMENT FAULTED CONDITION N +SSE + SRV+ LOCA SHOREHAM NUCLEAR POWER STATION-UNIT I PL ANT DESIGN ASSESSMENT ' FOR SRV AND LOCA LOADS REVISION 4 FEBRUARY 1981
O O O 2.5 -
4 2 -
1 y
o --
l.5 -
3 0
o / \
b f k HORIZONTAL -RH a
o \ l 6, 9 I -f k y g g- HORIZONTA L -T-Q 5
4
/-Q VERTIC AL- RH
.5 -
~----~ / -4 7 -
VERTICAL-T-Q '
I I I I O
a 2 5 10 20 30 FREQUENCY ~ HZ FIG. G-20 STATIC COEFFICIENTS-SECONDARY CONTAINMENT i UPSET CONDITION N+0BE+SRV.
SHOREHAM NUCLEAR POWER STATION-UNIT I PLANT DESIGN ASSESSMENT FOR SRV AND l.OCA LOADS REVISION 4- FEBRUARY 1981 i
O O O 2.5 -
I I
- 2 t
.* f T I\
. 6,5 _ l\
4 e / \
9 / \
gHORIZONTAL -RH I
I N /v (N (,
l O I -
H ORIZONTAL- T-Q
< \g --
\ VERTICAL- RH t_
5 -
VERTICAL - T-Q i
, O I I I I 2 5 10 20 30 FREQUENCY ~ HZ FIG. G-21 STATIC COEFFICIENTS-SECONDARY CONTAINMENT FAULTED CONDITION N +SSE + SRV + LOCA SHOREHAM NUCLEAR POWER STATION-UNIT I PL ANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS 4
REVISION 4-FEBh0ARY 19{
. O O O 5-HORIZONTAL - RH 4 -
HORIZONTAL -SEISMIC DESIGN CURVE i -M 4 o HORIZONTAL - T- Q M
2
\ c VERTICAL - SEISMIC z3 -
DESIGN CURVE W
g - - -
1 g
v N -
r - -- - -- - -
, t V E RTICAL - RH l
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- , _ /
/ \
\ /
/
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i 2 5 10 20 30 FREQUENCY ~ HZ FIG. G-22 STATIC COEFFICIENTS- PRIMARY CONTAINMENT UPSET CONDITION DESIGN CURVES SHOREHAM NUCLE AR POWER STATION -UNIT I PL ANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 4-FEBRUARY 1981
O O O 5 -
HORIZONTAL-SEISMIC DESIGN CURVE 4 -
HORIZON TAL- RH VERTICAL-SEISMIC DESIGN CURVE z
3 o
[ HORI:'ONTAL-TO 4
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$ \
~
r-- - - - - - - - 4 o x NI \ l o2 - \ l I p - - -
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N g VERTICAL-T-Q
- \ / N 1 -
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.i o I I I ~ l 2 5 10 20 30 FREQUENCY ~ HZ FIG. G-2 3 STATIC COEFFICIENTS-PRIMARY CONTAINMENT FAULTED CONDITION DESIGN CURVES SHOREHAM NUCLEAR POWER STATION-UNIT 1 PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 4-FEBRUARY 1981
O O O 10 9 -
e -
y .
s -
5 --
4 -
3 .
2 .
e .
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z 9 } \
ii a
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bs \
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8 .
< e -
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EQUIPM ENT FREQUENCY 1r O.1 1 ! ' ' ' ' ' ' ' ' ' ' I I I i 1 ' ' '
t 2 5 4 5 *7 8 9 go 2 3 4 5 r 7 e s100 2 3 4 5 e 7 e9 1000 FREQUENCY -HZ LEGEND FIG. G-24
. . . TEST RESPONSE SPECTRA MULTIFREQUENCY QUALIFICATION TESTING-
--- REQUIRED RESPONSE -SPECTRA SEISMIC HORIZONTAL REQUIRED RESPONSE - SPECTR A FAULTED (WiTH RH) SHOREHAM NUCLEAR POWER STATION-UNIT l
- REQUIRED RESPONSE - SPECTR A FAULTED (WITH TO) PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 4-FEBRUARY 1991
O O O 10 9 -
8 -
7 -
6 - *
- 3 -
m
. \
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e e
z 1 g I s
. $, : $ [ .
\
o
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FN
, EQUIPMENT FREQUENCY r , , e e ii i i i e i e ii i i i e i i i i i i n,, a 3 4 s s 7 es a 3 4 s s 7 es a 3 4 s s 7es ig ino igon i
. FREQUENCY-HZ LEGEND FIG. G-25 MULTIFREQUENCY QUALIFICATION TESTING-
- *
- TEST RESPONSE SPECTRA
--- REQUIRED RESPONSE -SPECTR A SEISMIC SHOREHAM NUCLEAR POWER STATION-UNIT I REQUIRED RESPONSE-SPECTRA FAULTED (WITH RH) PLANT DESIGN ASSESSMENTFOR SRV AND LOCA LOADS
- REQUIRED RESPONSE -SPECTRA FAULTED (WITH TQ)
REVISION 4-FEBRUARY 1981 t
i APPENDIX H POOL SWEIJ, MODELING t
l l
Revision 4 - February 1981
APPENDIX H POOL SWELL MODELING In order to calculate the suppression pool surface velocity and elevation as a function of time, Stone & Webster has developed a computer code POSH (now incorporated into LOCTVS(13) . This code employs the pool swell model from Reference 2 as further described in Reference 3. Input to the code includes the drywell pressure transient, drywell initial temperature, suppression chamber initial pressure, drywell and suppression chamber free volume, pool surface area, vent area and vent loss coefficient initial vent submergence, and time of vent clearing. Results include pool surface velocity, pool surface elevation, bubble pressure, and suppression chaaber freespace pressure as functions of time.
Figures H-1 through 6 are plots of pool surface velocity and pool surf ace elevation obtained with POSH for the three classes of plants presented in Reference,2 These are provided for comparison with the plots included in Keterence 2 as bencnmark pronlems for the POSH code and show good agreement.
O
, O l H-1 Revision 4 - February 1981 i
REFERENCES
- 1. A Stone & Webster Computer Code to Determine the Pressure and )
Temperature Response of Pressure Suppression Containments to a Icss of Coolant Accident, SNNO2, October 1969 and Supplement 1, August 1973.
- 2. Mark II Containment Dynamic Forcing Functions Information )
Report (DFFR Rev. 2) , NEDO-21061, September 1976. i i
- 3. Analytical Model, Mark II Pool Swell Phenomena, NEDE-21544-P, j l December 1976.
O O
H-2 Revision 4 - February 1981
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- APPENDIX I TEMPERATURE LIMIT FOR STABLE SRV DISCHARGE THROUGH A QUENCHER i
PROPRIETARY - See Proprietary Supplement l to this Report l
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APPENDIX J SNPS SUPPRESSION POOL TEMPERATURE TRANSIENTS ,
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APPENDIX'J SNPS SUPPRESSION POOL TEMPERATURE TRANSIENTS This appendix presents the results of analysis conducted by General Electric Company of SNPS suppression pool temperature response to plant transients involving SRV discharge. These transients have been analyzed in accordance with Reference 1.
Table J-1 presents the input data for SNPS used in the analysis.
~
Charts 1 through 6 and the accompanying figs. J-1 through 6 present a brief description of each case and the corresponding reactor vessel pressure / suppression pool temperature transient.
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J-1 Revision 4 - February 1981
REFERENCES APPENDIX J
- 1. Assumptions for Use in Analyzing Mark II BWR Suppression Pool Temperature Response to Plant Transients Involving Saf ety/ Relief Valve Discharge, Mass / Energy Subconnaittee, Mark I
II Owners Group, March 24, 1980.
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J- 2 Revision 4 - February 1981
TABLE C+1 SHORi, HAM IMPORTANT SYSTEM CHARACTERISTICS -
POOL TEMPERATURE TRANSIENT ANALYSIS Initial Pool Mass 4.53 x 106 lbm Initial Pool Temperature 900F Initial RPV Liquid Mass 465,548 lbm Initial RPV Steam Mass 17,069 lbm RPV S Internals Mass 2.209 x 10* lbm Initial Vessel Preasure 1,020 psia Initial Core Power (105% rated) 2.417 x 106 Btu /sec (2550 MWt)
Initial Steamflow (105% rated) 3,055 lbm/sec Initial CRD Flow 6.39 lbm/sec CRD Flow After Scram (P =0 psig) 21.5 lbm/sec CRD Enthalpy (From CSD) RPV 68 Btu /lbm HPCI On Volume 8,749 ft3 HPCI Off Volume 10,494 ft3 Vessel Max P For Shutdown Cooling 150 psia RHR K in Shutdown Cooling 231.7 Btu /sec 0F RHR in Pool Cooling 231.7 Btu /sec 0F RHR Flowrate in Pool Cooling (2 pumps) 1599 lbm/sec l RHR Pump Horsepower 1000 hp/ pump RHR Flowrate in Shutdown Cooling 1599 lbm/sec Service Water Temp 800F S/RV FLOW (122.5% ASME) P psia FLOW lbm/sec 0 0 1214.7 315.8 ENTHALPY Feedwater MASS lbm (pipe and fluid) 136,631 a02.3 Btu /lbm 193,583 353 Btu /lbm 315,936 287 Btu /lbm i
%/
1 of 1 Revision 4 - February 1981
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i CHART 1 O INPUT ASSUMPTIONS
- 1. Case 1A Stuck Open Relief Valve (SORV) at full power, 1 RHR available
. Manual Scram at T pool = 1100F
. Mechanistic closure of the turbine stop and bypass valves
. One RHR in pool cooling 10 minutes after high temperature alarm
. Main condenser reestablished through bypass system 20 minutes after scram Main condenser available using full bypass capacity until reactor vessel permissive for RHR shutdown cooling
. RHR out of pcc,1 couling when pressure permissive for RHR onutdown cooling is reached. Sixteen N
minute delay for RHR transfer to shutdown cooling.
(Additional SRV's opened as required during
( switchover to assure no repressurization during switchover)
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Revision 4 - February 1981
O O O I,
220 1300 210 - -
1200 200 - -
110 0 19 0 - -
1000
^
i
! '. 180 o
900 Q W m O 17 0 - -
800 b w W
' E E o 16 0 - - .
l
& "I E
i w 15 0 - -
600 [
2 J N 140 -
500 $
. a O >
O l30 - -
400 120 - -
300 4
4
- 11 0 - -
200 100 - -
100
, I ,,iiI , i , I i,iit , t .I , ....I , I . I i . . . . - - . .
