ML20235D869
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(5 AFAR c w DESIGN REPORT N O.
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5 Carolina Power & Light Company B R U N S WIC K STEAM ELECTRIC PLANT U NITS 1&2 I
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26 FEBRUARY 1971
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RSEP.1 & 2 I
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CAROLINA POWER AllD LIGHT COMPANT BRUNSVICK STEAM ELECTRIC FLANT UNITS 1 AND 2 DESIGN RDOET NO. 8 i
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SMALL STEAM LINE BREAK i
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ESD-1 & 2 TABLE OF OtMTENTS Section g
g 1.0 DrT30 DUCT!aN 1-1 1.1 Ceneral 1-1 1.2 Accident Evaluation 1-1 3
2.0 IETHOD OF ANALYSIS 2-1 3.0 FWYSICAL D5CRIPTION OF BRYWELL 3-1 3.1 Physical Description of Brywell Steel Liner 3-1 4.0 DESIGN AND ANALYSIS 4-1 1
l 4.1 Idealization of Structure 4-1 l
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4.2 leading Criteria 4-1 4.3 Material Constants 4-2 4.4 Besults of Drywell Analysis 4-2 j
4.4.1 Maximum Liner Stresses 4-3 1
4.4.2 Maximum Reinforcing Stresses 4-3 4.4.3 Soundary conditions 4-3 i
4.4.4 Stress Plots 4-4 f
5.0 LINER SUCKLING 5-1 5.1 Effect of Shear om Bucklias Strength of Liner 5-1 5.2 Buckling Behavior of Liner 5-2 i
i 5.2.1 No Buckling et Liner 5-3 5.2.2 Randon Buckling of Liner 5-4
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5.2.3 Complete Buckling of Liner 5-4 5.3 Behavior of Top Head and Liner Interface 5-5 i
6.0 DtTWELL PENETRATIVE SLEEVIS 6-1 6.1 Materials 6-1 6.2 General 6-1 6.3 Effects of Governing Load 6-1 7.0 RESIDUAL EFFECTS Opi LINER 7-1 8.0 SWMARY AND CONCLUS10ES 8-1 I
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If LIST OF FICURES i
Pisure No.
Tlt1e t]
1 Section Through Primary Containment j
2 Details of Stud Anchor Spacing 3
Drywell Axisymmetric Finite Element Ihdel 4
Enlarged Pinite Element Model of Drywell Emed and Dome 5
Drywell Thermal Gradients and Associated Pressures 6
Drywell Stress Plot "20 (CrackW Concrete) 7 Dryws11 Stress Plot P20 (Uncracked Concrete) 8 Drywell Stress Plot TM0 (Cracked Concrete) j 9
Dryve11 Stress Plot TM0 (Uncracked Concrete) 10 Drywell Stress Plot 1.1D + P20 + T340 IC'** 'd
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11 Drywell Stress Plot 1.1D + P20 + T340 (hene ad Concrete) j 12 Drywell Stress Plot P35 (Cracked concrete) l 13 Drywell Stress Plot P35 (Uncracked comenete) 14 Drywell Stress Plot T320 (Cracked Concrete) j 15 Drywell Stress Plot T320 (Uncracked Comente) 16 Drywell Stress Plot 1.1D + P33 + T33 (Cracked Concrete) j 17 Drywell Stress Plot 1.1D + P35 + T320 I"C"# *d0 "*****I 1
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1.0 Df730 DUCTION I
1.1 General i
j This report discusses the effect on the drywell portion of the primary containment when subjected to an accident which results in a higher temperature and lower pressure than associated with a Loss of Coolant Accident (LDCA), as defined in the Brunswick Steam Electric Plant PSAR. This report includes loading criteria and a discussion and graphical presentation of the analytic results.
i Particular attention has been given to the of fect on the drywell steel liner during this accident. 1he LOCA (Design Basis Accident) results in more severe loadings on the supression chamber than this acci' at and, therefore, governs the design of the suppression chamber. The drywell concrete will not be subjected to prolonged thermal loadings so that the thermal gradients discussed below realistically describe the thermal effects that the concrete is exposed to.
