ML19269C444
| ML19269C444 | |
| Person / Time | |
|---|---|
| Site: | North Anna  |
| Issue date: | 01/31/1979 |
| From: | VIRGINIA POWER (VIRGINIA ELECTRIC & POWER CO.) |
| To: | |
| Shared Package | |
| ML19269C443 | List: |
| References | |
| NUDOCS 7902020242 | |
| Download: ML19269C444 (89) | |
Text
{{#Wiki_filter:f r . DESCRIPTION OF NEUTRON SUPPLEMENTARY SHIELD Prepared for: VIRGINIA ELECTRIC AND POWER COMPANY By: STONE & WEBSTER ENGINEERING CORPORATION 7902020AO
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t I-1 I. Introduction The radiation levels inside the reactor containment determined by radiation surveys on Unit 1, are greater than the design levels presented in Section 12 of the FSAR, at two locations. These locations are as follows:
- 1. The a nnulus area between the crane wall and the containment wall on the operating floor (El 291 f t-10 in) at crane wall openings
- 2. Inside the personnel lock The survey results indicate dose rates on the operating floor in the annulus area at openings in the crane wall on the order of 2,500 mrem /hr neutron and 200 mrem /hr gamma. The gamma radiation levels are primarily attributable to neutron capture reactions in the containment concrete and steel structures. This conclusion is consistent with thermal neutron flux measurements on the order of 3 x 10* n/cm2 sec using thermoluminescent dosimetry. The survey results indicate dose rates in the personnel lock on the order of 40 mrem /hr neutron and 2 mrem /hr gamma.
Based on the higher than anticipated radiation levels inside the containment, additional neutron shielding has been designed for installation in Unit 2. The shield's neutron attenuation effectiveness has been conservatively calculated, and the safety analysis demonstrates that the installation of the proposed shielding will not have any ef fect on the safety of the plant or the integrity of reactor vessel support system and will substantially reduce t combined neutron and gamma dose rates in the personnel lock and in those areas required for general containment access.
II-1 II. Neutron Shield Design criteria , The neutron shield is designed to:
- 1. Reduce radiation levels in that portion of the annulus area between the crane wall and the containment wall on the operating floor that is required for general containment access and in the personnel lock to those presented in Section 12 of the FSAR.
- 2. Be a structure that does not require removal during refueling with the concurrent personnel radiation exposure.
- 3. Have no effect on the safety of the plant or the integrity of the reactor vessel support system and reactor coolant system, including the effects of reactor pressure vessel internals response and cavity pressure, as presented in Appendix SA of the FSAR.
- 4. Be a structure incapable of becoming a potential missile that could adversely affect any safety-related equipment.
- 5. Permit required Inservice Inspection of reactor vessel nozzle and piping welds.
III-1 III. Ef fectiveness of Collar / Saddle Shield A. Summary The ef fectiveness of the collar / saddle shield in reducing neutron streaming from the reactor cavity was assessed by two distinctly different calculational meth ods. The first method involved the use of the COHORT-II Monte Carlo programCt) in an analog mode starting with an isotropic surface source at the outside surf ace of the reactor pressure vessel. The second method involved the use of the MORSE Monte Carlo program (2) with neutron albedo representations of surface scattering and an anisotropic source at the outer surface of the reactor pressure vessel. The approach used in each method is described in Section III.B. The dose rates in the crane wall openings were calculated using both Monte Carl' programs without the collar / saddle shield in place and compared to measurements at North Anna Unit 1. The results of these calculations are tabulated in Table III-1. The calculated dose rates using both MORSE and COHORT II are in agreement with the measured values; however, there is evidence that the PNE-4 detector used in the radiation survey for Unit 1 overpredicts the neutron dose rates'by as much as 92 percent due to an overresponse to intermediate energy neutrons (3). Since the calculated neutron spectra for Unit 2 is similar to those measured at other PWRs, we can expect that the actual neutron dose rates will be significantly lower than the calculated values indicate. The neutron dose rates were then calculated for the same detector locations (as the unshielded case) with the collar / saddle shield in place using both Monte Carlo computer programs. Table III-2 compares the neutron dose rates for the two calculational methods. In the MORSE analysis, the dose rates were calculated by a boundary crossing estimator using the multicollision Snyder-Neufeld tissue dose response function. By contrast, the COHORT II analysis used a volume detector to calculate the neutron doce rates and the American National Standard (ANSI / ANS-6.1.1-19 7 7) for tissue dose response functionC*). Based on the results presented in Tables III-1 and III-2 and choosing the highest calculated dose rate around the peripheral region of the containment (detector location 5) , the neutron dose equivalent rate is expected to be between 25 and 50 mrem /hr, taking credit for instrument overresponse. B. Methodology The calculation of the dose rates in the reactor containment requires the use of three-dimensional Monte Carlo radiation transport meth od s. The problem lends itself, however, to decomposition into several stages which substantially improve the J
4 III-2 computational e f ficiency. The problem has been decomposed into three steps as shown in Figure III-1.
- 1. Calculation of the neutron transport from the reactor pressure vessel (RPV) to the top of the neutron shield tank (NST) , (Leg 1) .
This segment of the analysis is the calculation of the radiation streaming in the gap between the RPV and UST. The source term for Leg 1 was the RPV leakage term. Each analysis used different methods to derive the RPV leakage term. COHORT II: In this method, the reactor pressure vessel source was modeled as an isotropic cylindrical source which was embedded 0.7 cm into the outer surface of the pressure vessel. This slight embedding was included to take out some of the conservatisms of the isotropic source assumption. The magnitude of the leakage was calculated with the COHORT II Monte Carlo prog ra m. The core midplane value was normalized to the total flux in each of four broad energy groups as calculated by the NSSS vendor. MORSE: In this analysis, one-dimensional discrete ordinates calculations were performed to obtain the BPV leakage term. NSSS vendor supplied data of the axial and azimuthal dependence of the fast neutron flux were then used partially to correct the results of the one-dimensional model. This procedure has been shown to be a valid one by comparison with more detailed two-dimensional discrete ordinates calculations (St. The neutron leakage was tallied in the gap at the top of the NST to serve as a source term for Leg 2.