0
] 2 4 10 20 40 10 0 200 400 103 2 x103 4 x 103 10* 2 x 10 * ' S a l0*
TIME AFTER SCRAM (SECONDS) i F IG. J- I POOL TEM'ERATURE AND PRESSURE VS. TIME SHOREHAive POOL HEATUP- CASE I A SHOREHAM NUCLEAR POWER STATION-UNIT I PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 4-FEBRUARY 1981 i
CHART 2 INPUT ASSUMPTIONS
- 2. Case IB SORVit) at full power - 2 RHR's available
. Manual Scram at T pool = 1100F Nonmechanistic isolation at scram with a 3.5 second main isolation valve closure
. Two RHR*s in pool cooling 10 minutes after high pool temperature alarm
. When T pool = 1200F, begin manual depressurization by opening additional SRV's so that SORV plus cycled SRV's result in depressurization rate of approximately 1000F/hr unless SORV alone causes faster depressurization
. RHR shutdown cooling not initiated (1)SORV - stuck open relief valve O
O Revision 4 - February 1981
O O O 220 1300 210 - -
1200 4
200 - -
310 0 19 0 - -
3000 i -
I 18 0 - -
900 ^
o -
W m O 17 0 - -
800 b W W E r;:
a 160 - -
700 a
) < $
w i
i m
w 150 - -
600 m a 1 2
- 140 p - -
500 a 0 13 0 - -
400 l
l 12 0 - -
300 l10 - -
200 100 - -
100 i ! ,.it . I i I ,,i.I . l .I .....I . I . I . , ,,,I%, '
O I 2 4 10 20 40 100 200 400 103 2:103 4 :103 th* 2:10' 5 10' 4
TIME AFTER SCRAM (SECONDS)
FIG. J-2
. POOL TEMPERATURE AND PRESSURE VS. TIME i SHOREHAM POOL HEATUP -CASE IB i
SHOREHAM NUCLEAR POWER STATION -UNIT I PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS 1
REVISION 4-FEBRUARY 1981
)
CHART 3 i
INPUT ASSUMPTIONS
- 3. Case 2A Isolation-Scram (nonnechanistic) 1 RHR available
. Isolation Scram at t= 0 nonmechanistic with 3.5 cecond main isolation valve closure One RHR in pool cooling 10 minutes after the event When T pool = 1200F, begin manual depressurization by opening additional valves as needed.
Depressurization at 1000F/hr
. RHR out of pool cooling when pressure permissive for RHR shutdown cooling is reached. Sixteen minute delay for RHR transfer to shutdown cooling.
(Additional SRV's opened as required during switchover to assure no repressurization during switchover l
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Revision 4 - February 1981
O O O 220 1300 210 -
1200
/'% % _
194 *F 190 -
- 1000 n
h 180 -
e 900 ^s W to O 170 -
O -
e00 b W W
$ 160 -
700 ti E 15 0 -
600 3 d s 14 0 -
500 cn 3
W 130 -
400 12 0 -
300 11 0 -
- 200 100 -
10 0 i I ,i... i l i I i...I i I ,I e i e i ,1 i I , I i e it.I . . . O 2 4 20 40 3 3 3 10 100 200 400 10 2 x 10 4x10 10* 2 10 ' 5:10' i
TIME AFTER SCRAM (SECONDS) l l FIG. J-3 t
POOL TEMPERATURE AND PRESSURE VS. TIME SHOREHAM POOL HEATUP- CASE 2A SHOREHAM NUCLEAR POWER STATION -UNIT I i
PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS I
REVISION 4-FEBRUARY 1981
CHART 4 INPUT ASSUMPTIONS
. Isolation Scram at t = 0 nonmechanistic with 3.5 second main isolation valve closure
. SORV at t = 0
. Two RHR's in Pool Cooling at 10 minutes af t.er the j event
. When T pool = 1200F begin manual depressurization by opening additional valves. Depressurize at 1000F/hr
. RHR shutdown cooling not initiated
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b M <-
0 =
s IJ > E Y O 8
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) AISP( ERUSSERP LESSEV m g g g < H c8 o
y o g g g g g g g m mE !
u n o
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S 2N c pf G g
I ' ' ' I I I I I i e I
- a
=
{mZ MmO< m m )k >2 E i_
5-
%(4x Ti o o T m Z c m ,OH 5z 5
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- C F>m
- ~ 4 8 nF S O 37Oy II -
E m C 2
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CHART S INPUT ASSUMPTIONS
- 5. Case 3A Small Break Event Accident Mode 1 RHR available
. Scram at t = 0 on high drysell pressure Isolation at t = 0 (nonmechanistic) with 3.5 second main isolation valve closure
. One RHR in pool cooling 10 minutes after high pool temperature alarm
. When T pool = 1200F, begin manual depressurization by opening additional SRV's as needed.
Depressurize at 1000F/hr
. RER out of pool cooling when pressure permissive for RHR shutdown cooling is reached. Sixteen minute delay for RHR transfer to shutdown cooling.
(SRV's opened as required during switchover to
, assure no repressurization during switchover)
. Automatic RHR switchover to LPCI mode when RPV pressure is less than RHR pump flow head. If switchover occurs after 10 minutes, assume 10 additional minutes to convert manually back to pool cooling. If switchover occurs during the first 10 minutes while operator is attempting to initiate pool cooling, assume no additional time l
lO Revision 4 - February 1981 l , _ . _ , ___ . . _ _ _ - - . _ . - - - - ---- - - - - - -- - -- -
O O O 220 1300 210 - -
1200 200 19 0 y
^w _
110 0 1000
' 18 0 - -
900 a d 4 W m 8 170 - -
600 $
w W E E 16 0 - -
700 g
< m
$ 150 - - 600 n.
2 J N 140 - -
500 $
8 0 0 13 0 - -
400 >
120 - -
300 11 0 -
p 200 10 0 - -
10 0 I ,,ii, i I . I , ..el , I iI i , ,iil i l i l i, ,il . . . O 3 3 3 5 al0*
2 4 10 20 40 100 200 400 10 2 m 10 4 a 10 10
- 2 a 10
- TIME AFTER SCRAM (SECONDS)
FIG. J-5 POOL TEMPERATURE AND PRESSURE VS. TIME SHOREHAM POOL HEATUP-CASE 3A SHOREHAM NUCLEAR POWER STATION-UNIT I PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 4-FEBRUARY 1988
CHART 6 4
INPUT ASSUMPTIONS
- 6. Case 3B Small Break Event 2 RHR's available
. Scram at t = 0 on high drywell pressure
. Isolation at c = o (nonmechanistic) with 3.5 main isolation valve closure
. Two RHRs in pool cooling 10 minutes after high pool temperature alarm
. Automatic RHR switchover to LPCI mode when RPV pressure is less than RHR pump flow head. It switchover occurs after 10 minutes, assume 10 additional minutes to convert manually back to pool cooling. If switchover occurs during first 10 minutes while operator is attempting to initiate pool cooling, assume no additional time
. When T pool = 1200F, begin manual depressurization by opening SRVPs as needed. Depressurize at
, 1000F/hr
. RHR shutdown cooling not initiated l
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Revision 4 - February 1981
O O O i
220 1300 210 - - 1200 200 - -
810 0
--*===~~ ,
19 0 - -
l(v)O
^
i'
' 180 - -
(6 W
900 I m
o
- 170 - -
800 n. -
w W E E g 160 700 g 4 M E W w 150 - -
600 E 2 a
, $ 140 - -
500 $
a 0 13 0 -
4oo 12 0 - -
300 110 -
200 10 0 - _
30 0 i I e i ii. . I , I i.,,il . I iI .,,,il , t . I ,,,,,1 . I ~, .