1.2 Accident Evaluation An incident occurred at a nuclear power plant in which primary steam leaked into the drywell through cocked open safety valves. Analysis l
of the containment response during the incident has resulted in the estab-lishment of two additional design conditions. These two additional design i
conditions have been analyzed in this design report and are included in the table below:
1 Tempe ra ture Pressure i
Cane 1 Recirculation line break 281 F 46 psig 1-1 l
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58EP-1 & 2
_ Tempera ture Pressure Case 2 Hsin store line break from 1050 psi 320' F 35 psig 1
Case 3 Sas11 steam line break from 500 pst 340 F 20 psig The Case 1 conditions are the original design conditions, and correspond to a guillotine break of a major recirculation line taking place with the reactor operating at full power. The Case 2 conditions are a new set of conditions determined for a guillotine break of a primary steam line 5
taking place with the reactor operating at full power. The Case 3 conditices would occur for a special set of reactor conditions is which a steam blow-down would taka place with the reactor pressure held constant at 500 pois, and the leakage taking place over a relatively long period of time.
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58EP-1 6 2 2.0 METHOD OF ANALYSIS The method of analysis used is a finite element direct stiffness method for axisymmetric solids of revolution. The program used in the analysis was developed by Dr. E. L. Wilson of the thiversity of California, i
Berkley. BSEP Design Report No. 7 (DR-7), Section 3.0 dated December 31, 1970 includes details of this method of analysis.
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_ PHYSICAL DESCRIPTION OF DRWELL A physical description of the drywell is included in DR 7
. Section 4.1.
Figure I shows a cross section of the primary containment The physical material properties for the various components of the containunt structure are:
1.
Liner Steel ASIM-A516 GR. 60 to A-300 F =28 ksi at T=340 F 1
2.
Drywell Head Steel ASTH-A516 CR. 70 to A-300 F =26 kai at T=340 F 1
3.
Reinforcing Steel 714 and #18 bars
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ASTM-615 - edified to CR. 50 F =50 kai 1
- 6 to #11 bars l
ASTH-A615 CR. 60 F =60 kai y
$5 and smaller bars ASTM-A615 CR. 40 F =40 kai
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1 4.
Concrete j
Minimum 28-day compressive stress = 3000 pst f
Capacity reduction factors used are those stated in the BSEP PSAR:
1.
6 =.9 for flerure and tension 2.
0 =.85 for diagonal tension, bond, archorage 1
3.1 Physical Description of Dryvell Steel Lines l
i The portion of the liner backed by concrete is ASTM-A516 GR l
. 60 and the top head is constructed of ASTM-A516 GR. 70 saterial l
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88tP.1 6 2 Typically the dryvell liner is 5/16" plate, but at intersections of cones and cylinders which form the drywell, 3/4" plate is used. In addition, around each penetration, there is a reinforcing pad, 5/8" minimum thickness.
l To insure elastic stability, the liner is anchored to the concrete by 1/2" diameter x 8" long Nelson studs spaced 12" horizontally and vertically.
Around each penetration and at each seas, a closer spaced pattern is used.
The stud patterns are shown in Figure 2.
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SgEP-1 6 2 4.0
_DESIGl AND ANALYSIS 4.1 Idealization of the Structure the same idealization of the structure is used as described in DE-7, Section 4.2.
Figure 3 shows a cross section of the axisymmetric finite element model. Figure A shows an enlarged model of the top head. A table of equivalent plate thiciusesses associated with all idealised liner reinforcing eleasats is also included in Pigure 4.
4.2 Loadina Criteria The following load equations were used in the analysis:
1.
U = (1.01 1) D + P20 + T340 A
2.
U = (1.0 i.1) D + P35 + T320 B
Where U = Ultissata required load capacity of the structure D = Dead load of structure faciuding fixed equipment P20 = Pressure of 20 psig
- Associated with a small steam 7340 gerature f 4 F
line bred
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P33 = Pressure of 35 psig
" Associated with a large steam line T
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break 320 A discussion of these pressures end temperatures is found in Sec tion 1.2.