- 2. Calcula tion s of the neutron transport from the top of the NST to the seal plate elevation (with and without the collar / saddle shield) , (Leg 2) .
This segment of the decomposition advanced the analysis through the region containing the shield. The calculation was performed using both Monte Carlo computer programs with their respective source terms as calculated in Leg 1. The neutron leakage from the reactor cavity was then talli'ed at the seal plate elevation for subsequent use as a source term for Leg 3.
- 3. Calculation of the neutron transport from the seal plate elevation to dose points on the operating floor, (Leg 3) .
The final segment of the decomposition was the calculation of neutron dose rates at selected points on the operating floor. In the MORSE analysis, the dose rates were calculated by a statisti-cal flux estimator using the multicollision Snyder-Neufeld tissue J
III-3 dose response function . By contrast, the COHORT II analysis used a volume detector to calculate the neutron dose rates in the crane wall openings above the operating floor and the ANSI /ANS-6.1.1-1977 for tissue dose response function") . 1 4
III-4 C. Geometry Description The geometry models used in the collar / saddle shield analysis for each step is described below. The geometry model for both analyses (COHORT II and MORSE) were virtually the same. All of the figures used in the shield analysis were genera ted by a computer program from the geometry input of the MORSE calculation.
- 1. Leg 1 the model for Leg 1 is shown in Figure III-2.
The model consists of several cylinders of differing radii one inside the other. The Leg 1 model stops at the top of the NST.
- 2. Leg 2 -
the RPV and surrounding shield material have cylindrical symmetry except for the protruding nozzles; however, the relative nozzle position repeats itself
- every- 120 deg so it was only necessary to model one-third of the problem. Figure III-3 shows a horizontal cross section of the model at an elevation of 256.33 ft (corresponding to Z=825 cm in the Combinatorial Geometry Model) . An elevation of 229.26 ft corresponds to z=0 cm. The figure shows the RPV, hot and cold leg nozzles, the concrete shield, and the collar / saddle shield. Figure III-4 sho.is a vertical cross section of the cold leg nozzle perpendicular to the nozzle axis.
The cut is at a radius of 260 cm. The nozzle was initially modeled as solid, but subsequent models incorporated a hollow nozzle. Figure III-5 is a vertical slice of the cold leg nozzle parallel to the nozzle axis. The hot leg nozzle was modeled in a similar fashion. Figure III-6 shows a vertical cross section through all of the Leg 2 model including the cold leg nozzle.
- 3. Leg 3 - Leg 3 consists of everything above the RPV seal plate elevation. This includes the top of the RPV, refueling channel, missile shield, control rod drive mechanisms ( CRDMs) , containment building, and major structures within the containment such as the steam generator and pressurizer housings and the crane wall.
Figure III-7 is a horizontal cross section of the refueling channel between the top of .the RPV and the bottom of the operating floor. The region containing the CRDMs (homogenized into one region) is not shown in this figure. Figure III-8 is a horizontal cross section at an elevation between the operating floor and the top of the lowest openings in the crane wall. Figure III-9 is a vertical cross section of the containment.
III-5 D. Conclusion The assessment of the effectiveness of the collar / saddle shielc was concentrated at the openings in the crane wall above the operating floor. The ef fect of the crane wall is such that the dose rates in the annular region between the crane wall and containment wall will be a fraction of those levels nredicted air for the openings. Similarly, the dose rates in the personnel lock are expected to be well within the 2.5 mrem /hr criterion at that location due to the effectiveness of the collar / saddle shield. It is also expected, as noted previously, that the actual neutron dose rates will fall in the range predicted by the two analyses. For the highest neutron radiation area in the annular region on the operating floor (detector location 5), this would indicate values ranging from 25 to 50 mrem /hr. Since the gamma dose rates on the operating floor are primarily attributable to (n,l) reactions with the containment concrete and liner, we expect the combined neutron-gamma dose rates in the annular region between the crane wall and containment wall to be below the 100 mrem /hr criterion. To reduce even further potential e::cosure rates, openings in the crane wall between the personnel lock and the elevator (as shown in Figure IV-4) will be blocked with 3 inches of Permali, Type JN. sl'
TABLE III-1 COMPARISON OF CALCULATED 1:EUTRON COSE PATES WITH _ - MEASUREME!TTS MACE AT NORTH ANNA UNIT 1 - ACJUSTED TO 100% POWER Nautron Core Rate (mrem /hr) Analytical Plux to Dose cose Point
- Approach Resporise Function 3 4 5 6 Type of Data _
ANSI /ANS-6.1.1-1977 1,920 2,570 2,930 2,410 Calculated Dose COIDRT II Snyder-Neufeld 2,260 3,300 2,420 2,300 Equivalent Rate MORSE , Measurement (Uncorrected for Instrumen t 2,640 2,860 1,430 Overresponse) 2,090
- Refer to Figure III-10 1 of 1 7
TABLE III-2 - CALCULATED NEUTRON DOSE RATES WITH SUPPLEMENTARY NEUTRON SIIIELDING Expected Neutron Dose Rate As Measured With PNR-4 Detector (mrem /hr) Analytical Approach Dose Point
- 2 3 4 5 6
_1 COHORT II Method - 190 82 77 96 66 MORSE Method 285 45 17 25 25 19
- Refer to Figure III-10 NOTE: Detector location 2 is on the inside of the crane wall (i.e. surface of Permali Shield, Type JN).
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IV-1 IV. Shield Design , A. Description The supplementary neutron shield is comprised of these main components:
- 1. Collar Assembly: As shown in Figure IV-1, the cylindrical collar assembly is comprised of six segments, each with an extended base and centering tabs.
The segments rest on the top of the usT and are f astened together by a metal strap to form the collar. The collar fits around the reactor pressure vessel (RPV) over the insulation and extends to the spaces between the nozzles. Each collar segment consists of an outer steel casing, filled on the inside with a silicon based neutron attenuating material.
