ML20091D767
| ML20091D767 | |
| Person / Time | |
|---|---|
| Site: | Fort Saint Vrain  |
| Issue date: | 10/01/1991 |
| From: | PUBLIC SERVICE CO. OF COLORADO |
| To: | |
| Shared Package | |
| ML20091D758 | List: |
| References | |
| EE-46-0007, EE-46-0007-R-B, EE-46-7, EE-46-7-R-B, NUDOCS 9110290014 | |
| Download: ML20091D767 (22) | |
Text
- .
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Primary Evaluation Number EE-46-0007 REV.B ENGINEERING EVALUATION OF PRESTRESSED
.l CONCRETE REACTOR VESSEL AND CORE SUPPORT FLOOR STRUCTURES FOR A i
PROPOSED SYSTEM 46 TEMPERATURE CH ANGE I
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d Prepared by:
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- I/21/9/
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' DATE Verified by:
M Mc/u 27.Ti
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_f. BATE Approved by:
abYa.$b ph /9 i DEPARTMENT MANAGER DATE 9110290014 911011 PDR. ADOCK 050001o7 P-ppg
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1.0 PURPOSE This Engineering Evaluation (EE) will determine the feasibility of allowing a decrease in the operating temperatures of System 46 cooling water provided to the Prestressed Concrete Reactor Vessel (PCRV) and the Core Support Floor (CSF).
Specifically, this EE will investigate the possibility of allowing the minimum average of the inlet and outlet cooling water temperature to decrease from the 100*F specified in '.CO 4.2.15 (e) to 85'F. This change is desirable because of present System 46 sperating considerations.
2.0
SUMMARY
Various areas of the PCRV and CSF judged to be the most critical were investigated to determine the effects of the proposed System 46 operating temperature change. Areas investigated were the PCRV liner, the PCRV liner anchor studs, the PCRV concrete, the PCRV penetrations, the PCRV tendons, the PCRV reinforcing rods, the CSF liner, the CSF concrete, and the CSF reinforcing rods.
The analyses performed to evaluate these various areas conservatively assumed that the proposed 1S*F temperature change occurs instantly in the item being investigated. This assumption provides conservative results as it creates the maximum differential thermal movement between adjacent items which are physically bonded, such as for example, the PCRV liner and concrete which are connected together by anchor studs embedded in the concrete and welded to the liner. The assumption of instantaneous temperature change is very conservative as the thermal masses involved are very large and temperature variations of the System 46 cooling water tend to occur slowly.
The stress levels in the various items due to the 1S*F temperature change were found to be relatively low even when based upon very conservative assumptions.
This is due to the fact that the proposed temperature variation of 15'F is very low and does not result in high levels of stress in even completely restrained structures.
The total stress levels in the PCRV and CSF structures due to the proposed temperature change and other previously analyzed loads were found to be acceptable. A summary of the stress levels in the areas investigated follows.
SUMMARY
OF PCRV AND CSF STRESSES DUE TO 1S*F SYSTEM 46 TEMPERATURE CHANGE PCRV LINER (PSI)
(PSI)
LOWER FLOOR COLUMN ANCHOR-59 4
O LOCAL STRESS INTENSITY IN LINER JUNCTION WITH BOTTOM ACCESS PENETRATION LINER STRESS 24.039 65,700 INTENSITY IN CAVITY LINER 1
l I
l l
l PCRV LINER MAXIML*M SHEAR ALLOWABLE SHEAR ANCHOR 5TUD5 LOAD LOAD 18,605 LB.
32.500 LB.
PCRV CONCRETE MAXIMUM TEN 51LE ALLOWABLE TENSILE STRESS (PSI)
STRESS (PSI)
(See Note 1) 220 233 PCRV PENETRAT10NS (PSI)
(PSI)
BOTTOM ACCESS PRIMARY 40,534 69,300 PENETRATION SECONDARY 41,318 52,500 l4LIUM CIRCULATOR PRIMARY 36,024 52,500 PENETRATION SECONDARY 36,722 52,500 REFUEbNG PRIMARY 19,419 52,500 PENETRATION SECONDARY 34,079 52,500 TOP ACCESS PRIMARY 41,879 69,300 PENETRATIONS SECONDARY 30.589 52,500 The stress levels in the tendons do not increase PCRV TENDONS due to this proposed temperature change.