O 2 4 10 20 40 10 0 200 400 103 2 xlO3 4 x 103 10* 2 x lO* 5 :10*
TIME AFTER SCRAM (SECONDS)
FIG. J-6 POOL TEMPERATURE AND PRESSURE VS. TIME SHOREHAM POOL HEATUP- CASE 3B SHOREHAM NUCLEAR POWER STATION -UNIT I PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 4-FEBRUARY 1981
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APPENDIX X SUBMERGED STRUCTURES l t
i Revision 4 - February 1981
TABLE OF CONTENTS O Section Title Page K
1.0 INTRODUCTION
x-1 K2.0 METHODOLOGY ON SUBMERGED STRUCTURE LOADS K-2 K2.1 LOCA Submerged Structure Loads K-2 K2.1.1 LOCA Water Jet Loads K-2 K2.1.2 LOCA Air Bubble Charging Loads K-2 K2.1.3 Pool Swell Loads K-3 K2.1.4 Pool Fallback Loads K-3 K2.1.5 Condensation Oscillation Loads K-3 K2.1.6 Chugging Loads K-4 K2.2 SRV Submerged Structure Loads K-4 K2.2.1 SRV Water Jet Loads K-5 K2.2.2 SRV Air Bubble Loads K-5 K2.3 Compliance with the NRC Acceptance Criteria on LOCA/SRV Sutznergcd Structure Loads K-6 K2.3.1 LOCA/SRV Jet Loads K-6 K2.3.2 LOCA/SRV Air hubble Drag Loads K-6 X2.3.3 Steam Condensation Drag Loads K-9 n2.4 Seismic Sloshing Submerged Structure Loads K-9 K
3.0 DESCRIPTION
OF SHOREHAM SUBMERGED STRUCTURES AND FORCING FUNCTIONS K-9 K3.1 Types, Dimensions, and Typical Mathematical hacels K-10 of Submerged Structures K3.2 Forcing Functions on Submerged Structures K-10 K
4.0 REFERENCES
K-11 l
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LIST OF TABLES Table Title O
K3.1-1 Suppression Pool Subnerged Structures K3.2-1 Shoreham Flow and Pool Characteristics K3.2-2 Typical Values of KM for Shorehem O
O K-ii Revision 4 - February 1981
1 LIST OF FIGURES
( Fiqure Title K3.1-1 Mathematical Model for Downcomer with Vacuum Breaker K3.1-2 Mathematical Model for Safety / Relief Valve Discharge Line K3.1-3 Mathematical Model for FPCC Piping K3.1-4 Mathematical Model for Core Spray Pump Suction Piping K3.1-5 Mathematical Model for HPCI Turbine Exhaust Piping K3.1-6 Mathematical Model for HPCI Turbine Suction Piping K3.1-7 Mathematical Model for RCIC Turbine Suction Piping K3.1-8 Mathematical Model for RCIC Turbine Exhaust Piping K3.1-9 Mathematical Model for RHR Pump Suction Piping K3.1-10 Mathematical Model for RHR Test Line Discharge Piping K3.1-11 Mathematical Model for RHR Test Line Discharge Piping K3.1-12 Mathematical Model for RCIC Small-Bore Piping K3.1-13 Mathematical Model for Drywell Floor Seal Pressure Monitoring Piping K3.1-14 Mathematical Model for RCIC Small-Dore Piping K3.1-15 Mathematical Model for RHR Small-Bore Piping K3.1-16 Mathematical Model for Level Sensor Piping K3.2-1 Top- View of Major Submerged Structures of Shoreham K3.2-2 Typical LOCA Water Jet Load in X-Direction on a Quencher Arm Segment 4
' K3.2-3 Typical LOCA Water Jet Load in Y-Direction on a Quencher Arm Segment K3.2-4 Typical LOCA Water Jet Load in Z-Direction on a Quencher Arm Segment K3.2-5 Typical LOCA Air Bubble Load on a Downcomer Segment K3.2-6 Typical LOCA Pool Swell Load on a Segment of RHR Test Line Discharge Piping K3.2-7 Typical LOCA Pool Fall Back Load on a Segment of RHR Test Line Discharge Piping K3.2-8 Typical' LOCA Condensation Load on a Segment of RHR Pump Suction Piping K3.2-9 Typical LOCA Chugging Load on a Segment or Downcomer K3.2-10 Typical SRV Air Bubble Load on X-Direction on a Downcomer Segment Due to All Valve Actuation K3.2-11 Typical SRV Air Bubble Load in Z-Direction on a Downcomer Segment Due to All Valve Actuation
() K-iii Revision 4 - February 1981
K
1.0 INTRODUCTION
During a postulated LOCA event or SRV discharge, water clearing, air expulsion and steam condensation in the Mark II suppression pool will create induced fluid motion which in turn may produce loads on the submerged structures. Similar loads are also produced by seismic sloshing. ' In order to evaluate these hydrodynamic loads, analytical models are utilized to predict velocity and acceleration in the entire flow field. The subsequent calculation of standard drag, acceleration drag, and lift force is carried out for each submerged structure as per DFFR Rev. 2(1) and NEDE-21730sa) as follows:
Standard drag = F, = Cri A,.6UlUf (1) 2g, Acceleration drag = Fa = Cm 64N (2) 9e Lift = Fj = Ci Ar 30l0l~ (3) 2ge where:
Cd = Standard Drag Coefficient Ax = Structure's Area Normal to the Flow Direction 6 = Water Density O U = Fluid Velocity Normal to the Structure Axis ge = Acceleration Constant Cm = Inertia Coef ficient V = Structure Volume 0 = Fluid Acceleration Normal to the Structure Axis C, = Lift Coefficient The standard drag acts in phase with the velocity of the flow and the acceleration drag acts in phase with the acceleration of the flow. The lift force exerted on a structure is along the direction perpendicular to the flow velocity. According to the nature of the load direction, the standard drag and acceleration drag are called in-line forces whereas the lift force is called a transverse torce. The total fluid force on a structure is the vector sum of the in-line and transverse forces.
In the process of computing the loads, Appendix D of the NRC Acceptance Criteria (3) concerning the unsteady flow etfect, interterence effeet, equivalent uniform flow, bubole asymmetric effect, etc., are addressed and considered whenever they are required.
Tne approaches to the calculatien or submerged structure loads and the item-by-item compliance with P.he NRC Acceptance Criteria for the Shoreham plant are discussed in Section K2.0. A detailed K-1 Revision 14 - February 1981 l
description of submerged structures and forcing functions is presented in Section K3.0.
K2.0 METHODOLOGY ON SUBMERGED STRUCTURE LOADS ,
The generic load methodology ior LOCA water jet, LOCA air bubble, IDCA pool swell, and LOCA fallback loads is described in DFFR, NEDE-21730, NEDE-21472(*), and NEDE-21471(5). However, no generic load specifications for LOCA steam condensation, SRV water jet, SRV air bubble, and seismic sloshing loads have been cubmitted to the NRC by the Mark II Owners. The latter cases are treated by plant unique analysis. In the following sections, the Shoreham methodology associated with submerged structure loads is pres ented.
K2.1 LOCA subnerged Structure Loads In this section, the methods used to predict the hydrodynamic loads associated with water jet, air bubble charging, pool swell, fallbacx, condensation oscillation (CO) , and chugging in the LOCA event are discussed.
K2.1.1 LOCA Water Jet Loads Shoreham adopts the method described in Reference 6 to meet the intent of the NRC Acceptance Criteria to calculate the LOCA water jet loads. First, the vent clearing transient is calculated as descriced in Section 4.2.1. Based on NEDE-21472(*), DFFR, and NEDE-21730, the jet front location, velocity, and acceleration are calcualted. The potential function from the NRC Acceptance Criteria is used to predict the induced velocity and acceleration in the flow field by the method of NEDE-21471(5). The standard and acceleration drags and lift are then calculated in accordance with Eqs. (1) , (2), and (3). For a structure which is fully engulfed or not fully submerged inside the jet boundary, the procedure outlined in NEDE-21730 is used to calculate the drag forces.
K2.1.2 LOCA Air Bubble Charging Loads Based on NEDE-21471, NEDE-21730, and DFFR, the air bubble charging loads are calculated. The following steps are taken to compute the loads:
l l 1. Use NEDE-21471 to determine the bubble source strength, the induced velocity, and acceleration in the flow field.
- 2. Use procedure in NEDE-21730 and DFFR to determine the drag loads. i K-2 Revision 4 - February 1981
- 3. "RC Acceptance Criteria are addressed according to Attachment 1.K of the Zimmer FSARC*).
K2.1.3 Pool Swell Loads NEDE-21544-Pt7), DFFR, NEDE-21730, and NRC Acceptance Criteria are used to develop the forcing functions on structures in the pool swell zone. Major steps taken are as follows:
- 1. Use NEDE-21544 and DFFR to evaluate the pool water slug velocity, acceleration, and elevation time histories.
- 2. Use the DFFR and NEDE-21730 to calculate the standard drag and acceleration drag loads.
- 3. Use NRC Acceptance Criteria (Section III.B.3.C.1) to calculate the impact loads.
- 4. Use Reference 6 to address the NRC concerns about lift force.
K2.1.4 Pool Fallback Loads The fallback loads on structures are computed based on the velocity and acceleration time histories of a free falling fluid slug. The procedure outlined in DFFR, NEDE-21730, and Reference 6 1s used to evaluate the fallbacx loads on structures located between the vent exit and the maximum pool swell height.
K2.1.5 Condensation Oscillation Loads Condensation oscillation produces an ef f ective unsteady source at the vent exit analogous to the LOCA air bubble source and can also be expected to generate submerged structure loads. Once the source strength is defined, the same basic approach and tundamentals that are applied to the LOCA air bubble load calculation can be utilized to compute loads.
Shoreham uses basically the approaches of General Electric Co.
(GE) dbcuments, DFFR(1), NEDE-21730ca), NEDE-21471(5),
NEDO-21669te), and NEDE-23617-P(*) with the tollowing features:
- 1. Based on analytical hydrodynamic models in NEDO-21669 and NEDE-23617-P(*), the source is detined as follows:
l (a) A point source is located at each vent (downcomer)
- tip.
l (b) The source strength, S, is derived from tull scale single vent data of the 4T test facility and is related to the wall pressure as follows:
K-3 Revision 4 - February 1981 L
S = P wall (4) where:
P wall = 4T C0(10) wall pressure at tank bottom center P= Fluid density t(r) = transfer factor determined by the method of images and 4T geometry
- 2. Aftel the source strength is defined, the worst combination of phasing for all vents is considered to calculate the submerged structure loads in accordance with the analysis and procedure outlined in DFFRC1),
NEDE-21730(2), and NEDE-21471(5).
- 3. Interference effect, unsteady flow effect, and equivalent velocity and acceleration are incorporated to comply with NRC Acceptance Criteria (3).
K2.1.6 Chuqqing Loads Essentially the same methodology used in the load definition for CO is adopted to compute the chugging loads. All documents mentioned'in Section K2.1.5 plus the shoreham response to NRC Question 020.75 form the design bases for load generation. The key ditterence from CO is described as follows:
- 1. Chug source strength is based on the 4T Application Memont) with the maximum source being that resulting in the +20 psi, -14 psi amplitude wall pressure. The frequency range used for chugging loads is 20 through 30 Hz.
- 2. When considering the influence of more than one vent, source amplitude is reduced by using a source amplitude multiplier versus number of vents function based on the asymmetric wall pressure distribution (see Fig. 020.75-1 in Shoreham response to NRC Question 020.75) .
K2.2 SRV Submerged Structure Loads In this section, che methodology adopted to predict SRV submerged structure loads related to the water jet and . ir bubble oscillation is discussed.
K-4 Revision 4 - February 1981
K2.2.1 SRV Water Jet Loads
/O According to the NRC Acceptance Criteria, the SRV water jet loads Q1 may be neglected for those structures located outside of a sphere circumscribed about the quencher arms. Shoreham's zone of influence as defined by the NRC Acceptance Criteria (3) is a sphere with radius of 5 feet. However, as proposed by the Mark II Lead Plant Owners, the zone of influence of SRV water jet is modified from a sphere to a cylinder tangent to the sphere.
Although the jet loads are small, the loads on structures inside the modified cylinder zone of influence are calculated as follows:
- 1. Standard drag is calculated from NEDE-23539(12) and NEDE-25090-Ptaa),
- 2. The induced acceleration is predicted by the line source method (**). The subsequent acceleration drag computation is obtained fro.n DFFR(1) and NEDE-21730(2).
K2.2.2 SRV Air Bubble Loads Shoreham uses a NRC/KWU hybrid method. This methodology includes the following features:
- 1. NRC Acceptance Critera(3)
() (a) Assume air bubble with radius of 5 feet located at quencher center.