- 2. Saddle _ Assembly: As shown in Figure IV-2, the saddle assembly consists of U-shaped blanket type covers for the RPV nozzles, extending from the collar interface to the primary shield wall. The saddles are comprised of approxima tely 130 one-quarter inch wide strips of silicon based neutron attenuating material por nozzle.
- 3. Dust Cover Blocks: The dust . cover blocks are silicone based neutron attenuating material blocks encased in stainless steel sheet metal. The blocks are shaped to cover the dust covers on the RPV nozzle support structure and to partially fill the space between the dust cover and the collar base underneath each nozzle as shown in Figures IV-1 and IV-3.
- 4. Crane Wall Area Shielding: Neutron attenuating shic1d material will be placed in the crane wall openings extending f rom directly opposite the personnel hatch to the elevator entrance and over the portion of the fuel transfer canal behind the crane wall, as shown in Figure IV-4.
B. Location The neutron shielding components, with the exception of the shielding in the crane wall openings, are all located inside the upper reactor cavity. The base of the six collar segments rest on the top of the NST. The collar segments are strapped together into contact with the RPV insulation. In this position, the callar segments are placed directly in the path of escaping neutrons emanating from the annulus between the RPV and the NST. The one quarter inch U-shaped saddles are positioned on the nozzles over the thermal insulation along the length of the nozzles, as shown in Figure IV-2.
IV-2 The dust cover blocks, shown in Figures IV-1 and IV-3, are positioned on top of NST around the dust covers underneath the nozzles. Shielding is located in those crane wall openings shown in Figure IV-4. The layout arrangement of the collar / saddle shield is shown in Figure IV-S. C. Materials The neutron attenuating material used in the collar, saddles, and dust cover blocks is a silicon based elastomer with a hydrogen density of approximately 0.06 gm/cc (4.3 percent by weight) . The shield material will be impregnated with boron carbide c (Ba ) to 2.0 percent by weight, with the resultant effective boron density of 0.02 gm/cc (1.5 percent by weight)-. The material used for attenuating neutrons in the crane wall openings in Permali, Type JM, a densified beechwood laminate which incorporates six percent hydrogen and three percent baron. The outer wall of the collar segments is constructed of 3/8 in carbon steel, and the inner wall is 10 gage stainless steel. The dust cover blocks are encapsulated with stainless steel. D. Supports The entire extended base of the collar rests on top of the NST. The inner cylindrical surface rests against the RPV insulation. Additionally, . collar segments are held together by a metal belt wrapped around the collars at the top. The saddle elements rest upon the top of the nozzle insulation and they are axially restrained by angle iron attached to the nozzle insulation shell. The dust cover blocks rest on top of the NST and RPV nozzle support structure dust covers and are laterally restrained by the collcr base. Shielding sections are supported in the crane wall openings by a steel franie work attached to the crane wall. E. Missile Effects The only credible missiles are the saddle strips on the nozzle of a postulated broken reactor coolant pipe. The jet force of the flow from the broken pipe ends may cause the strips to exit the upper reactor cavity. These missiles would be long thin (1/4 in wide) low mass strips of silicon elastomer with low rigidity.
IV-3 Ile n c e , they will not adversely affect any safety-related equipment. The collar segments are not expected to be potential missiles because of the following reasons:
- 1. The collar is located so that it is not subjected to direct jet impingement forces from the postulated limited displacement breaks.
- 2. The pressurization of the reactor cavity due to the mass and energy released from the break would force the collar segments down against the NST, against each other, and against the RPV insulation.
- 3. The metal belt around the collar, together with centering tabs at the base of each segment, will keep the collar assembly in place.
Under LOCA conditions, the dust cover blocks will not become missiles because they are not exposed to lifting forces on any surface. F. Effect on Containment Sumn Saddle strips that may be propelled by the jet force onto the operating floor would be required to follow a complex and tortuous path through gratings or down stairwells to reach the sump level. The strips have a density greater than that of water and will not float. If any of them reach the containment floor, it is unlikely that they will be transported toward the sump by the containment spray water because of the low water velocity to the cump. The sumps have screens provided as described in Section 6. 2. 2.2 of the FSAR and, thus, would prevent any caddle strips from reaching the sump.
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V-1 V. Reactor cavity Pressure Transient Analysis A 27-node model was used to calculate the pressure-time history in the reactor cavity following a postulated 150-in2 cold-leg, limited-displacement rupture (LDR) . The 27-node model is an extension of the 21-node model that was previously used to calculate the pressure transient reported in the FSAR. Six additional nodes were added in the refueling cavity to more completely nodalize the volume around the reactor vessel. The refueling cavity was axially divided by a horizontal plane through the centerline of the vent ducts that are connected to the control rod drive shroud, as shown in Figures V-3 and V-4. The refueling cavity was circumferential1y divided by extending the vertical planes that pass through the centerline of each nozzle. Plan and elevation views showing the general arrangement of the reactor cavity structures and piping are presented in Figures V-1, V-2, and V-3. This model is the same as the model shown in Figures 6.2-29 and 30 of the FSAR, except for the six additional nodes in the refueling cavity. A block diagra m showing the nodalization of the reactor cavity with the net free nodal volumes and interconnecting flow paths is given in Figure V-4. This is the same diagram as FSAR Figure 6.2-28, except for the six additional nodes in the refueling cavity. Reduced nodal volumes in the upper reactor cavity in Figure V-4 are due to the addition of the neutron shield in these nodes. The computer c ode RELAP4, MODS ( 8 ) (with air) was used to calculate the pressure-time transients. The vent areas, effective L/As, and vent loss coefficients used to calculate flow between nodal volumes are presented in Table V-1. This differs from FSAR Table 6.2-21 as follows: (i) Data for the six additional nodes in the refueling cavity are included. (ii) Effective L/As are included. (iii) All data related to the THREED code has been deleted. (iv) Vent areas and loss coefficients for nodes in the upper reactor cavity have been modified to reficct the installation of the neutron shield. The vent loss coefficients were obtained from References 10 and 11. No credit was taken for venting out of the shield wall penetration of the broken pipe. This is similar to the analysis presented in Section 6.2.1.3.2 of the FSAR. No credit was taken 4
v.2 for venting into the reactor annulus under the broken pipe. The insula tion and shielding material on the unbroken nozzles were assumed to remain in place for vent area and volume calculations. The reactor flange-mounted ventilation seals at the + ap of the upper reactor cavity were assumed to completely block flow out of the top of the cavity while initially in place. They were assumed to be instantaneously blown away when the design differential pressure of 18 psid, a final design (nominal) value across the seals was reached. For the analysis presented in FSAR Section 6.2.1.3.2, a blowout pressure of 10 psid was assumed. Studies have shown that peak pressures in the upper reactor cavity are insensitive to the blowout pressure. The postulated break is a 150-square inch LDR at the reactor vessel inlet nozzle safe end. The mass release rates and blowdown enthalpy are presented in Table V-2. The mass release rates are the same as those given in FSAR Table 6.2-7. Due to code input requirements, a constant enthalpy of 561.2 Btu /lbm was used, whereas the blowdown enthalpy as given in Table 6.2-7 varies from 556.4 to 561.2 Btu /lbm. Figure s V-5 through V-8 present the pressure transients in the reactor cavity for selected nodes.