Figure 5 shows thermal gradients and associated pressures used for the analysis.
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F 4.3 Msterial Constants The material constants used in the analysis are listed below:
_ Conc re te : (uneracked)
Concrete: (cracked)
E,. and/or E,. and/or E, = 3.0 x 10 psi E,. and/or E,. and/or E, = 0.0 psi w
=.15 v
= 0.0 6
C
= 1.3 x 10 pet G
= 1. 3 x 10 psi i
Liner 0
E, and Eg = 27.4 x 10 psi E,. = 0.0 (not applicable. *n = 0.0)
- se =.30 6
C
= 10.5 x 10 p,g Reinforcing (Hoop)
Reinforcing (Meridional) 6 6
E,.= 29.0 x 10 psi E,
= 29.0 x 10 p,g E, = E, = 0.0 E,
=E
= 0.0 9
w = 0.0 w
= 0.0 C = 0.0 C
= 0.0 Head Bolts 6
E,
= 29.7 x 10 p,g E
= E, = 0.0 v
= 0.0 G
= 0.0 4.4 Results of Drywell Analysis A series of 12 computer runs were made to maximize liner and rein-forcing stresses.
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4.4.1 Maximum Liner Stresses The liner stresses were maximised by considering that no concrete i
in the structure was cracked.
This saintaised the free thermal growth of the liner and consequently maximised its compressive stresses.
This is a very conservative asseption because the structure is subjected to thermal bending moments that cau.se stresses in the concrete which exceed its tensile capacity, i.e., the concrete would crack.
This was verified by the analysis.
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As a result of the concrete cracking the structure would grow l
more than assumed, thus relieving the thernst loads aed-reducing the com-j pressive stress in the liner.
4.4.2 Maximum Reinforcing Stresses Fully cracked concrete maximised reinforcing stresses because the reinforcing alone resists the entire load with no assistance from the concrete.
As will be noted in Section 4.4.4, af ter an examination of re-inforcing stresses using 0.9D in the load equations, it was determined that the 1.0CA loads governed the design of the reinforcing.
Therefore, in order to maximize the temperature induced compressive stresses in the liner,all final computer runs were made using 1.1D in the load equations for the low-pressure, high-temperature accident conditions.
4.4.3
_B_oundary Conditions To maximize liner and reinforcing stresses, the drywell pedestal was assumed to be restrained vertically and radially. The experience gained
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4-3 1
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i httP.1&2 in analyzing the structure as reported in Design asport No. 7 was used in determining which boundary conditices anximized stresses.
4.4.4 Drywell Stress Plots The results of the analysis are presented graphically in the lorm of stress plots for the liner and reinforcing. Positive and negative stresses were plotted respectively to the right and lef t of the centroidal axis of the structure.
Stresses were plotted rt 14 representative sections in the drywell wall, six sections in the drvwell pedestal, and three sectiona ilF in the drvwell head.
i Stress components were plotted for the meridional and hoop stress in the liner and for each layer of maridional and hoop reinforcing.
Pressure lead Stresses 1,
Figures 6, 7,12 and 13 show the effect se the structure of 20 peig and 35 pois pressure loading. hto sets of stresses were plotted for each pressure loading case, one for the uncracked concrete case and one for the cracked concrete case.
Thermal Load St ressen Figures 8, 9,14, and 15 show the effect of the 320 F and 340 F thermal loads on the structure. Two sets of stresses were plotted for each thermal loading case, one for the erscked concrete case and one for the uncracked concrete case.
Combined leads Figures 10,11,16, and 17 show the ef fec: of the combined loads on the structure according to the following conditionn:
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. l risure 10 1.1D + P20 + T340 (Concrete Crachd)
Finure 11 1.1D + P20 + T340 Finure 16 1.1D + P35 + T320 (concrete Crachd)
Flaure 17 1.1D + P35 + T320 (Concrete Uncracked)
The conclusions based on a comparison of the above stress plots and those presented in DR-7 are as follows:
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1.