The stress levels in the reinforcing rods do not PCRV REINFORONG RODS increase due to the proposed temperature change.
S STRESS AROWABLE CSF LINER (PSI)
(PSI)
Liner masimum tensile stress 6,639 23,100 Liner support columnjunction 62,539 69,300 2
CSF CONCRETE MAXIMUM TEN 51LE Al,LOWABLE TEN 51LE
$1REs5 (P5!)
STRF.55 (PSI)
(See Note 1) 382 233 The stress levels in the reinforcing rods do not CSF REINFORONG RODS increase due to this proposed temperature change.
NOTE 1:
The 233 PSI allowable is for unreinforced concrete. The PCRV and CSF are constructed with bonded reinforcing steel and this 1
low tensile stress is not significant.
The effects of the lower temperature upon the ability of the PCRV liner materials to resist fracture were investigated. The PCRV liner is subjected to
)
neutron irradiation with the top head liner receiving the highest dose. This
)
irradiation tends to increase the Nil Ductility Transition (NDT) temperature j
of the steelliner. LCO 4.2.15 specifies a minimum liner temperature of 100*F so as to maintain a 60*F margin above the NDT at the plant end of life after 30 years of operation.
The reactor was permanently shut down on August 18, 1989 having accumulated 890 Effective Full Power Days, which represents approximately one tenth of the design lifetime and which corresponds to a maximum integrated neutron dose of 2.4E E17 n/cm'. This neutron exposure would cause a shift in the NDT temperatu-a of approximately one tenth of the experimentally determined NDT temperature shifts.
The fracture transition elastic (FTE) temperature is approximately NDT +
60*F and this is the lowest allowable temperature. The end-of life FTEs were calculated to be 10*F for the liner material anci -2*? for the liner weldment material. These FTE temperatures are well below the proposed value of 85'F and it was concluded that operation at 85'F is acceptable for the liner materials.
~
3.0 SCOPE The scope of this EE includes the structural evaluation of the PCRV and CSF for loads imposed by the proposed temperature change. Additionally, the scope of this EE includes the fracture mechanics evaluation of the PCRV liner for the proposed decrease in liner temperature.
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i Figure 4.1 PCRV General ConGguration Ril01-100C 1
4 0
4.0 BACKGROUND
The Prestressed Concrete Reactor Vessel (PCRV) is the structure that contains the reactor core and the entire primary coolant system. It functions as the primary coolant pressure boundary. The general configuration of the PCRV is shown in Fig ure 4.1.
The PCRV is constructed of high strength concrete reinforced with bonded reinforcing steel and prestressed by means of the post tension method with steel tendons. The main cavity of the PCRV is an upright circular cylinder 31 ft. in diameter and 75 ft. high. The exterior vertical surface approximates a hexagonal prism with vertical pilasters at each corner which accommodate anchors for circumferential prestressing tendons. Both PCRV heads are flat. The external PCRV dimensions are 49 ft. across the flats of the hexagon and 106 ft. high. The wall and heads are 9 ft. and 15 ft. 6 in, thick, respectively.
The concrete walls and heads are constructw around a 3/4 inch thick carbon steel liner which provides a leak tight membrane' for containing the primary coolant within the PCRV cavity. This liner is anchored to the concrete by means of studs welded to the liner and embedded in the conctete.
Prestressing tendons, located in conduits ernbedded in the concrete, are used to prestress the entire structure prior to pressurization. Prestressing forces are oriented such that they oppose internal pressure.