TABLE V-1 Vent Areas, Effective L/A's, and K-Factors Used ii. the Reactor Cavity Analysis With RELAP4, MODS Loss Coefficients, K From To Effec-Vol. Vol. Vent tive Con- Ex- Input Node tiode Area L/A trac- Fric- pan- to No. No. ft2 ft-t , tion tion Bend sion RELAP(21 1 2 13.03 .25 .2 .048 -
.21 .458 e 3 12.54 .603 .05 .075 .084 .061 .27 1 11 13.71 .595 .03 .075 .084 .04 .23 1 22 20.1 .23 .02 .05 -
.32 .39 2 4 3.08 1.43 .24 .118 .15 .32 . 83 2 12 3.63 1.36 .2 .118 . 15 .28 .75 2 13 .15 6.07 .5 .057 1.78 -
2.33 3 4 21.88 .17 .11 .038 -
.12 .27 3 5 13.0 .608 .04 .075 .084 .04 .24 3 23 28.14 .14 .02 .048 -
.53 .598 4 6 3.63 1.48 .2 .118 .15 .19 .66 4 14 .33 4.34 .5 .073 1.78 -
2.35 5 6 12.31 .26 .2 .048 -
.24 .488 5 7 12.54 .604 .06 .075 .084 .05 .27 5 24 21.88(1) .2 02 . 05 -
53 . 60 6 8 3.08 1.44 .25 .118 .15 .27 .79 6 15 .33 6.08 .5 .057 1.78 - 2.33 7 8 21.88 .17 . 11 .038 -
.12 .27 7 9 13.0 .604 .03 .075 .084 .05 .24 7 25 29.92(1) .14 .02 .048 -
. 53 .598 8 10 3.63 1.44 .17 .118 .15 .26 .7 8 16 .33 4.34 .5 .073 1.78 -
2.35 9 10 12.31 .26 .2 .048 -
.24 .488 9 11 12.54 .6 .06 .075 .084 .05 .27 9 26 20.1 .23 .02 .05 -
.32 .39 10 12 3.08 1.41 .27 .118 .15 .27 .81 10 17 .33 6.08 .5 .057 1.78 -
2.33 11 12 22.6 .16 .1 .038 -
.11 .25 11 27 28.14 .16 .02 .048 -
.31 .378 12 18 .15 4.33 .5 .073 1.78 -
2.35 13 14 8.711 1.84 -
.482 .056 -
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.54 13 19 1.561 6.03 - . 04 -
.98 1.02 14 15 8.711 1.84 -
.482 .056 -
. 54 14 19 2.186 4.34 -
06 -
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.482 .056 -
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.482 .056 -
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.06 -
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.98 1.02 1 of 2
TABLE V-1 (CONT ' D) Loss Coefficients t K From To Effec-Vol. Vol. Vent tive Con- Ex- Input Node Node Area L/A trac- Fric- pan- to No. No. ft2 ft-1 tion tion Bend nion REL AP (2) 18 19 2.186 4.34 -
.06 -
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.017 -
1.0 1.02 23 21 268.8 . 04 -
.01 -
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.011 -
10 1.01 25 21 268.8 .04 -
. 01 -
10 1.01 26 21 58.7 .09 -
.017 -
1.0 1.02 27 21 82.2 .06 -
.013 -
1.0 1.013 22 23 62.75 .22 -
.045 .24 .085 .37 23 24 62.75 .20 -
.045 .24 -
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.045 .24 -
.29 25 26 62.75 .22 -
.045 .24 .085 .37 26 27 38.75 .25 -
.061 .107 -
.17 27 22 38.75 .25 .061 .107 -
.17 1 21 .923 4.91 .o .066 -
1.0 1.57 2 21 .923 4.95 .5 .066 - 1.0 1.57 3 21 1.783 2.54 .5 .09 - 1.0 1.59 4 21 1.783 2.57 5 .09 - 1.0 1.59 6 21 1.783 2.6 .5 . 09 - 1.0 1.59 8 21 1.783 2.57 .5 .09 - 1.0 1.59 9 21 1.783 2.55 .5 .09 - 1.0 1.59 10 21 1.783 2.6 .5 .09 - 1.0 1.59 11 21 .923 4. 9 .5 .066 - 1.0 1.57 12 21 .923 4.92 .5 .066 - 1.0 1.57 NOTES:
- 1. The code RELAP4, MOD 5 allows a total of 21 junctions for any two connecting volumes. To input all vent areas, the shield wall penetrations from Nodes 5 and 7 were added to 'the vent areas from Nodes 24 and 25, respectively.
- 2. The loss coefficients between two connecting nodes were assumed to be equal in the forward and reverse direction.