(b) Use Mark II submerged structure load calculation methodology in DFFR(1), NEDE-21730(2), and NEDE-21471(5).
(c) Fulfill the NRC required modifications, such as oubble asymmetric affect, interference effect, etc.
- 2. Single exception to the NRC Acceptance Criterla(3): use l KWU specification (25) to define the bubble pressure, 1.e., bubble pressure is equal to 1.5 KKB (3 KWU-PPL l
pressure traces) .
- 3. Application
, (a) Source strength is based on the product of bubble l
radius (5 feet) and pressure (1.5 KKB) .
(b) Frequency range is covered by using various time scale factors of 3 to 10 Hz on KWU pressure time histories. Worst frequency is considered for the structure loads.
K-5 Revision 4 - February 1981 l
(c) Use analysis and procedure in DFFR(1),
NEDE-21730(a), and NEDE-21471(5) to calculate the submerged structu.re loads.
(d) Four cases have been evaluated, namely single valve, all valve, ADS, and asymmetrical SRV actuations.
K2.3 compliance with the NRC Acceptance Criteria on LOCA/SRV Submerged Structure Loads The Shoreham plant has responded to the NRC Acceptance Criteria (3,163 by either adopting the NRC criteria directly or by meeting the intent of the criteria through modified methodology.
The following subsections describe Shoreham's position on the NRC Acceptance Criteria (3 ).
K2.3.1 IDCA/SRV Jet Ioads
- 1. LOCA Water Jet The Shoreham plant has met the intent of the criteria dealing with LOCA jet loads to include the induced acceleration drags and lift (6).
- 2. SRV Water Jet A T-Quencher cylinder zone of influence with a 5 foot radius proposed by the Mark II owners which is accepted by the NRC staff is utilized.
K2.3.2 LOCA/SRV Air Bubble Draq Eoads
- 1. LOCA Air Bubble Loads (a) Bubble asymmetry ef fect:
A conservative estimate of bubble asymmetry effect has been added by increasing acceleration and velocity by 10 percent in accordance with the NRC Acceptance Criteria (3,(16).
(b) Flow unsteady effect:
Consideration of the unsteady effects on drag and lift coefficients is presented in Attachment 1.K of the Zimmer FSAR(6) with the following NRC constraints:
(i) Delete g and use CH =Cg=2 in the FA formula on page 1.K.2-7 of Reference 6.
K-6 Revision 4 - February 1961 O
(ii) For noncylindrical structures the lift coefficient of 1.6 is used.
(iii) The standard drag coefficient tor pool swell and SRV ocillatf.ng bubbles is obtained from data as suggested by the NRC for structures with sha'.p edges.
(c) Equivalent uniform flow velocity and acceleration:
Structures are segmented into small sections such that 1.05 L/D 51.5 (where L is segment length and D is the diameter of the structure) . The velocity and acceleration are then calculated at the geometric center of each segment.
(d) Interference effects:
The detailed methodology presented in Attachment 1.K of the 2,immer FSAR( O is employed to modify the drag and list coefficients due to interf erence ef f ects.
(e) Blockage Eff ect on Downcomer Bracing:
Reference 3 identifies Chapter 8 of Reference 18 as an acceptable method of modifying the vent bracing mystem standard drag coefficient to account tor
' O clockage. Chapter 8 of Reference 18 describes five ooundary constraint effects, two of which are appropriate for vent bracing; solid blockage and wake blockage.
In Reference 18, solid blockage and wake blocxage are both handled the same way by correcting the
" tunnel speed" U T
to the " free-air speed" Up as tollows:
Up=UT (1 + e) where e = cs + ew where the subscript "s" referes to solid blockage and "w", wake blockage.
UT . corresponds to the dynamic head on which the standard drag coefficient is based, while Up refers to the actual velocity at the obstruction in the tunnel. Since C D2 a U2, the correction to the standard drag coefficient is (1 +e)a; that is, the actual C D presented by the obstruction in the tunnel is (1 + c ) a greater than the free-air value.
(1 + e ) a is the drag coefficient multiplier for blockage.
K-7 Revision 4 - February 1981
1 The Shoreham vent bracing system is composed of collars approximately 4 feet in diameter conaected by flat and circular section bracing elements. The total projected area is approximately 30 percent of h the free surface area of the pool. Using Chapter 8 of Reference 17, e for a 2 flat disk and 30 percent blockage is 0.22 and the Cb (corrected) is 3.0. e for a long flat plate with the same blockage is 0.34 and for a long cylinder, 0.16. The w correspoding values for Cp (corrected) are 3.6 and
\ 1.5, respectively.
Prior to publication of Reference 3, Shoreham had used the momentum balance principle to modify standard drag coefficient, C, d to account for the blockage effect. Based on that principle, Cd can be determined by using the following equation:
Cd=kb_ (5)
Ad where Cd = standard drag coefficient k = velocity head loss coefficient A = flow area as if no blockage present Ad = obstruction area Reference 17 is used to obtain k for complex tiow geometries. It was found that the results from this approach are more conservative than those from the NRC suggested method. The overall Cd for the Shoreham bracing system using the momentum balance principle and velocity head loss coefficient data from Reference 16 is 5.27 compared to 3.0 using the methods of Chapter 8 of Reference 18.
The NRC criteria is acceptable and is used to prcvide additional justification for the more conservative vent bracing drag coefficients used in the Shoreham assessment.
(t) Replacing Mg bypFB V: A The criteria is acceptable and is included in the load calculation.
- 2. SRV-Ramshead Air Bubble Loads (a) Neglecting the standard drag:
The standard drag is calculated and included for all submerged structure loads.
K-8 Revision 4 - February 1981
(b) Other constraints and modifications:
,O The constraints and modifications described above V for LOCA air bubble loads are considered and addressed.
- 3. Quencher Air Bubble Loads Shoreham uses T-Quencher device for steam discharge.
The associated air bubble loads are calculated on the basis of Section K2.2.2 which identifies a single exception to the NRC Acceptance Criteria (3), i.e., the bubble pressure of the quencher is assumed to be 1.5 times the wall pressures specified by the KWU specification ( W .
K2.3.3 Steaa Condensation Draq Loads No generic load methodology has been provided by the Mark II Owners. Shoreham uses the approaches described in Sections K2.1.4 and K2.1.5 for CO and chugging submerged structure loads, respectively. Drag and. lift coefficients due to unsteady flow and interference effects are calculated in accordance with Attachment 1.K of the Zingner FSAR (6 3 Also, the equivalent uniform flow velocity and acceleration on a structure are properly determined by segmenting the structure into small sections.
K2.4 Seismic Sloshing Submerged Structure Loads Seismically induced vibration of the wetwell will generate pool sloshing suomerged structure loads. Thes e are considered as secondary loads. The methodology for establishing the induced flow field resulting from seismic sicshing is described in Section 2.1.3 of Chapter 2 and Reference 19. After the velocity and acceleration time histories are determined, the drag and lift loads are calculated by using Eqs. (1) , (2) , and (3). The loads have been generated for both the operating basis earthquake (OBE) and the safe shutdown earthquake (SSE). It was found as expected that these loads are not significant.
K
3.0 DESCRIPTION
OF SHOREHAM SUBMERGED STRUCUTRES AND FORCING FUNCTIONS Sutmerged structures in the Shoreham suppression pool are oefined
- as those subject to the submerged structure loads as described in Section K2.0. These structures are either normally submerged within the pool or located in the pool swell zone (from el 18-0 to el 46-75) above the pool surface. In this section, the characteristics of these structures, associated flows, and induced forcing functions are delineated.
V K-9 Revision 4 - February 1981
K3.1 Types, Dimensions, and Typical Mathematical Models of Submerged Structures The submerged structures in suppression pool consist or the following types:
- 1. Drywell floor column,
- 2. downcomer,
- 3. safety / relief valve discharge line (SRV DL) ,
- 4. fuel pool cooling and clean up piping (FPCC),
- 5. core spray suction piping,
- b. high pressure coolant injection (HPCI) turbine exhaust piping,
- 7. HPCI turbine suction piping,
- 8. reactor core isolation cooling system (RCIC) turbine suction piping,
- 9. RCIC turbine exhaust piping,
- 10. residual heat removal system (Ium) test line discharge,
- 11. RHR system relief valve discharge,
- 12. RHR pump suction,
- 13. temperature elements,
- 14. other small bore piping, 15 submerged pipe supports, 16, downcomr bracing, and
- 11. miscellaneous bracing members and beams.
The number, size, and location of the above-mentioned submerged ctructures are listed in Table K3.1.1. Typical mathematical models of submerged structures used for dynamic stress analysis are shown on Figs. K3.1.1 through 16. In these figures, the coordinate system is defined as follows:
-z: in the direction of " CALLED NORTH" per piping drawing
+y: vertically up
+x: according to right hand rule 1
i K3.2 Forcing Functions on Submerged Structures The submerged structure loads are calculated in accordance with l Eqs. (1) , (2) , and (3). The drag and lift coefficients are i properly modified to account for unsteady flow and interference eftects. The velocity and acceleration at segment centers are evaluated by using adequate nodalization of structures and multiplying by necessary multipliers to account for the bubble asymmetry effects. Table K3.2.1 lists the key parameters tor load calculation, while Table K3.2.2 tabulates the multipliers l
ror acceleration drag coefficients due to interference effects.
Figure K3.2.1 shows the top view of major submerged structures for Shoreham. It d epicts the relative distance between major K-10 Revision 4 - February 1981 l
l l
submerged structures for the consideration of the interference effect.
Typical t:une histories of submerged structure loads are shown on Figs. K3.2.2 through 11. In these figures, the coordinate system is defined as follows:
+x: radially outward from reactor pressure vessel center line
+y: vertically up
+z: by right hand rule origin: at the center of structure segment K4.0 REFERENCE
- 1. " Mark II Containment Dynamic Forcing Functions Information Report (DFFR) ," NEDO-21061-P, NEDO-21061, Revision 2, Class 1, June 1978, (GE Report) .
- 2. " Mark II Pressure Suppression Containment System Loads on Submerged Structures -
An Application Memorandum,"
NEDE-21730, December 1977. (GE Report) .
- 3. " Mark 11 Containment Lead Plant Program Load Evaluation and Acceptance Criteria," USNRC, NUREG-0487, October 1978.