The concrete is reinforced with bonded reinforcing steel which provides the added tensile strength needed where discontinuities cause unavoidable secondary tensile strain, distributes and minimizes width and depth of the minor cracking caused by concrete shrinkage and tensile strain from thermal gradients, and resists localized high compressive and shear stresses that would otherwise overload the concrete.
The concrete, reinforcing steel, prestressing system, and steel liner function as a composite structure.
The temperature of the PCRV concrete is controlled by means of insulation mounted on the inside surface of the liner, and cooling tubes welded to the concrete side of the liner, in general the cooling tubes are spaced on approximately 7.5 in, centers.
However, additional cooling system capacity is provided in the cylindrical section of the PCRV liner from just below the core support floor to just above the top of the core barrel. Here the spacing is 3.75 in. Maximum spacing of cooling tubes on the top head liner is 3 in. Cooling tubes are arranged so that alternate tubes are connected to redundant supply and return headers.
The whole of the internal surface of the liner is covered by the thermal barrier which uses Kaowoolinsulation, a ceramic fiber blanket material of high chemical purity. The blankets are compressed against the PCRV liner by 1/4 in carbon steel 4
cover plates and studs that are attached to the liner. This causes Kaowool to conform very closely to liner surface irregularities and provides an effective seal against helium flowing in those irregularities, thus preventing heat transfer to the liner by forced or natural convection.
Due to operational considerations it is desirable to reduce the operating temperature of System 46 which provides cooling water for the PCRV and the CSF.
The Technical Specifications LCO 4.2.15 establishes the following limits for System 46 operating temperatures:
a)
The maximum temperature difference between the outlet water temperature of the PCRV cooling water system, and the PCRV external concrete surface temperature, averaged over 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />, shall not exceed 50*F.
b)
The maximum outlet water temperature of the PCRV cooling water system shall not exceed 120*F.
c)
The maximum temperature difference between the outlet water temperature and the inlet water temperature of the PCRV cooling water system shall not exceed 20*F.
d)
The maximum rate of change of the PCRV concrete temperature shall not exceed 14'F per week, as indicated by the weekly average outlet water temperature of the PCRV cooling water system.
e)
The minimum average of the inlet and outlet cooling water temperatures shall be greater than or equal to 100*F.
This Engineering Evaluation willinvestigate the effects of lowering the average of the inlet and outlet temperatures from 100*F (condition "e" above) to 85'F. This would provide the additional working margin desired for System 46 operating considerations. Additionally, the maximum outlet water temperature will be lowered from 120'F to 105'F (condition "b" above). No changes are proposed for the other conditions (a, c and d above).
The average of the inlet and outlet cooling water temperatures is presently approximately 107 F. The PCRV internal pressure is limited to less than 1 psig by the requirements of LCO 4.7.1, but a maximum credible pressure of 5 psig was assumed for this analysis per section 9.5.5 of Ref. 8.1.
It is of interest to note that the proposed System 46 operating temperature change can be viewed as being very close to a previously analyzed condition.
Consider Table 5.3 2 of Ref. 8.1 Here are reported some of the stress results for a variety of loading conditions which cover the life of the plant. Of interest are the following cases: 1) the PCRV prestressed, unheated (70*F), and at atmospheric pressure 2) the PCRV prestressed, at design temperature, and at atmospheric 5
pressute 3) the PCRV at the end of operation, at design temperature, and at reference pressure, and 4) the PCRV at the end of operation, unheated, and at atmospheric pressure. These cases contain the history of the PCRV as it was prestressed, heated, pressurized, depressurized, and allowed to cool to atmospheric temperature.
The proposed loading case is very close to case 4 mentioned above with the primary differences being that the PCRV will not have cooled completely down to the case 4 level and the f act that it is possible that the PCRV could be pressurized to a low level (5 PSI per Ref. 8.1) versus atmospheric pressure of case 4. The PCRV also has not been pressurized for as long as was assumed in case 4 resulting in less of a loss of prestress in the concrete and tendons due to concrete creep than was predicted in the case 4 analysis.