2 of 2
TABLE V-2 Mass Release Rates and Enthalpy For A 150 In2 Cold Leg LDRCE) Time Mass Rate (sec) (lbm/sec) 0.0 0.0 0.251E-2 1.24792E+4 0.500E-2 1.66373E+4 0.751E-2 1.91765E+4 0.1004E-1 2.13965E+4 0.2005E-1 2.54591E+4 0.2501E-1 2.46335E+4 0.3257E-1 2.55243E+4 0.4501E-1 2.73888E+4 0.5513E-1 2.68367E+4 0.7003E-1 2.54029E+4 0.7750E-1 2.59409E+4 0.9254E-1 2. 3 7 9 3 5 E+ 4 0.11003 2.29650E+4 0.12516 2.39874E+4 0.14013 2.31753E+4 0.15502 '2.27749E+4 0.17012 2.21401E+4 0.19006 2.26099E+4 NOTE:
- 1. Enthalpy is conctant at 561.2 Btu /lbm through the transient.
1 of 1
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FIGURE V-4 BLCCK DIAGRAM FCR THE REACTM CA'/ITY (27 !; ode Mcdel) 21 Ccntainment 1,V10,C M ft3 El.270'-10" 0 50* 120" 170 240' 290" 360" l 22 I aj l c4 l Z:, } T6 .
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3 Tvelve vents are not shown. These are the vents from nodes 1 through 12 to node 21 through the piping penetrations in the chic]d vall. 4 The following symbol representations are used: CL - Cold Leg IIL - IIot Leg
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VI-1 VI Reactor Pre ssure Vessel _ support Intearity Review A comprehensive recvaluation of the RPV support system was conducted in a manner similar to that described in Appendix SA, Section SA.4.4 of the FSAR, entitled " Integral Support - Phase IIIC. The pressure transients (Figures V-5 through V-8) developed in accordance with Part V of this report were transformed into asymmetric force-time-histories and moment-time-histories for application to both the RPV and internal structures (Figures VI-2 through VI-7,. These figures are analogous to FSAR Figures 5A.4-69 and 5A.4-70. In this regard, the unbalanced forces on the RPV and primary shield wall (PSF) were higher tha- previously determined. Peak horizontal RPV force increased from 1,540 kins to 1,660 kins and peak moment increased from 26 x 103 inch-kips to 49.5 x 103 inch-kips. A recalculated RPV support stiffness, utilizing additional flexibility in the sliding block, was used in the development of reactor pressure vessel and primary shield wall motion in response to forces on the RPV. Results, developed by Westinghouce, are shown in Figures VI-8 through VI-12. The most important changes involved the so-called Case 1 (maximum horizontal RPV displacement) . The maximum horizontal displacement in fact was relatively unchanged (from .072 inch to
.071 inch) , but had to be combined with RPV rocking vs 0.000517 radians) present at this now, (0.00038 radians slightly shifted time point (f rom 0.070 see to 0.0737 sec) .
These new displacements were combined with revised primary shield vall (PSW) asymmetric pressure response data in the same manner as for the FSAR. New loads for the RFV support and neutron shield tank (NST) were developed and are presented in Tables VI-1 and VI-2. These are analagous to Tables 5A 4-14 and SA.4-15 in the FSAR. RPV nozzle support loads are shown to be higher than previously reported. None are concluded, however, to exceed the integrity definition inherent in Figure VI-1. This figure, analogous to FSAR Figure SA. 4-7 4 shows that the new load data remains within the structural integrity limit envelope developed in the FSAR. Revised relative displacement data are presented in Table VI-3 and relate to FSAR Table SA.4-16. While these data again show differences with the FSAR, these values are shown to have little effect when ccmpared with the allowable displacement envelope developed in FSAR Figure SA.4-76. The FSAR reports, that for two of the six supports, maximum calculated cap screw strain is 8.3 percent. This maximum value, when computed by the bounding conservative method used for Appendix SA, increases to a value of 12 percent. This value represents a change from 57 percent to 80 percent of ultimate
VI-2 strain. Ilowever, it is approprit.te to reemphasize the point that these strains are determined conservatively on a load basis from the clastic system model. The local plasticities in the cap screws permit relatively large displacements before load based strains of this magnitude can develop. Reference to FSAR Figure SA. 4-76 presents this point, indicating displacements of the order of .4 inch'to develop such strain levels. A more realistic strain could be estimated from the imposed RPV nozzle displacement, relative to the NST which can be obtained from Table VI-3. These relative displacements, if plotted, would still fall within the approximate bounds of the " worst case envelope" designated on FSAR Figure SA.4-76, and would indicate displacement based strains certainly less than 2 percent. This would clearly be within the rationale, developed in Part 5 of Appendix SA, with respect to ASME Section III, Appendix F. It is, therefore, concluded that fundamental conclusions relating to the integrity of RPV supports and to the extent of permissible local plasticity are unchanged f rom those reported in the FSAR. The reevaluation of the system included assessment of changes in load effects in the steam generator and reactor coolant pump supports. No design basin loads were affected and no changes to data reported in Section 5.5.9 of the FSAR are required. Analysis of the neutron shield tank and primary shield wall showed that the applied loads are within the material capability of these components. The ECCS branch piping for Unit 2 was stress analyzed by the method described in Appendix SA of the FSAR. This evaluation showed that the ECCS branch piping remains integral. The results of the Unit 2 ECCS piping stress analysis are given in Tables VI-4 through VI-13. The line numbers listed in these tables are shown in Figures VI-13A through VI-24. FSAh Tables SA.4-19 through SA.4-27 and Figures SA.4-68 through SA. 4-85 show North Anna Unit 1 lines only.
TABLE VI-1 PPV SUPPORT AND NST LCADS PHASE . . ytt) FCk) V M P V M T Load Tyoe M V SW SW B n 3,509 370,268 1,067 268 31,897 7,296 Pipe Rupture (s) 1,253 1,249 232,467 t316 t278 184,658 t3,883 Seis mic t121 181 2259 3,768 f,02,735 1,383 546 116,555 11,179 Total 1,374 .1,330 Design capability of 10,433 6,260 545,964 745,955 NST/RPV Support 844 1,000 25,748(3) 617,993(3) tioTES:
- 1. Per support foot as shown in Figure SA.3-3. See Figure SA.3-3 for load application locations.