Design of reinforcoasnt is governed by the 10CA combined loads (DR-7). This is because the higher pressures and only slightly reduced temperatures associated with the LDCA combined loads result in higher reinforcing stresses than the combined loads considsred in this report.
2.
The two load conditions considsred in this report result in higher compressive stresses is the liner, and consequently a more severe condition for the liner than the 14CA load conditions (DR-7).
3.
The load condition of P = 20 peig and T = 340' F results in higher compressive liner stresses than the P = 35 peig and T = 320 F loading.
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BSEP-1 6 2 5.0 LINER BUCKLING This section of the report will be erected to a detailed investiga-tion of the liner when it is subjected to a cen61ned load of 1.1D + P 20 + T340 (governing load condition for liner.)
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The primary purpose of the liner is en function as a lenktight mem-brane under the specified loading conditions.
The liner is anchored to the concrete with stude as described in Section 3.1.
References 1, 2. 3. 4 and 5 were med for the design of the an-chorage system.
The use of 1/2" diameter x 8" long Nelson studs spaced 12" vertical-ly and horizontally on the 5/16" liner plate results in a critical bucklina stress, associated with equal bi-axial compressm, of 40.2 ksi.
In order to minimise the critical buckling stress, the largest diarnecer of the drywell, i.e., 64'-0", was used in the calculation of the critical buckling stress.
The critical buckling stress of 40.2 kai was used to conservatively estimate the buckling characteristics of the entire line.
Various imperfections in the liner plm:c, along with stud misalign-ment, were considered.
Since the yield strengtt of the liner is 28 ksi at a temperature of 340 F. the anchorage system insares that elastic buckling will not occur.
5.1 Effect of Shear on Buckling Stre gth ed Liner The effect of shear on reducing the cr.tical buckling stress has been considered.
The amount of seismic shear prmortioned between the liner and concrete shell depends upon the relative stiffness of each.
Conserva ti ve1v 5-1
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88EP.1 & 2 assuming only the stif fness of the reinforcing in the concrete shell and the stiffness of the limer, approximately 30% of the seismic shear could be trans-mitted into the liner. This results is a liner stress of 5.5 kei.
The limiting case, governing the contribution of shear otress to axial stress in the liner, to determine the critical buckling stress shows the liner car de of carrving 50% of the seismic shear stress before the critical buckling stress is reduced.
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i entire seismic shear load.
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It is concluded that shear leads will not reduce the critical buck-
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ling stress froa its calculated value of 40.2 kat.
i 5.2 Bucklina anhavior of, Liner The basic design concept for the liner, utilising stud anchorage ductility, is that under the specified load conditions the stude will not fail under shear, tension or bending loads causing the stud connection to fail or f
i tear the liner plate.
Tests have shown that to develop the ultimate capacity of a 1/2" diameter stud, an embedment length of approximately 5" is required. As an additional measure of conservatism, the studs used are embedded 8", thus insuring that thev v.11 develop their ultimate capacity.
To insure that the studs provide lateral support for the liner, they must supply a tensile capacity of 1% of the maximum compressive load on the liner For the governing load case, using conservatively 2% of the maximum 3
5-2
.R8EF-1 & 2 i
compressive liner load, the required tensile capacity of a stud is 2.2 kips.
I Since 85% of the ultimate capacity of 1/2" diameter studs is 11.8 kips, this design insures the studs provide lateral support for the liner. This criterion is also adequate if a stud is not installed and the adjacent stud i
has to carry additional loads. The pressure associated with the governing load case provides an additional margin to insure lateral support for the liner.
j The liner strains multiplied by the stud spacing result in deflec-
)i tions which, under all loading conditions, must be less than the stud capacity.
f Tes ts show that a 1/2" diameter stud will deflect (slip).069" at 85% of i
its ultimate load. Based on a stud specing of 12" vertically and horizontally, this criterion results in a maximum liner strain of.00813 inch / inch before i
J the studs will reach 85% of their ultimate load. As will be noted below, the liner strains are below this maximus value under all load conditiono, and thus failure will not result.