- 4. " Analytical Model for Liquid Jet Properties for Predicting Forces on Rigid Submerged Structures," NEDE-21472, September 1977. (GE Report) .
- 5. " Analytical Model for Estimating Drag Forces on Rigid Sutunerged Structures Caused by LOCA and Safety Valve Ramshead Air Discharge," GE Report, NEDE-21471, September 1977.
- 6. Zimmer Nuclear Power Station -
Unit 1 - Attachment 1K -
Amendment 99 -
Submittal of Revision 61 to the FSAR, September 28, 1979.
- 7. " Mark II Pressure Suppression Containment System: An Analytical Model of the Pool Swell Phenomenon," GE Report NEDE-215 4 4-P , December 1976.
- 8. "The Multivent Hydrodynamic Model for Calculating Pool Boundary Loads Due to Chugging - Mark II Containments," GE Report NEDE-21669, June 1977.
- 9. " Mark II Lead Plant '1bpical Report Pool boundary and Main Vent Chugging Loads Justification," GE Report NhDE-23617-P, July 1977.
O V K-11 Revision 4 - February 1981
- 10. 4T Condensation Oscillation (4TCO) Test Program Final Test Report, General Electric Co. Report NEDE-24811-P, May 1980.
- 11. " Mark II Phase I, II, and III Temporary Tall Tank Test Application Memorandum," January 1977, Letter from L.J. Sobon (G .E . ) to O. Parr (NRC).
- 12. " Analytical Model for Quencher Water Jet Loads on Rigid Submerged Structures," GE Report NEDE-23539-P, Class III, FxT: t ??"'9 (Oratt).
- 13. " Analytical Model for T-Quencher Water Jet loads on Submerged Structures," GE Report NEDE-25090-P, Class III, May 1979.
- 14. Shames, I.H., " Mechanics of Fluids," McGraw-Hill Book Company, 1962.
- 15. "Thermo-Hydraulic Quencher Design of the Safety Relief System," Kraftwork Union Report R14-25/1978, Revision 1, April 1978.
- 16. " Supplement No. 1 to the Mark II Containment Lead Plant Program Load Evaluation and Acceptance Criteria," USNRC, NUREG-0487, Supplement 1, May 1980.
- 17. I.E. Idel*chik, " Handbook of Hydraulic Resistance -
Coefficients of Local Resistance and of Friction," 1966, l AEC-TR-6630.
- 18. Pankhurst, R.C. and Holder, D.W., " Wind Tunnel Technique,"
Chapter 8, Pitman 6 Sons, Ltd., London, 1952.
- 19. T.F.Li, C.C. Line and C.H. Luk, "Three-Directional Fluid Pool Seismic Sloshing Analysis," 80-c2/PVP-49, Presented at the ASME Century 2 Emerging Technology Conferences, August 1980.
l l
i K-12 Revision 4 - February 1981
. . . . _ _ . . . _ . _ . . _ - - .. ._ . _ _ m i \
i 4
TABLE K3.1-1 l
SUPPRESSION POOL SUBMERGED STRUC'NRES Size Elevation j Structures Number finchl Frosi 3
- 1. Column 14 36 59-3.5 9.0
- 2. Downcomer 88 24 59-3.5 9.0
- 3. SRV Discharge Line 11 10,12,15 62-8 0-9 i 4
- 4. FPCC 1 10 29-0 8-9
- 5. Core Spray Suction 2 14 24-0 19-0
- 6. HPCI Turbine Exhaust 1 18 30-0 10-0
- 7. HPCI Turbine Suction 1 16 24-0 18-2
- 8. RCIC Turbine Suction 1 6 24-0 21-4
- 9. RCIC Turbine Exhaust 1 8 31-0 17-0.5
- 10 . RHR Test Line Discharge 2 16 29-0 14-6,20-O
- 13. Tesaperature Elements 26 1/2,2 31-0 24-6 5
4 14 Other Small Bore Piping 11 1/2,3/4,1,11/2,2 59-0 19-0 j- 15. Subanerged Pipe Supports 2 11/2,21/2 25-3 22-0 16 . Downcesaer Bracing -
Various 27-9 27-9
- 17. Miscellaneous Bracing Members -
Various 46-9 32-0 arut Beanas i
i 1 of 1 Revision 4 - February 1981
I TABLE K3.2-1 SHOREHAM FLOW AND POOL CHARACTERISTICS Flow Characteristics l Phenomena Flow Type LOCH Air Bubble Charging Constant Acceleration LOCA Pool Swell Oscillatory Flow LOCA Fallback Constant Acceleration LOCA Condensation Oscillation Oscillatory Flow LOCA Chugging Oscillatory Flow SRV Air Bubble Oscillation Decaying Oscillation Pool Characteristics Basement at Elevation 9-J High Water Level at Elevation 27-0 Low Water Level at Elevation 26-0 Dimension of Annular Suppression Pool Tank:
Inside radius = 13-0 Outside radius = 39.5-0 r
l l
l 1
i l
lO 1 of 1 Revision 4 - February 1981 l
l
TABLE K3.2-2 TYPICAL VALUE OF K, FOR SHOREHAM Target EM Target Target EM gg DP S 1.0394 DP 33 1.0266 DP 61 1.255 DP 6 1.0394 DP 34 1.0231 DP 62 1.255 DP 7 1.0064 DP 35 1.033 DP 63 1.2555 DP 8 1.0455 DP 36 1.0334 DP 64 1.2552 DP 9 1.0 DP 37 1.035 DP 65 1.2552 DP 10 1.0394 DP 38 1.083 DP 66 1.2552 DP 11 1.0394 DP 39 1.037 DP 67 1.2552 DP 12 1.0064 DP 40 1.037 DP 68 1.2552 DP 13 1.0455 DP 41 1.0463 DP 69 1.2555 DP 14 1.0 DD 42 1.0438 DP 70 1.2552 DP 15 1.0389 DP 43 1.0754 DP 71 1.2552 DP 16 1.013 DP 44 1.038 DP 72 1.2552 DP 17 1.0389 DP 45 1.075 DP 73 1.2552 DP 18 1.013 DP 46 1.0476 DP 74 1.2552 DP 19 1.0231 DP 47 1.0297 DP 75 1.2552 DP 20 1.0476 DP 48 1.0335 DP 76 1.2549 DP 21 1.0127 DP 49 1.0231 DP 77 1.2558 DP 22 1.051 DP 50 1.023 DP 78 1.2552 DP 23 1.031 DP 51 1.035 DP 79 1.2552 DP 24 1.031 DP 52 1.083 DP 80 1.2555 DP 25 1.081 DP 53 1.037 O DP 26 DP 27 1.0162 1.051 DP 54 DP 55 1.037 1.079 DP 81 DP 82 DP 83 1.2552 1.2552 1.2552 DP 28 1.1095 DP 56 1.072 DP 84 1.2552 DP 29 1.0476 DP 57 1.075 DP 85 1.2552 l
DP 30 1.031 DP 58 1.038 DP 86 1.2552 DP 31 1.031 DP 59 1.075 DP 87 1.2552 DP 32 1.081 DP 60 1.047 DP 88 1.2552 Col 1 1.12772 SRV A 1.18717 Col 2 1.12776 SRV B 1.20335 Col 3 1.13488 SRV C 1.20371 Col 4 1.1349 SRV D 1.20359 Col 5 1.14183 SRV E 1.1872 l Col 6 1.14183 SRV F 1.20334 Col 7 1.13469 SRV G 1.18717 Col 8 1.12772 SRV H 1.20358 Col 9 1.12796 SRV J 1.2175 Col 10 1.13488 SRV K 1.20371 Col 11 1.13513 SRV L 1.18715 Col 12 1.14162 Col 13 1.14183 Col 14 1.13489 1 of 2 "evision 4 - February 1981 er - - ,--
l l
TABLE K3.2-2 (CONT *D)
NOTES:
DP = Downcomer Pipe, Col = Colunal SRV = SRV Line K = C'M CM where:
C'M= acceleration drag (inertia drag) coetficient with interference effects Cg = acceleration drag coefficient without interference effects i
O l
O 2 of 2 Revision 4 - February 1981
is'
[o* , .i, 3 ga DRYWELL FLOOR EL.59.292' ANCHOR (- 21.HX,59.292Y,16.830 kD4 2.13' OD 24.0"
+ Fy (Da THs.375"
+0y 2'.13'
- I* VACUUM BREAKER
+F l.O ,
+4a I ",,
(( ,,4
+Fr Z
.; 21 is
+At ( g ,, to 2.8' (b23 2.8' (D24
- 2. 8*
(Das 2 . 8' PLATFORM E L . 38*-9"
--4Das 2.01' 1 (D27 V 2.01' (has 2.Ol'
' (Das DOWNCOMER BRACING AT EL 37*-9" 2.Ol' 30 8 I Mess / _D 7 WATER SURFACE EL. 27.O' m- - = - _
,, , ,ss
- 1. 8, i . ----()42 I
I . 8'
_ __ qh**
l 1.8 g __ _ . _4 5s 4
t8 .
I I,. _j )sa 09
-4 ) ss FIG. K3.1.1 MATHEMATICAL MODEL FOR DOWNCOMER WITH A VACUUM BREAKER SHOREHAM NUCLEAR POWER STATION - UNIT I PLANT DESIGN ASSESSMENT FOR SRV AND LOCA LOADS REVISION 4-FEBRUARY 1981
l in-h' v *(* [ o' M,;'31 DRYWELL CONT.ON
+FY .267 FLOOR M Se#K -1245
+Ay s73 g EL. 62 -8 82.45'
+My Q 167' w I Mx +Fx 2 57* ' sao I s78 y 267.' SEE DETAIL"A" 4M1 VA PSA-042 I
+3* DETAll "B"
+M2 h src .5 IO"-SLP 206-3OI-5 435' df 2 sot PSR 153
+F2
- E H1.00PSR153 L.C.
+A2 N.' 12" X 10" SEE DETAIL"8" DETAIL "A" 24o
?.' p *
,, REDUCING 23s ELBOW '
SEE DETAIL"E" asoi, 23o AX AZ 2" -
6 AX: .060' PT FM PT TO 15.68' R AD.-- 22T ; i6
.368' 230 227 . 0 78' l.75 8END ,2jo 23
.060' AY 227 224 .056' l.33 als ; ,442' AZs .516 224 220 .17' l.32 210 1.25, d: .,g 4 7 '
- 714' A X: .147' 220 217 .28' l.29' D UM L " E "
A Y = . 4 4 2, 217 213 .39' l.23' .266' AZ= 1.25 213 210 .502' l.24' -
PSR 119 CONT. FROM 200 *>? v.C.