- 2. J.scludes internals due to break No. 2 plus dead weight plus asynnetric pressurization loading on the PSW, RPV, and NST.
- 3. Based on weighted average of mill test reports.
- 4. Units: Inches, kips e
9 1 of 1
~
r.
~
r TABLE VI-2 p# RPV NOZZLE SUPPORT LOADS PHEE
- INCLUDE RFV I!;rEH! ALG MOV E:iEN T ASYK4ETRIC PhESSUPE, UE.A3 WI:IGHT, A :D SEISMIC O
3(18 4(13 St il 6(13 1(t) 2( 1 ) F e F F F F F F F " F F F F 'I V Ti:-e V H V H V H V H V H coment (Sec) 925 -986 58 239 -1,647 O
.07373 1,253 291 -1,224 910 -321 1,249 - 1, 2 38 Maxinus Horizontal 666 488 380 -171 302 -509 403 -403 610 -158 Maximum Vertical Up .1650 517 551 O
-962 -549 -749 -1,500 264 -2,120
-1,275 925 -597 -252 -76 Maximum Vertical Down .1400 974
-1,126 -011 -880 -1,004 232 -1,382 O 1,139 -973 1,090 -886 -264 -663 Maximum Relative !!orizontal .1350
-1,216 841 -962 746 291 -2,643
-318 1,197 702 -312 1,151
.0800 1,233 O Maximum Potation NOTES: O
- 1. Nozzle Support Designations (see Figure 6 A. 4-75) .
g
- 2. Ref. Figure SA.3-3 for Load Application Directions.
- 3. Units: Kips C C
'o a
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TABLE VI-4 ECCS STRESS
SUMMARY
RC Loco 1 r, RHR System Faulted End Displ. Stress (Vsi) Sum of Stress (Msi) Allow. Line Pt. Component Stress Unbroken DrcKen Unbroken Eroken Stress Nu-ber & Descriction (Ksi) Loop Loon jog _ Loot 3 S,jxsil 14"-RH-401-164 Loop Nozzle' 16.9 1.6 1.8 18.5 18.8 49.0 1501-Q1 14"-RH-401-165 Elbow 1.4 1.5 17.5 17.7 49.0 1501-01 16.1 14"-RH-401-166 Elbow 15.7 1.1 1.3 16.9 17.1 49.0 1501-Q1 14"-RH-401-1501-Q1 168 Sockolet 18.6 1.1 1.2 19.8 19.9 49.0 14"-RH-401-169 Elbow 15.6 1.0 1.2 16.7 16.8 49.0 1501-Q1 14"-RH-401-1501-Q1 170 Elbow '65.5 0.7 0.8 16.2 16.3 49.0 14"-RH-401-1501-Q1 171 Elbow 15.7 0.4 0.4 16.1 16,2 49.0 14"-RH-401-1501-Q1 174 Elbow 16.9 0.7 0.9 17.6 17.8 49.0 14"-RH-401-1501-Q1 175 Elbow 17.6 1.0 1.3 18.6 13.9 49.0 14"-RH-401-1501-Q1 176 Elbow 17.9 1.3 1.6 19.2 19.5 49.0 NOTES:
- 1. Refer to Figures VI-19A and 193.
- 2. Worst case for either unbroken loop.
1 of 1
TABLE VI-5
' ~
ECCS STRESS
SUMMARY
LMSI & RHR System Faulted End DisSI. Strenn (Ksi) Sun of Stresn (Kni) Allow. Line Pt. Component Stress Unbroken Broken Unbroken Broken Stress Nor.ber No Cescription (Kni) Loon Loor_ Loon Loop 3 Sm (Knil 12"-RC-424 250 Loop Nozzle, 10.6 1.1 8.6 11.7 19.2 51.6 12"-RC-424 250 Elbow 19.6 8.4 6.2 20.4 25.8 51.6 12"-RC-424 252 Elbow 20.2 1.3 9.9 21.5 30.1 51.6 12"-RC-424 253 Elbow 20.0 1.1 8.4 21.1 28.4 51.6 12"-SI-469 256 Elbow 18.8 0.4 3.1 19.2 21.9 60.0 12"-SI-469 257 Elbow 18.1 0.3 2.4 18.4 20.5 60.0 12"-SI-469 258 Elbow 17.8 0.3 2.2 18.1 20.0 60.0 12"-SI-469 265 Pipe 9.4 0.8 6.1 10.2 15.5 60.0 12"-SI-469 259 Elbow 17.6 0.2 1.9 17.8 19.5 60.0 12"-SI-469 260 Elbow 17.2 0.2 1. 7 17.4 18.9 60.0 NOTES:
- 1. Refer to Figure VI-23.
- 2. Worst case for either unbroken loop.
e 1 of 1
s TABLE VI-6 ECCS STRESS
SUMMARY
RC Looo 1, Low Head Safety Iniection Systen Faulted End Discl. Strens (Ksit Sum of Stress (Kni) Allow. Stress Pt. Ccaponent Stress Unbroken Broken Unbroken Broken Line Loop 3 Sm (Ksi) Description Loop Loop T.o o n Number No. (MsiL_ 1.1 9.6 9.6 18.1 49.8 12"-PC-422 250 Loop Nozzle 8.6 10.7 17.9 27.3 49.8 12"- F;C- 4 2 2 251 Elbow 16.6 1.3 8.7 17.5 25.1 49.8 12"-PC-422 252 Elbow 16.4 1.1 9.0 25.0 25.0 60.0 12"-SI-467 270 Elbow 15.2 9.8 17.3 0.012 0.2 17.3 17.5 60.0 12"-SI-467 291 Elbow 12"-SI-467 292 Elbow 17.4 0.017 0.2 17.4 17.6 60.0 0.017 0.2 17.5 17.7 60.0 12"-SI-467 293 Elbow 17.5 0.030 0.3 20.5 20.8 60.0 12"-SI-467 295 Sockolet 20.5 12"-SI-467 299 Elbow 17.8 0.041 0.5 17.0 18.3 60.0 0.036 0.4 17.2 17.6 60.0 12"-SI-467 300 Elbow 17.2 NOTES:
- 1. Refer to Figures VI-19A, 19D, 20h, and 203.