5.0 APPROACH The PCRV and CSF structures were previously analyzed for a variety of load cases and were found to be structorally adequate as reported in Ref. 8.1. The evaluation of this EE builds upon the results of the previous analyses.
Stresses which could arise due to the proposed temperature change are conservatively calculated for various PCRV and CSF components which are considered to be the most critical. In general the assumption is made that the temperature change is instantaneous, resulting in the maximum differential thermal movements and corresponding thermal stresses. These stresses are added to previously calculated stresses in a conservative manner and the total stress is compared to the allowable stress.
The PCRV liner is subjected to neutron irradiation during plant operation.
Calculations are performed to determine the Nil Ductility Temperature of the liner materials based upon actual plant operating history. The new NDT values are compared to the proposed reduced liner temperature of 85'F to ensure that a 60*F margin exists between the NDT and the minimum liner tempvature.
6.0 EVALUATION OF PCRV AND CSF STRUCTURES 6.1 PCRV Liner The whole of the internal surface of the PCRV, that is exposed to primary coolant,is covered by a continuous 3/4 inch thick carbon steelliner. Welded studs are attached to the outside surface of the liner on a 7-1/2 in. x 7-1/2 in. pitch and are embedded in the PCRV concrete.The liner, studs, and PCRV concrete act as a composite structure and the liner follows the major concrete strains. (Ref. 8.1, Section 5.7)
During the proposed System 46 operating temperature change it is possible that the liner would be subjected to some thermally induced stresses. These stresses would occur due to the differential thermal growth of the liner 6
s 4
relative to the PCRV concrete which would be restrained by the anchor bolts.
For a limiting case the liner will be assumed to instantly cool off the entire proposed decrease of 15*F from the conditions of I.CO 4.2.15 (e) while the concrete remains at the initial temperature. The liner will t'.ien be conservatively treated as a uniform flat plate held at the edges subjected to a uniform temperature decrease of 15'F. Such a plate would attempt to contract and, being restrained, would develop tensile stresses. The worst case magnitude of these tensile stresses can be calculated as follows:
ATaE (Ref. 8.3, Pg. 374, Article 88, Case 2) o=
ST = 15'F a = Coefficientofthermalerpansion
= 6.5 x 10 ~' Inlin
- F E = modulusofelasticity 29 x 10' PSI (Ref. 8.4, Pg. 4) p = Poissons Ratio = 0.3 15*F 6.5 x 10 ~' In/In
- F 29 x 10 L6/In 6
2 a=
1 -.3
?
o = 4,039 PSI (Tension) b The liner is in general in a state of compressive stress for all plant operating conditions due to the prestress of the PCRV tendons. The level of compressive stress in the liner is increased when, as in the case being considered, the internal pressure of the PCRV is decreased (Refer to Ref. 8.1,
. Table 5.3 2 for various load case liner stresses). The liner tensile stresses
- calculated above will tend to negate a portion of these compressive stresses which exist in the bulk of the liner. This fact will be conservatively neglected and the calculated liner tensile stresses will be added directly to the liner stresses as reported in Table 5.71 of Ref. 8.1. It is also noted that the calculated thermal stress is secondary in nature but will be conservatively added to certain primary stresses reported in Table 5.7-1.
7
Arca Stress Intensity (PSI)
Stress Allowable (PSI)
Lower floor column anchor -local stress 18,500 + 4,039 = 22,539 34,600 intensity in liner junction with bottom access penetration 20,000 + 4,039 = 24,039 65,700 liner - stress intensity in cavity liner junction with core support column - stress 24,800 + 4,039 = 28,839 69,300 intensity in cavity liner junction with top access penetration -
55,900 + 4,039 = 59,939 65,700 liner - stress intensity in cavity liner it is concluded that the PCRV liner is structurally adequate for loads imposed by the proposed System 46 operating temperature decrease.