1.45, B E L,0W LEGEND A s MASS POINT 89 5 39 g 6 ,
- =CO RAINT g,45' .42 ;$aX12" pg ,7A ANSITION
.33 Q 50' ies - 6.28* n, , 99 187' 872
+2 'o 1.18' l.02'
- )" 17 s D ETA ll,, C,, .102 1.18 ss [2.18 l.28 PSR 120 g, g 7' 1.278' loo lis' l.792' 7' 1.875' ,V .(PT. lO)
j [.320 is n 5. I' SEE DETAIL"C" 107 2026 5.O' l TEE QUENCHER SN-OO6-QEN OOlO I829' \ 14 0.D. p(
as a5.O', 'PSR 120 3.75 / I , i21 y ,Q L.C., L.C $
3 SEE DETAIL"D" .O 3
.3715' 2.O' isz "5' '4.O' EL 2 8' O" 1.0,
" 2 0 "O.D.+
SC H. 80 ,7 0 ' Iss E. .708 Sodh
" 172 o
- 2. O')
- 4. 8 ' E L . 9'-il l{s j;3o, sys 1,4c { i,42 s5d' 5.25, N.S. ,43 [ ' N2.0NIS. ~
, .708 iso
)
70d6 kqE 3,60 15 2 2.0,
.3 3' 66.90 E.W.
~ N .708, CONT.
D ETA I L "D" ABOVE PSR121 FIG. K 3.1. 2 MATHEMATICAL MODEL FOR SAFETY /
p'"j s
RELIEF VALVE DISCHARGE LINE SHOhEHAM NUCLEAR POWER STATION - UNIT I PLANT DESIGN ASSESSMENT FOR SRVAND LOCOLOADS REVISION 4-FEBRUARY 1981
+ Fy
+ 6y ,3
+My b +M a e, PENERATION X5-6
+Fz ,
+0Z
.058' .089, i .3697 f[.1841'
.1260 '1423'f.2856 .2156' 1 # .5103' / ,,Ms, -. 6645' 10175'
- j A CHOR
.0616
' .0 31
.2333
.010 0 'O s S6W 42 275
.1188'
.1237' FOR CONT.
124f SEE PR08. 74 3633' l' AX-7E-l 0521'
.3677' AL
~7II3 isa g EL.27'-O" V HIGH WATER LEVEL
.6085' -
led L I.217' IG 41-204 b(TYP) t o". FC- 124 151-2 22 a g 24j g (j LEG EN D:
A = MASS PolNT re; 6
- ^ '
IG41-soJg PSR 089 NS-EW stag 3,
1.217' W/N 6 8LIND FLANGE
.6555', ,EL 16'-0" osf*',
.5615' ',*, ,
.3744' 3 31 3
382' 5l, L ANSITEE 56 EL 8.75' o l
l FIG. K 3.1. 3 MATHEMATICAL MODEL FOR FPCC PIPING (O_) SHOREHAM NUCLEAR POWER STATION -UNIT I PLANT DESIGN ASSESSMENT FOR SRV AND LOCA 1. DADS REVISION 4-FEBRUARY 1981
+ Fy O +0F
+Myb +u x
- g. Cb'
+An ggo *
+M g,0 go
+F: Z Eb 7 5 TON
+02 A1 17 2,
.427, gb ggiB Agg
. 682' ANCHOR
.275 4
.176'
.071' . FOR CONT.
SEE PROB 110 113, 4 7 7* AX-II A 88 0
' iss' IE2l-
.296*
.375 t4"-WR-21 151 2 18 5' *
.348
.075'
.435'
.032'
.013'
.219' CLl51 #
,7474 g ,, W/N FLANGES CL153 '4 ;20
<> is SUCTION STR AINER
.330' ,
IE21 FL-082 t> l9
.670' sgzo LEGEND:
= MASS POINT
.706 e : CONSTR AINT POINT r,
.202' JL22
.556 EL IO'- O" j 24 FIG. K 3.1. 4 MATHEMATICAL MODEL FOR fl k/
CORE SPRAY PUMP SUCTION PIPING SHOREHAM NUCLEAR POWER STATION -UNIT I PLANT DESIGN ASSESSMENTFOR SRV AND LOCA LOADS REVISION 4-FEBRUARY 1981
+v-+ w y m* ~ e w w - "-~
ve " -
m'- 4-g e- s 4 w-
+ Fy g .
/~N +My 6 +Ma ANCHOR CONT. ON Y[g", $o* ao X i8
.+. yg as .1861
+Fz 40
.0229'
+Az IEji- 357' .6410 18 -SLP 7-151-2 s
.3504' So .043'
.0457
.3722, 7e .043'
.0979' 797
.3504 h v
iso s so, X s . 0 50 4,' ,
Y s.1712 .130, f Z a ,410 4' 185
{ W/N FLANGE R/F# 150 esg_ ) X s.0 4 3 9.'
y,,; 4 9g '
iso ag.1352 Zs 3573 .2314' s/ Q:g;f .Z .0622
?'.25, ZsN 7 [ o'S$'
.8281' X 3.003' Y .3607' - -
X 3. OO71, ano s, Z s.0244' r .4750'
, lX s.0531,, Y z .0614, trosg 1 Ys.3029 Z s.0579
,Z s. 2 8 5 8'
' .8702
,3q X s.Ol95 8 .5208' 1os3, Y s.3 8 61' 2*0d' s Z s.15 81' .8334, 2so n
{sb) tioj'
.1458' d'
42 o E L. 2 7 '- O" Q HIGH WATER LEVEL .4208' d
- == =e:g_c=, _.=c- _.
ass a 18'- S PAR GER
,3p d' (S P-OOI) 27on .3083
'*U d' l.0 917' IE51 teo n d ' PSRO55 .3667' aso ,
E L. 2 4'-O" E W. NS
.1667' 330 h
n .0833, 31o j, U0Jb .0633 EL. f 6 '- O, 320j ,
tooa g g LEGEND As M ASS POINT e CONSTRAINT POINT FIG. K 3.1.5 MATHEMATICAL MODEL FOR HPCI TURBlNE EXHAUST PIPING f) v SHOREHAM NUCLEAR POWER STATION -UNIT I PLANT DESIGN ASSESSMENT FOR SRV AND LOCO LOADS REVISION 4 FEBRUARY 1981
+ Fy
+0V g. N P +M y b +M A s MASS POINT , syn 9*
0
- : CONSTRAINT POINT p/ +M,
+4r 4' O" EL.2 5
N (X TIO AZ 2101 ) '
W
/p,9; S6 ENETRA
/g+N. ~lO/.g? ~ 5__P xf
.= 8 p
- g ; ((
p .py,,
% '#9?g, ,
o
O24 7 676 2.00' Ss - L.R. ELBOW .1522 a 6 ig y L.22'-O ,
'4"7' IE41 -lC179-BJ-Of t ; L[ 150 # W/N FLANGE ').2815,
.4167,
() - .2813 I,3333
- 3675' sL2 9 r ' .15 22
.9167 3675' sL22 SUCTION FILTER .1522 1.1666' ,,, ,
L JL24 EL.18'- f has NOTE SUPP RESSION POOL j HIGH WATER LEVEL 27'-O" l
l l
l FIG. K 3.1.6 MATHEMATICAL MODEL FOR HPCI
^ TURBINE SUCTION PIPING SHOREH AM NUCLEAR POWER STATION - UNIT I PLANT DESIGN ASSESSMENT FOR SRV AND LOCO LOADS REVISION 4-FEBRUARY 1981
+ Fy
+ Ay ch" V
+u V S +Mx Y x l0
+,, +=>
+0z CONT ON PROB . 022
/ AX-2C-l
/
/
[ XY=: .0516'
.0571, ,
N
.12 2' kt 2
Z : .127 7 ' .244' I 5 \
.6 '
.244' PEN ETR ATION Z .2255,'J , b D is \ X-19 g .244 -
20
, .14 0' N as
\
X =.0609',) t \
Y : .2 4 3 3, sc g Z : .15 07 , g 35 I E \
X s.0214 6'- WR 2-151-2 PRIMARY WALL Y a.1378' 40 Z z.0529',
/\ <l64s O
tbs 0 4408' ss A 150 PSI WELDING NECK FLANGE H> 60
.358' JL 6s LEGEND:
, ASTRAIN ER A = M ASS POINT e : CONSTRAINT POINT JL 70 NOTE SUPPRESSION POOL
.325' HIGH WATER LEVEL 27*-O" 75 l FI G. K 3.1.7 MATHEMATICAL MODEL FOR HPCI O TURBINE SUCTION PIPING C/ SHOREHAM NUCLEAR POWER STATION - UNIT I PLANT DESIGN ASSESSMENTFOR SRV AND LOCO LOADS REVISION 4-FEBRUARY 1981
+ Fy g /
+ Ly i5- , p
+1 ,S' O f b 5 g a
+r,+"*
s to '
/
is ' CONT.ON
[XY s .0.123' Y . 759
/ PROB.023
/ [20 76' .-
X: . 3 5' ,Z: .13 7' .15 167' Y .325, 25 ~
.18 6,- PRIMARY Z .389' 3o .167' WALL 55 .167 '
40304' .337'
- 45 X .256' S c, X : . 35, Y a .184' 550 , Y a . 3 25, Z : .2 84'
.698' [Z:.384, 50Jk ^
-lE-51 1.0, 8 "- S L P 151 - 2
- 1. O' 7eag EL . 27'- O" Q HIGH WATER LEVEL 75JL t .864' esag 864'
'0JL
. 8 4' EL. 24.0" IE-51
.893' PSR 039
'00Jk N.S.E.W
~I04 LEGEND:
A = M ASS POINT b'
815 4D e : CONSTRAINT POINT 092' 120JL
.857' i25;L
.864' SPARGER PIPE isoag
.864 1354L
.864' FIG. K 3.1.8 4o26 M ATHEMATICAL MODEL FOR RCIC TURBINE EXHAUST PIPING
,r'l,
\
.864' SHOREHAM NUCLEAR POWER STATION - UNIT I
' ' ' ' PLANT DESIGN ASSESSMENTFOR SRVANDLOCO LOADS
.432' iso J k REVISION 4-FEBRUARY 1981
y is*
N V Y b +M a [o "O
+bz 20"-W R -205 -151- 2 X= 488' Y .4 893' s ,Z3.0382' 4 .026'
% 329' ' s Ng.9A 42
.66 8 .0523' \ N
,, pgggTRgTtO gt..zA'-A 2 se,4* $ S.R. EL80W s6* M ioj l.16 6'
, 'dk CONTAIN MENT WALL '
'474' 20"- 150# W/N - s 12 FLANGE TYPE-R.F " 333,
. l
,3
.7455' n
LEGEND:
A MASS POINT
' [ '
e : CONSTR AINT POINT
'8n NOTE SUPPRESSION POOt. EL.18t O/4" ta gg HIGH WATER LEVEL:27'-O" FIG. K 3.l.9 MATHEMATICAL MODEL FOR RHR (3
w)
PUMP SUCTION PIPING SHOREHAM NUCLEAR POWER STATION -UNIT I '
PLANT DESIGN ASSESSMENTFOR SRVAND LOCOLOADS REVISION 4-FE8RUARY 1981
+ Fy -
+ Ay ,3 (O)
+MY b +M,
+, l$o *
+p, +Mr ,
CO N T. O N
+A-* PRIMARY Y .2629 ,
AX-8F CONTAINM ENT PR OB- 805] Z .0347') X = .4 62 8'
' v. > vl,- X : . 5 34 3', (Y.2679, Z .0347' Z.0402 s ' '-
- l . ' v. b.?'. .