- 2. Worst case for either unbroken loop. .
9 e M 1 of 1
e TABLE VI-7 . ECCS STRESS
SUMMARY
7 RC Loop 2, RHR and LHSI C Faulted End Displ. Stress (Ksi) Sun of Stress (KsLL Allow. Line P t. Component Stress Unbroken Broken Unbroken Brcken Stress e* Ne Descriotion (Ksil Loop Logg_ Loop Loog_ 3 Sn (Ksi) Norber 12"-RC-423 250 Loop Nozzle 10.4 1.0 8.8 11.5 19.2 51.6 g 0.9 7.8 20.1 27.0 51.6 12"-RC-423 253 Elbow 19.2 r 18.8 0.8 6.5 19.6 25.4 51.6 12"-RC-423 254 Elbow 0.3 2.3 19.8 21.8 51.6 12"-RC-423 262 Valve 19.5 1.9 19.1 20.8 60.0 12"-SI-468 263 Elbow 10.9 0.2 I 0.1 2.1 19.7 21.7 60.0 12"-SI-468 278 Elbow 19.6 0.1 1.5 20.2 21.5 60.0 12"-SI-468 279 Elbow 20.1 0.1 0.6 21.0 21.6 60.0 12"-SI-468 288 Tee 21.0 12"x12"x10" ( Elbow 0.0 0.3 17.3 17.6 60.0 10"- R!!- 412 294 17.3 I 336 Sockolet 20.6 0.0 0.4 20.6 21.0 60.0 12"-SI-468 ( FSIE.E*
- 1. Ref er to Figures VI-21 and VI-22.
- 2. Worst case for either unbroken loop.
e 1 of 1
b
- . r TABLE VI-8 g
ECCS STRESS
SUMMARY
Safety Iniection system, Hot Le2 O Faulted End Displ. Stress (Kni) sun of Stress (Knit Allow. Stress () Unbroken croken Unbroken Broken Line P t. Component Stress 3 rn (Ksi)- r.co p Looe_ Loon Loon _ Nu?her No. Description _JKe'L_ 15.5 60.0 q) 0.1 15.5 6"-SI-419 25 Pipe 15.4 0.1 20.7 20.7 60.0 20.6 0.1 0.1 6"-SI-419 30 Pipe 0.2 16.2 16.2 60.0 6"-SI-419 38 Elbow 16.1 0.1 16.5 16.6 60.0 q) 16.4 0.1 0.2 6"-SI-419 39 Elbow 0.2 16.6 16.6 60.0 40 Elbow 16.4 0.2 ( 6"-S I- 419 0.2 20.3 20.3 60.0 46 Sockolet 20.0 0.3 6"-SI-419 21.1 20.9 60.0 19.7 1.2 6"-SI-419 77 Tee 6 x 6 x 2 1.4 18.2 17.7 50.4 15.9 2.4 1. 9 6"-RC-418 102 Elbow 2.9 19.9 19.1 50.4 6"-RC-418 100 Pipe 16.3 3.6 2.9 20.0 19.2 50.4 6"-RC-418 99 Loop Nozzle 16.3 3.7 NOTES:
- 1. Refer to Figures VI-13A and VI-13D. ,
- 2. Worst case for either unbroken loop.
t 1 of 1
e TABLE VI-9
,e ECCS STRESS
SUMMARY
Safety Iniection System e Faulted End Disol. Stress (Ksi) Sum of Stress (Ksij, Allow. Stress Unbroken Broken Unbroken Broken Stress Line P t. Component Loop Loop _ 3 Sn (Ksi) Number & Description (Ksi) Loop _ Loop 22.6 0.6 0.7 23.2 23.3 60.0 6"-SI-421 134 Tee 6"x6"x2" .
- 2. 2 17.0 16.8 60.0 2"-S I- 4 61 139 Elbow 14.6 2.4 2"-SI-461 140 Elbow 15.4 2.3 2.1 17.7 17.5 60.0 15.3 2.1 2.0 17.4 17.2 60.0 2"-SI-161 141 Elbow 8.5 1.0 1.1 9.5 9.6 50.4 6"-RC-416 172 Loop Nozzle 1.1 1.2 17.4 17.5 50.4 6"-RC-416 174 Elbow 16.3 16.3 1.1 1.2 17.3 17.4 50.4 6"-RC-416 175 Elbow L1 bow 16.2 1.0 1.1 17.2 17. 3 50.4 6"-RC-416 176 0.9 16.7 16.7 50.4 6"-RC-416 177 Elbow 15.4 0.8 15.3 0.0 0.9 16.1 16.2 50.4 6"-RC-416 178 Elbow NOTES:
Refer to Figures VI-14A and %I-14B. (
- 1. -
- 2. Worst case for either unbroken loop.
i e 1 of 1
TABLE VI-10 - ECCS STRESS
SUMMARY
Satety Iniection synten, Hot Leq End Displ. Stress ( fisi) Sum of Strens (Ksi) Allow. Faulted Stress Co.ponent Stress Unbroken Droken Unbroken Broken Line Pt. Loop Loop 3 SM (X91) Nu-ber No. Descrirtion (Msi) Lcon Loon _ 0.1 0.2 17.3 17.3 60.0 6"-SI-416 7 Sockolet 17.1 1.0 1.3 14.7 15.0 60.0 6"-SI-416 21 Elbow 13.6 1.5 1.7 18.5 18.3 60.0 6"-SI-416 50 Tee 6x6x2 17.1 4.7 5.? 18.2 18.7 51.6 6 "-R C- 4 21 56 Elbow 13.5 13.6 4.5 4.3 18.2 18.0 51.6 6"-RC-421 54 Valve 14.3 1.8 1.4 16.2 15.8 51.6 6"-RC-421 68 Elbow 2.6 2.1 17.4 16.8 51.6 6"-RC-421 67 Elbow 14.6 3.7 3.0 18.6 18.0 51.6 6"-R C- 4 21 66 Elbow 14.9 3.9 3.3 18.9 18.3 51.6 6"-RC-421 65 Loop 15.0 Nozzle 3.4 18.9 18.4 51.6 6"-RC-421 64 Elbow 15.6 3.9 NOTES:
- 1. Refer to Figures VI-15A and VI-15B. -
- 2. Worst case for either unbroken loop.