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i The effect of the load equations on the buckling behavior of the
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vestigated:
1.
No buckling of liner.
T 2.
Random buckling of liner.
3.
Complete buckling of itner.
5.2.1 No Buckling of Liner The stud spacing insures that the liner will yield before it buckles.
The critical buckling stress is 40.2 kai. This results in a liner strain of l
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.00147 inch / inch which is significantly below the critical strain of.00813 f
inch / inch.
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l It is concluded that for a loading condition which results in a i
liner vield and the onset of buckling, the stud pattern provides an adequate margin of safety against stud failure.
5.2.2 Randon Buckling of 1.iner I
i Figure 10 (load equation 1.1D + P20.+ T340; Cracked Concrete) shows j
a local area of buckling at the large cylinder and lower cone sections of the drywell. The associated liner stress is 44 kai which results in a liner strain below the critical buckling strain. This buckling will not result in stud fa!!ure and consequently no lou-of-function of the line vi i result.
i 5.2.3 Complete Bucklina of Liner j
Figure 11 (load equation 1.1D + P20 + TM0; Unctacked Concrete) shows stresses which result in general liner buckling. A maximum stress of 64.3 ksi j
occurs in the lower section of the upper cone. This results is a liner strain fi
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of.0035 inch / inch significantly below the critical buckling strain of.00813 i
inch / inch.
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Even considering zero pressure and 340 F temperature (Figure 9, T340; Uncracked Concrete) the liner strains will be below the critical buckling s t rain.
As noted in Section 4.4.1, this analysis assuming uncracked concrete is very conservative because, due to the thermal bending moments, the concrete does crack; this, if taken into consideration, would relieve $ne liner. Main-taining liner strain levels significantly below those that would cause stud failure insures that an unbuttoning failure of the liner will not occur.
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t 5.3 Behavior of Top Head and Liner Interface The intersection of the top head and the concrete drvwell were in-vestigated for ef fects of the governing load equation.
The top head is constructed of steel and, under the predominantly thermal effect of the load, experiences approximately free thermal growth, i. e., ve ry low stress levels.
Because the top head is less stiff than the concrete dryvell, the top section of concrete grows more than the concrete section below the inter-section.
This results in lower stresses in the top section of liner plate, than liner stresses below the intersection, as noted on Figures 10 and 11.
The dryvell meridional reinforcement f a cadwelded to a 1-1/2" cap plats which is, in turn, connected to the top head. In addition the liner is thickened to 1-1/2" at the top of ths drywell and connected to the top head.
This insures continuity at the intersection of the top head and concrete drywell.
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1 RSEP-1 & 2 6.0 DRWELL PENETRATION SLEE1ES 6.1 MaterisIn Penetration sleeves 12" diameter and smaller are ASIN-A333 Cr.1
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l material, while penetration sleeves larger that 12" diameter are ASTM-A516 Cr. 60 natorial.
6.2 General Continuity between the linar, reinforcing pads and penetration sleeves to insure that loss of function does not occur, is maintained by an-choring the penetration assemblies to the concrete with a series of flat bar ard bearing plate anchorages.
In addition, at each reinforcing pad the stud spacing is reduced from the typical spacing of 12" x 12" used en the liner to accomodate the stress concentration effect associ.ted with the penetration sleeve.
Size and thickness of reinforcement pade is deterair.ed by replace-ment of area design procedure.
Load transfer around penetrations is maintained by establishing con-l tinuity of the main reinforcement. This is accomplished by bending or splav-ing the main reinforcement around the penetration sleeve. Supplemental rein-forcement is added where required.
A detailed analysis of the large openings will be reported in a future RSEP Design Report to be submitted to the AEC in the near future.