6.2 PCRV Liner Anchor Studs The liner anchor stss are attached to the autside surface of the liner on a 7f inch circe:.nerential by 7f inch axial pitch (Ref. 8.1, Section 5.7.1). The anchor studs act as shear anchors which force the liner an i PCRV concrete to act as a composite structure.
The proposed reduction in the System 46 operating temperatures could potentially increase the loads acting upon the anchor studs by two mechanisms. The first mechanism is due to the fact that when the liner cools off it will shrink circumferentially and attempt to move in the inward radial direction away from the PCRV concrete. This radial movement will be resisted by the anchor studs resulting in a tensile axial force in the studs.
The second mechanism is due to axial shrinkage of the liner which could renit in shear forces acting upon the studs as discussed below.
The worst case scenario for both of the mechanisms is that the liner instantly cools off the full postulated 15'F. This will maximize the differential thermal movement between the PCRV concrete and liner and consequently maximize the forces acting upon the anchor studs which resist these differential movements.
The axial tensile forces acting upon the anchor studs due to circumferential thermal shrinkage of the liner will be conservatively calculated as follows:
Since the studs are located on a 7t x 7t inch pitch,in order to calculate the loads on a circumferential row of anchor studs the liner can be treated as a 1
8
- - - - - - - - - - - - ~ - - - - - -
~
7t" high section of the liner (ring) which is assumed to be subjected to a 15'F temperature drop.
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The unrearained thermal contraction of this ring is then:
6 = a AT D (Ref. 8.5, Pg. 82) where 6 = nniculatedchangein diameter a = cuelficientofthermalexpansion
= 6.5 x 10-8in/in
- F Re 8.4, Pg. 4 ST = 15'F (Anumed)
D = LinerDiameter = 31'= 31 x 12 = 372-(Ref. 8.1, Section 5.1) 6 = (6.5 x 10-8)(l5)(372) = 0.0303" The radial pressure acting upon the ring to produce this deflection will now be calculated. The pressure will then be multiplied by the effective contributory area of 7t x 7f inches per stud. This will give the axial force acting on each stud.
RPn (Ref. 8.3, Pg. 298. Case 1)
RadialDisplacement Es solving for P:
9
i SKt i
P= n' where R
Radsus = (31 X 12)(1/2) = 186" (Ref. 8.1, Section 5.1)
=
P PressureiTo be calculated)
=
E Modulus ofElasticity
=
29 x 106 PSI (Ref. 8.4, Vg. 4)
=
3/4 "
(Ref. 8.1, Section 5.7.1) t
=
RadialDisplacement 1/2 diameter change calculated above
=
(1/2)(.03G3)" = 0.0181"
=
f0.75 0.0181 29x10 P=
= 11.4 PSI 2
186 Radialload! Stud
= Area x Pressure = (7)~ x 7)")(I1.4 PSI) = 641 Lo.
The resulting stress in the 3/4 inch diameter studs froro this load is 2
o = P/A = 641 Lb / 1/4 n
3/4
= 1,4 50 PSI-For the liner stud spacing and embedment lengths the stud strength will be fully developed and the stud will fail before the concrete. The above calculated stud axial stiess is very low and is evaluated as being acceptable. No other axial stresses were reported in Ref. 8.1 and the axial / shear interaction is evaluated as being acceptable based on the shear force margins calculated below.
The second mechanism to be discussed is the axial thermal shrinkage of the liner. Due to the general symmetry of the stud placements, the net shear force on a given stud due to this shrinkage will be close to zero due to the fact that there will be two equal and opposing forces acting at the stud.
For the purposes of this evaluation it will be conservatively assumed that the shear forces are not balanced but are reacted by the studs.
This postulated condition will be modeled as a section of liner plate 7t" x 7t" with two opposite edges anchored and subjected to a 15*F temperature drop.
10
- - ~
~
The force required to resist the shrinkage will be calculated and will be that assumed to be acting upon the studs.
-P P
1
]
3/4" l:
74"
- lY PL 6 = - = n h TL P = a bTAR P = (6.5 x 10-6 8
15 71 x 3/4 29 x 10
= 15,905 LH This very conservative shear force will be added to the previously calculated Ref. 8.1 maximum shear force and compared to the allowable.