1 of 1
s
. s
.r TABLE VI-11 ECCS STRESS
SUMMARY
r Safety Iniection Sveter, Cold Leq ( Faulted End Digpl. Streog,fKni) Sun of Streco (Ksi) Allow. Line Pt. Component Stress Unbroken Droken Unbroken Broken Stress , ( Number No. Description (Ksi) Loop Loon Loop Loon 3 Sm (Kni) 6"-RC-417- 52 Loop Nozzle 21.3 0.7 6.6 22.0 27.8 51.6 ( 1502-Q1 6"-RC-417- 53 Elbow 20.2 0.6 7.0 21.2 27.4 51.6 g 1502-Q1 6"-RC-417- 54 Elbow 19.0 0.7 6.8 19.7 25.8 51.6 g 1502-Q1 6"-SI-531- 46 valve 7.9 0.5 4.9 8.5 12.9 60.0 1502-01 6"-SI-531- 44 Elbow 15.9 0.6 5. 5 16.4 21.3 60.0 1502-Q1 6"-SI-531- 43 Elbow 16.2 0.6 5.6 16.8 21.7 60.0 1502-Q1 6"-SI-531- 42 Elbow 15.9 0.5 5.2 16.3 21.1 60.0 1502-Q1 6"-SI-531- 41 Pipe 7. 0 0.5 3.4 7.4 10.4 60.0 1502-01 6"-SI-531- 39 Elbow 13.0 0.8 6.6 13.8 19.0 60.0 1502-Q1 6 "-S I- 531- 30 Elbow 13.3 0.8 7.0 14.0 20.3 60.0 1502-Q1 NOTES: -
- 1. Refer to Figures VI-16A and VI-16B.
- 2. Worst case for either unbroken loop.
1 of 1 . d
e e
)
i
.ss wsK 0 0 0 0 0 0 0 6 6 6 oe( 0 0 0 0 0 0 1 1 1 lr 0 ltm 6 6 6 6 6 6 6 5 5 5 ASS 3
)
i s eE n_ 1 9 6 5 0 4 8 2 7 7 K ko . . . . . ( oo 8 7 8 8 0 2 0 2 9 0 3 9 2 7 2 rL 1 1 1 1 1 s B s . e r t S n e f kp 6 2 5 5 3 5 9 8 9 2 o oo . ro 4 4 4 4 6 6 5 2 1 2 0 2 m bL 1 1 1 1 1 1 1 2 u n SU o e L
)
d i l s n _ o K eg Y C ( ko 0 2 6 5 3 9 8 8 2 8 R oo . . 4 4 4 8 9 8 A . srL 4 4 4 4 M m sB 2 M e e 1 U t r
- S s t 1 I y S V S S f S .
o E E n l L R o e n 1 B T i s e A S t ikp T c Deo 5 5 6 5 6 6 6 4 4 3 S e ro C i dbl 0 0 0 0 0 0 0 1 1 1 C n nn .p E I EU Bo y 7o t 1l e d , - f esl I n a tsi 1 7 0 0 7 9 3 5 5 9 Ve S les . . . k urK 4 3 4 4 5 5 5 1 0 8 do at( 1 1 1 1 1 1 1 2 2 1 nr FS ab n Au 7 n e 1 r o l e ti e z Ih nt e z Vt eo T o i ni N se e cr 2 w w w w w w w w rr pc x o o c o c o p o o ms 6 b b b b b b o b b uo oe x l l l l l l o l l gf CC 6 E E E E E E L E E i Fe s oa
. 1 3 4 5 8 9 0 5 6 7 tc to 5 7 7 7 8 8 9 9 9 9 Pit rt es fr
. eo RW 2 2 2 2 2 2 2 0 0 0 3 3 3 3 3 3 3 2 2 2 5 5 5 5 5 5 5 4 4 4 . . r - - - - - - - - - - : 12 ee I I I I I I I C C C S nb S S S S S S S R R R E ir - - - - - - - - - - T Lu " " " " O N 6 6 6 6" 6" "b 6" 6" 6" 6 N
TABLE VI-13 ECCS STRESS
SUMMARY
Safety Iniection Syster, Cold Leq Faulted End Displ. Stress (Ksi) Sum of Stress (Kai) Allow. Line Pt. Component Stress Unbroken Broken Unbroxen Broken Stress Nu-ber No Description (Knil Loco __ _ Loon _ Loon Loog_ 3 Sm (Kai) 6"-RC-419 99 Elbow 17.0 1.8 18.5 18.8 35.6 51.6 6"-RC-419 98 Elbow 18.0 1.9 20.1 20.0 38.2 51.6 6"-SI-533 89 Pipe 7.4 1.1 12.4 8.5 19.8 60.0 6"-SI-533 Ub Elbow 14.4 0.8 7.6 15.3 22.1 60.0 6*-SI-533 86 Elbow 14.5 0.6 6.3 15.2 20.9 60.0 6 "-S I- 5 3 3 71 Elbow 14.3 0.8 6.0 15.0 20.3 60.0 6"-SI-533 52 Sockclet 15.0 1.8 11.5 16.4 26.6 60.0 6"-SI-533 50 Pipe 9.2 0.9 7.5 10.2 16.7 60.0 2"-SI-453 67 Pipe 5.6 0.2 1.7 5.8 7.2 60.0 6"-SI-453 64 Elbow 10.0 0.2 1.1 10.2 11.1 60.0 6"-RC-419 97 Loop Nozzle 9.7 1.7 18.4 11.5 28.2 51.6 NOTES: ,
- 1. Shown on Figures VI-18A and VI-18B.
- 2. Worst - se for either unbroken loop.
e i 1 of 1
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