6.3 Et'fects of Coverning Load Becauce of the predominant thermal effects of the governing load, the 6-1
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BSD.1 6 2 thermal stresses at the discontinuity between the penetration sleeve and the liner were investigated. Also, consewativelv assuming the concrete uncracked, the stresnes due to restraining free thermal growth were determined.
Timoshenko & Coodier ( I was used to determine the stress concentra-tion ef fects around the openings due to the penetration sleeves.
Because the vent pipe penetration ased the personnel lock penetration will be the only sleeves subjected to the governing load equation, their res-ponse was determined.
To reduce the thermal stress and to eliminate buckling, a laver of compressible material will bw icatalled aromand the vent line penetrations and personnel lock penetration. This material will permit free thermal growth of the sleeves. Tb6 thickness of this material will be based on the calculated f ree thermal growth of the penetration with allowance for construction loads.
Under the governing load conditions with uncracked concrete. It was determined that the penetration sleeves will not buckle, and thua wl!!
remain functional.
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BSEF-1 & 2 7.0 3ESIDUAL EFFECTS ON LINER The ability to return the plant to operation should a high-tempera-ture low-pressure accident occur is of paramount importance. One of the asjor structural requirements is that the liner maintain its function as a leaktight memb rane.
This report has shown that under the governing load equation (1.1D +
P20 + TMO), e me buckling of the liner may take place; however, in no in-l stance is liner leaktightness compromised.
I While the pressure associated with this accident wi!! reduce the de-formation of the liner, some small perusnent set coald possibly remain. This would reduce the 11nar stiffness in some areas, but, as has been shown in 3t-7, 4
the reinforcing can carry LOCA loads, even if liner stif fness is completely eliminated.
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8.0 Sl4 MARY AND CONCLUSIONS I
The drywell structure, in particular the liner and penetration j
sleeves, has been investigated for response to the temperature and pressures t
associated with a small and large steam line break. These postulated accidents i
I result in higher temperatures and lower pressures than the LOCA.
l It has been determined that the design of the reinforcing is govern-ed by the Design Basis LDCA loads discussed in DR-7, while the design of the i
liner, studs and stud spacing is governed by the loads discussed ir. This re-
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1 Conservatism is introduced in the analysis by assuming uncracked Even with this added conservatism, it is shown that the liner and concrer e.
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the stud connections tc the liner will eat fail, and consequently, the liner will maintain its leakti htness. Exzmining the results containing the more F
rastistic aesumption of cracked cone: rete indicates a small section of the liner will buckle but esill function satisfactortiv.
I tened on thir analysis, it is concluded that the drvvell structure will maintain its integrity during and after this cecident, can be safeli returned to operation, and subsequently respond satisfactorily to a Design Easta LOCA or to additional small steam line break accidents.
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REFERENCES
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1.
Theory of Elastic Stability by Timoshenko & Gere. McGraw-Hill,1961, 2nd Edition.
2.
Effect of Imperfections on Bucklina of Thin Cylinders & Columns Under Arial Compression by Donnell & Wan, Journal of Applied Mechanics (ASME),
March 1950.
3.
The Stabilization of the Steel 1.iner of a Prestressed Caecrete Pressure vessel by Chan & McMin, Nuclear Engineering & Design 3,1966.
l
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4 Theory of Plates & Shells by Tinochenko & Woinowsky-Kriegre McGraw-Hill,
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1959, 2nd Edition.
5.
The Effect of cene_ral Imperfections on the Bucklina er Cylindrical snells by Ath:s & Babcock, Journal of Applied Mechanics (ASME) March 1969.
6.
Containment Desian Report, BSEP DR-7, Carolirea Power & Light Co.,
December 31, 1970.
7.
Engineering Desian Data for Nelson Concrete Anchors, Nelson Stud Welding Co., May, 1968.
8.
Theory of Elasticity by Tinoshenko & Coodier, McGraw-Hill,1951, 2nd Edition.
9.
Plastic Desian of Steel Frames by Beedle, John Wiley & Sons,1961.
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