% X .2672'
s'% . . Y 3.46 41' IE Il 16"WR 222151-2 /
[. s
.498' 3
j y.,. v.,. .0375 ,
PENETRATION ' V: p /, i .* t .'
X IO A ** '
3 X :.2672, e
p ', , ,e y a,4 641,
- p. .. \
'F Z .0 2 Ol' , sjg EL. 27.OO' V HIGH WATER LEVEL
_ - = -- --
X . 00 0'. '659' 8, Ys.5359 e 2 =.000* *0eg 47g iso,g .,""--
.659' 400gg
.659'
.659
{p}
v
'00 dL llojg
.659, "
.659' IEll isodL PSR 329 '85JL
/
.659 / EW,NS ,
865
[19.00' '
120jg
.659' .659' i2Sj 170jg g
.659' 659' 175aL 130 -dL-FLANGE TO .420'
.659, ACCEPT '3'd L STRAINER '400 .239' 18 0;g 14 5 -dW.181'
.659' ~w issag h
.659' 19 0 ;L
! .659' FI G. K 3.1.10 MATHEMATICAL MODEL FOR RHR
'"ak TEST LiNE DISCHARGE PIPING l (~}
(/ .659' SHOREHAM NUCLEAR POWER STATION -UNIT I 200 EL.14.50' PLANT DESIGN ASSESSMENT FOR SRV AND LOCO LOADS 3
REVISION 4-FEBRUARY l981
+ Fy
+ 4y . CT bi .g y VS +M V
Y"x $o *
+F,l+ ~ p
+02 W CONT. PROB.
\o o # 809. AX-8K 9
IE ll y c, 16"W R- 308 151-2 (TYP
.747 FOR 4) [$- ANCHOR 3
.060' s
PRIMARY 7
CONTAIN MENT io WALL h'* IE X = .5 34' Z s .043, los a
16'-WR-3lO 151-2 ,
X z.464' illGH WATER LEVEL 7 dL tos E L . 27 ' O Y:.268
---l=5-- 7.3.037'] go
.613 ,,,
no ,
X =.267,
.613' Yz.464 ios 30,30 Z z.0 21,'
,,3
.645' O*
Z:.536 A r s n'10 nos k.) .645' its DETAIL OF ELBOW (10-106) h
.645' s n'30 \45 Y .828'
.645' 4 jg l3s iso e iss X =,29 3,
.775 al40 X=.414., Y : .586
1 2 .433 EL. 2 2.O' Z :.717, ,
43A t .420'gFLANGE ACCEPT TO
"' 5' STRAINER j
DETAll OF ELBOW (155-165) issv__
is x 4 /s, LEGEND:
A = MASS POINT 4 = CONSTRAINT POINT e
FIG. K 3.1. Il MATHEMATICAL MODEL FOR RHR p)
(, '
TEST LINE DISCHARGE PIPING SHOREHAM NUCLEAR POWER STATION -UNIT I PLANT DESIGN ASSESSMENT FOR SRV AND LOCO LOADS REVISION 4-FEBRUARY 1981
+ Fy g A + Ay . # > .
+My6 +u, d'1*t 4+rx 8 g#oc* ) s#
+M Z 6
/
1f0 ,
FIELO TO ANALYSE
+Fz
+g, 15p' , CONT PROB.
e' zoy '
e PRIMARY as, / ,
WALL
/ ' O/*
l 0 ' f
/ ss 382' 7 2"- A-lO 151-2
(,EL.29 5"0 f 50 , t EL.27'-il'/z" d L55 90
. L95
.5 21' es
.467' M
.7 5'
,EL.27 10"' 7HIGH WATER LEVEL
- ~-
601 > 100' - -
so O. N' .50,,,,__
.312' ios;g
.5 21,
- )g .50' hes 75
.312 iloti izoit . 5 0' s' 467'
' ' 'd' 1
it.
E L.24'-II' /
izs/
N(. 5, 50'
'30
.50, ANCHOR LJ t35;g I404 ,. 50'
. 50' I 4 Sag
. 5 0' t50< i
. 5 0' LEGEND:
155 d' A: M ASS POINT
. 50'
- = CONSTRAINT POINT
'80 iD
.208' k
EL .21'-6"
,,3 8 g M. lESI-PSR
- N. S. 059 ttoa ,
.50' 17 5 J L l .50' l iso < >
is s, ,' 5 0' FIG. K 3. l .12
'50' MATHEMATICAL MODEL FOR RCIC (7 ISO <> SMALL- BORE PIPING I (l .. .2 5' SHOREHAM NUCLEAR POWER STATION -UNIT I l (E L .18,-1 n PLANT DESIGN ASSESSMENT FOR SRV AND LOCO LOADS )
REVISION 4-FEBRUARY 1981
) f C[\ U v h 3 - 90* LR. ELBOW
&o f ,
ANCHOR 3 @ ,,
DRYWELL 20 @
FLOOR 4 8 90 'J
- [io3o SEAL OUTER 4" 3 ,
l2 ,7lo<
JL350 aL4so
,333.s gsro ][e4o STRUCTURE ,o jg ,333 ,333 .333
,733 j- 3 y 4 h (TYP) 3so (TYP) I JL5to (T Y P) Jgs8o aisso (TYP)d d
L L*'O (TYP)J Llo4o JL'oSo f 733' f i dL'7 " ero jgioso 8 TH REDUC.
750' aszo n700 d L CAP s 6 33T N' LOW h@8/2"VOS- 2"
/ so I 50- '
a L3 eo 250' ' = 0
.333
[ .
60Y-2 ?*
-p .13 5' 153-2 '710' 33o . E L 50'-7"
.083*n M' dL720 1723-
^*
ism
'25 ~ a dk' '
'o' O EL.27'O" V LEVEL 22
~
ri[T2( .3 3'
),
I2O,459' no EL 58'-O" .639' I# T "'
( );isso PSR 050
'dL74o EW-NS 505I d LO (fff)eJk"lo
- o seo iro iso iso l40 ' iT23 n47o gro
- 'dk27o 4,o PSR .333' 73o IT23* nn2o aLsso
,4 59, E L 57'-i" 5049 251 EL 37'-O"
.083' 2eo- -
EW-NS 7so: ( 22 L', o EW-NS JLuso
.135 .625' JL42o sL59o .0 8 2' ag77o
,333 j g g4o 299 _ ,$o
-;17 2, ags m .252 seo h,EL 31'-O"
.3, .,333 yp 7so JLuso l',-VOS-60Y-2 # _l72, sio d L*'O l n7so .082';gsro o jg uso g
dL Vo .333 J', e o .333 LEGEND: .7 o'
".jg,3o dL* ' di soo (ryP3 .085' d L"7g g.,3 A z M ASS POINT (TY P3 L330 .167', ,,so e ; geso jge.o JL" IT23-8 3 +333, JL'Mo PSR 5053 e Cr JSTRAINT POINT g ,3,33 , g4 yo .17 8 ;
gs4o dkero '333 EW-NS sso gl2oo n340 d Li n,o .333,J
.f L4ao 155 EL 40*-7" ,
j'.3o
.710
- 33 3.a suo giozo .167'JL R M'- 5"
.3 33,j g ,7o .E.jk,4g m dL3So jg4so ss J 333.sALiO3O LJ s,
+ Fy
+ Ay ,3 FIG. K 3.1.13
+My 6 +Mx MATHEMATICAL MODEL FOR DRYWELL pg=" &l.o*'*
6*'
FLOOR SEAL PRESSURE MONITORING PIPING SHOREHAM NUCLEAR POWER STATION -UNIT I
+M Z
+Fz PLANT DESIGN ASSESSMENT FOR SRV AND LOCO LOADS
+Az REVISION 4-FEBRUARY 1981
O :2, ,,.
~
+My$ +Ma
+x o* j
+P, +~- c
+0x f i
1.5'DIA 1
l s' '
AG FIELD TO .
, ANALYSED 10 0 g
il! CONT.
PROB.
084' l IE51 95 I l '/"
2 -S LP -lO-151-2
.084, g/
lp s
.084*
as O PRIM ARY
.084 WALL
- PENETR ATION X-42 so 09
.084*
35 9
,0
.084' NOTE SUPPRESSION POOL HIGH WATER LEVEL LEGEND: 2 7 '- o" i Aa MASS POINT o CONSTR AINT POIN T FIG. K3.1.14 MATHEM ATICAL MODEL FOR RCIC SMALL - BORE PIPING
( SHOREHAM NUCLEAR POWER STATION-UNIT I PLANT DESIGN ASSESSMENT FOR SRV AND LDCO LOADS REVISION 4-FEBRUARY 1981
4 Ch^
'5'
+ ry
\- + Ay ,
+u y ") +u, ql o** /
x , / to
+F: + 2 j h
+Az , f 25 l dos oh PENETRATION X 43
/ \ FIELD TO 5 ANALYZED
.005' l
, I/
- l
. s ('
- PRIMARY 33
.01C' yo" d. WALL l/
o4 io l/.