Previous maximum shear load
= 2,700 LB (Ref. 8.1, Table 5.7-1) 32,500 LB (Ref. 8.1, Table 5.7-1)
Allowable
=
15,905 + 2,700 = 18,605 LB < 32,500 lb Total Load
=
Acceptable 6.3 PCRV Concrete The PCRV concrete was designed to be in a not compressive state of stress during the life of the plant. Previously analyzed loading conditions include initial prestress at atmospheric pressure, before heating at atmospheric pressure, design temperature at reference pressure, design temperature at 1.15X reference pressure, start of operation at working pressure, end of operation at reference pressure, and end of operation at atmospheric pressure (Ref. 8.1, Section 5.3). This compressive stress is due to the tensioning of the tendons which envelope the PCRV.
The effects of the proposed System 46 temperature change upon the PCRV concrete will now be investigated. Both long term and short term effects will be considered.
I 11
First, consider the case in which the System 46 temperature change has been in effect for a period of time such that thermal equilibrium has been established. This would have the effect of reducing the net effective bulk temperature of the PCRV concrete.
A bulk temperature decrease of the PCRV would have the effect of shrinking the PCRV and causing a decrease in the concrete compressive stress if the tendons were assumed to remain at the original higher temperature by some mechanism. A conservative measure of this decrease in compressive stress can be made as follows:
J The strain in a segment of concrete subjected to a 15'F temperature drop is Changein Length Stru m. =
Originallength a AT L L
(
= cat where a' = coefficientofthermalespansion
= 4.4 x 10'8 l a
'F (Ref.1, Section E.6.4) bT = 15'F Strain =
M.4 x 104HIS) = 0.000066 < < om3 allowableper Act318 89 or 0 0066%
The tendons were initially strained 0.61% (Ref. 8.1, Table 5.6 5). The change in the initial concrete compressive stress would then be on the order of the concrete strai 1 divided by the initial tendon strain or
= 0.011 or 1.1%
This change was conservatively calculated and is considered to be negligible, particularly in view of the fact that the PCRV is pressurized to a very low level for the proposed case (5 PSI maximum, Ref. 8.1).
Now, consider effects which could occur shortly afte'r the beginning of temperature chango. During the initiation of the proposed temperature change it is possible that some tensile stresses woulo develop on the inner surface of the PCRV concrete due to a non uniform temperature
-distribution. The limiting case for this stress would occur-if the PCRV concrete adjacent to the System 46 cooling tubes was assurned to instantly cool off the 15*F proposed for the event. In this case the resulting maximum 12
tensile stress at the concrete surface due to the restraint of the adjacent concrete would be:
o=
1-LN "
(Ref. 8.3, Pg 377, Article 3
6 88, Case 16) c-6 2
1-LN where hT = 1ST Anutned 4.4 x 10-6fn/In - T (Ref. 81, $ection E.10.3) a 2:
6 E = 4.82 x 10 pgj p = 0.1667 c = OutirltadiusofPCitV = 186 + (9 x 12?= 294" 6 = Innerliadiuso/PClf V = 31'x 12 x 1/2 = 186" 8
2 15 4.4 x 10' 4.82 x 10 2
294
,,v o=
l-2 2
2 1 - 0.1667 LN 294 - 186 o = 220 PSI < 233 PSI (Ref. 8.1, Section 5.4.4.3)
The calwlated worst case surface tensile stress is very conservative and is lower than the Ref. 8.1 allowable for unreinforced concrete. Thi; stress would tend to partially relieve the compressive stress which exists in the depressurized PCRV, This stress is not critical due to the presence of bonded reinforcing steel.