.032 /o 40
.01 ,
o w'
O / .....
< r,5
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D ETAll " A IEt'-
go 2" S Lf' 151 - 2
.064' D.
NOTE .054 25 SUPPRESSION POOL HIGH WATER LEVEL ;
27'- O" d 45 SEE D ETA ll "A" l LEGEND:
As MASS POINT 0: CONSTRAINT POINT FIG. K 3. l.15 l MATHEMATICAL MODEL FOR RHR .
T SMALL- BORE PIPING
\ SHOREH AM NUCLEAR POWER STATION - UNIT I PLANT DESIGN ASSESSMENT FOR SRV AND LOCO LOADS REVISION 4-FEBRUARY 198I
.. ,. .., - - _. ---, . . . , , . . . . - . - - --1
O O O
+%
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.o
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f36 3. b +Ma
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s
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l I
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I NOTE SUPPRESSION POOL l .: 'Nk.dd Jo HIGH WATER LEVEL 2 7 '- O '* N s %
4
$e,
} LEGEND:
A MASS POINT
%\
s j 0 CONSTR AINT POINT 4
FIG. K 3.1.16 M ATHEM ATICAL MODEL' FOR 4
LEVEL SENSOR PIPING
, SHOREHAM NUCLEAR POWER STATION - UNIT I j PLANT DESIGN ASSESSMENT FOR SRV AND LOCO 1.OADS REVISION 4-FEBRUARY 1981
']
1 l
o
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4 .E' I A
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_ . . . .. ; e , , . _ _ . . . . , '. . . . .. __
-.y , :.
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- .- - .6 n .. .
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( *;.9.. ;. ^ -
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n
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.s v Q) v *
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- ' O JD'. s
, i . .... , .
LEGEND 3.g
, ;. .* 6, ). -
l O cO'uun O DOWNCOMER O SRV LINE FIG. K 3.2.1 A TOP VIEW OF MAJOR SUBMERGED STRUCTURES SHOREH AM - NUCLEAR POWER STATION - UNIT I PLANT DESIGN ASSESSMENT FOR SRV AND LOCO LOADS REVISION 4-FEBRUARY 1981
1 8 _
9 _
1 Y
0 S R 8 T D A N A U
O 0 EI LO M TO R
B E
F NGIC I
E ULNO 4 N
S O D D I S
0 AMNN I V
7 ORI OA E 0 LATV AR R
O T RTS EESR
- J H RFO C E RN T
- 0 EE W N H 6 T U OE P
0 A Q SM WARS AE AO NES LS C C A U
ON NN LO G
- 0 5
I
- 2. L T MSI 0 2. A CAE 3C KI R EHDET S .PIDOARN
_m D GY I HL N FTXSP O
0 C E
- 4. S 0
O _m E M
TI 0
- 3
. 0 0
.. 2
- 0 0
1 O
t O 0 0 0 0 0 8 3 8 3 2 1
1 O -9 mJ $ o4 J.. X~.I5,O)u
i 1
8 9
1 Y
S R 0
8 T D A U
N A R 0 EI O B L E M TO GI N F
E U O C 4 N
I SD MN N L N O
S I
DRO A I
0 V 7 I E
OR L ESTS b TH R
)
0 RU OP N E 6 EQ M 0 T A AR A S S
E WNE S w
A OLC S A CNU OO N N G 0
5
- 3. LTMS I I 2.L CAE 0 3 AEH D K CRET I I RN S
D G.PD OA I Y HL
- N FTYSP O
C 0 E 4 S 0 -
C E M
I T
0 3
C 0 0
l 2 C 0 0
C 1 0
3
( 0 0 0 0 0 4 O 4 8 8 - -
> _. FZh 2 $"
J
$ s z$<
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1 8
19 Y
R 0 T SD A O 8 U 0 N A R E ILO B E
M F GTC O -
EIO SNL 4
N U O INM RN N D I S
I V
R 0 A ITR
]
[ - ORAS LET HSR O TC RF
- E J NET EWN 0
6 RUOE EQPM S
0 TARS A AE WNES OLS A CA CNU N OON G 0 4.L TMS I I 5
- 2. L C A E 0 3AEHD KCRET IRN S G IPDOA Y
C HL D
N IFTZSP O
0C
- 4. E 0S O C E M
I T
0 3
C 0 0
2 C O O
C 1
O 3
[ O 0 0 0 0 0 7 5 3 1 O l O z- 81- N_ _
zw 2oW g 4 ) i . j i'i i* l , i ' .
s v v 4
~ l20 , ,
70 u L t m 3
e 20 m !
@ C O b O O O O O O O O O O S
N l p -30 z
W !
2 o
W m
-80 ;
[L _.,
f, w C C w
-130 0 8 16 24 32 40 48 56 TIM E - SECON DS FIG. K 3.2.5 A=x TYPICAL LOCA AIR BUBBLE LOAD ON A O=Y DOWNCOMER SEGMENT SHOREHAM NUCLEAR POWER STATION- UNIT l' O:z PLANT DESIGN ASSESSMENT FOR SRV AND LOCO LOADS i
REVISION 4-FE8RUARY 1981 l
l
O O O d
16 0 l
- u. 80 J
o 8'
40 _._ . - - -
C C, 3 2 us 2
O r' r e t L m r.,
, O w w w
_ m l
i I
-40 i O.56 0.64 0.72 0.30 0.88 0.96 1.04 1.12 1.20 TIME - SECON DS i
OY FIG. K 3.2.6 Oz TYPICAL LOCA POOL SWELL LOAD ON A SEGMENTOF RHR TEST LINE DISCHARGE PlPING SHOREHAM NUCLEAR POWER STATION-UNIT I PLANT DESIGN ASSESSMENT FOR SRV AND LOCO LOADS REVISION 4-FEBRUARY 1981
O O O 80 1
I o[I C C C C C
? pJ V \
m
! o i 8J -80 -
b H
1 z
hJ l
2 uj -iso
%\ <
! m
)
~ \ '
, o 0.20 0.40 0.60 0.80 1.00' l.20 1.40
- TIME - SECONDS i
! O=Y O=Z FIG. K3. 2.7
]-
TYPlCAL LOCA POOL FALLBACK LOAD ON A SEGMENT OF RHR TEST UNE DISCHARGE PIPING SHOREHAM NUCLEAR POWER STATION -UNIT 'l PLANT DESIGN ASSESSMENT FOR SRV AND LOCO LOADS
? REVIStoN 4-FEBRUARY 1981
)
i
( .
, O O O 120 1
1 80 k I i u. 1 y
m 40 E- -
f
{
b
)
\
m oY . n , - -
~v :
3
- 3-
' w 2 -40 - - -
- E ; I
-80
-12 0 0 0.40 0.80 1.20 1.60 2.00 2.40 2.80 3.20 3.60 4DO TIME - SECONDS
=x FIG. K 3.2.8 oaY TYPICAL LOCA CONDENSATION LOAD ON A d'Z SEGMENTOF RHR PUMP SUCTION PIPING SHOREHAM NUCLEAR POWER STATION -UNIT I PLANT DESIGN ASSESSMENTFOR SRVAND LOCO LDADS REVISION 4-FEBRUARY 1981
.~. -
4 80
- / l 40 -
I
$ J rYl i
mvvm
^
x
^ ^ '
Ot
,C s l
i \
- c W
o \ t y -40 i -80
'[f O
O O.05 0.10 0.15 0.20 0.25 0.30 0.35 TIME - SECONDS 0=X 0*Y A:Z FIG. K 3.2.9 TYPICAL LOCA CHUGGING LOAD ON A SEGMENT OF DOWNCOMER SHOREHAM NUCLEAR POWER STATION - UNIT I PLANT DESIGN ASSESSMENT FOR SRV AND LOCO LOADS REVISION.4-FEBRUARY 1981
l
)
1 EkS*
to
$.p'fp 0l' j%
///
Y;>)tfp *
%d.7 TEST TARGET (MT-3) i l.0 lf a EA l8&lf !_ ole
- j. b ;g lil 2.0 I.l U .s c.
i l.8
[
. 1.25 1.4 1.6
- 6" - -- * ,
)
\
4
- w[3#'4p <[>[d4 u
'g 7/ <#. %
f%l*g;9
.gp 4
.sp 'gp e
- Ols'@ %Wf TM+ ___
%d<7 s TEST TARGET (MT-3)
- 1.0 ,T a !?4 o a p'"=
a: i e t y; {!ll2.0 l.l L 4 <
18 ll l l.25 1.4 1.6 1
I i
4 6" > ,
m,% N 3p /$
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o
/ #y%y#
go I
.m e
1 8
9 1
Y S R 0 T D A 8
N A U O O R l
0 E I L B t E M T O F
/ G I C N O -
y N I S E U L 4 N
D O I
NN S o DR NOA I V
7 AE OMI OITV AR E
R 0
u L CAS OTTS E R L NU RO BWTE F BO CWT 0
U D AON PE 6 B AE M 0 RS R NVAS I
l.
SIT L GI 0 AMS 5 2,L C AE 0 3 AEOHD S K CRTET I I RN D .
PDE OA N GY I U HL O FTXD SP C
E oS 4
0 E
O M I
T 0
l' 3
0
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- G o
I.
0
- - f 1
- h
[
0 0 o o
0 0 0 0 6 4 2 o 2 4 6 s U I
O $JI wo ' x. Fb2b*
. i 4 .i ! ! ,i 1 is ! i!j!i I i ,!1:
4 O O O i 80 1
40 - -- - - ---- - - - - - - -- --
4 0
! h N O(> - - ---- -- -- -- - -- - - --
c
$ l in 2
S c -40 --- - - - - - -----
4
-80 0 0.10 0.20 0.30 0.40 0.50 0.60 . 0.70 0.80 TIM E - SECONDS 4
- FIG. K 3. 2.11 TYPICAL SRV AIR BUBBLE LOAD IN Z-DIRECTION ON A DOWNCOMER SEGMENT DUE TO ALL' VALVE ACTU ATION SHOREHAM NUCLEAR POWER STATION - UNIT I
- PLANT DESIGN ASSESSMENT FOR SRV AND LOCO LOADS
, REVISl0N 4- FEBRUARY 1988
.