6.4 P_CRV Penetrations The General Arrangement Dravvings (Ref. 8.1, Figures 5.1-1 and 5.8 5) show the locations of the various penetrations through the walls and heads of the PCRV. All of the PCRV penetrations have carbon steel liners approximately 1/2 inch to 2 inch thick. The penetration liners are welded directly to the PCRV liner to maintain the continuity of the membrane enclosing the primary coolant (Ref. 8.1, Section 5.8.1).
All original primary closures are designed in accordance with the principles of ASME Boiler and Pressure Vessel Code,Section III, Class A, with a design pressure equal to the PCRV Reference Pressure of 845 psig.
The combination of the various loads is in accordance with the rules of ASME Section III fcr Class A vessels for primary closures.
13
During the proposed reduction in System 46 operating temperatures the PCRV penetrations could possibly be subjected to an incrcase in stress levels.
This stress could develop due to restraint of the penetration's thermal i
contraction due to the restraint offered by the adjacent concrete or by the penetration material itself. The limiting case for this differential thermal expansion induced stress can be calculated by assuming that the 15'F temperature change occurs instantly and that complete restraint is provided to any thermal movement. In this case, the maximum stress can be calculated as follows:
(Ref. 8.3, Case 2, Pg. 374) where AT = 15*F (Assumed, WorstCase) a = CoefficientofThermeiExpansion
= 6.5 x 10 ~' Inlin
- F E = modulusofelasticity = 30 x 10' PSI p = Poissons Ratio = 0.3 The values of a, E, and p above are typical values for carbon steel. The penetrations are made of various grades of carbon steel (Ref.1, Section 5.8.2.1) which may have slightly different values. The total stress levels in the penetrations are insensitive to these small variations.
o = (15)(6.5 x 10-8)(30 x 10') / (1 - 0.3) = 4,179 PSI (TENSION)
The tensile stress of 4,179 PSI calculated above is very conservative and represents the worst case value. This value will be added to the typical penetration stress analysis results found in Table 5.8 3 of Ref. 8.1 in a conservative manner. As the calculated stress is self-limiting it is seondary in nature. It will be added to primary + tecondary membrane + bending stresses of line 5 of Table 5.8.-3.
I 14
_ ST7ESSINTENSITY(Pi]
STRESS LIMIT (PSI)
BOTTOM ACCESS ~
PRIMARY 36,355 + L 4,179 40,534 69,300
=
. PENETRATION SECONDARY 37,139 + 4,179 41,318 52,500
=
e HELIUM CIRCULATOR PRIMARY 31,845 + 4,179 36,024
'52,500
=
- PENETRAWN SECONDARY 32,593 + 4,179 36,722 52,500
=
REFUELING PRIMARY 15,240 + 4,179 19,419 52,500
=
PENETRATION SECOND/aY 29,900 + 4,179 '=
34,079 52,500
- TOP ACCESS PRIMAR1' 47,700 + 4,179 41,879 69,300
=
PENETRATIONS SECONDARY E410 + 4,179 = 30,589 52,500 s_
lt is seen from above that with the addition of the conservatively calculated
_ thermal stresses that there is still a wide margin between the total stresses e
and the allowable stresses. It is anticipated that fewer than 10_ cycles of the
_ Ths fatigue damage due to
_ proposed temperature variation will occur.
even 10 times this amount (100 cycles) will be insignificant. It is concluded
- that the _ penetrations are structurally adequate for the proposed
' temperature variation.
3
' 6.5 _
PCRV Tendons The PCRV tendons were tensioned after the concrete was placed and before the. initial pressurization of the PCRV. The tensioning of the tendons serves to yield a net compressive stress in the PCRV concrete and liner during the entire life history of the plant.
1 The' highest state of tensile stress in _the tendons existed immediately after
- the initial prestressing. After this time the effects of concrete shrinkage and 2
creep tend to decrease somewhat the tendon tension. (See Ref. 8.1, Table 5.32).
The piiposed temperature changc is enveloped by the PCRV analysis previously performed. The results of this analysis are repor*2d in Ref. 8.1,
. Table 5.3-2. It is concluded that the PCRV tendons are acceptable for the proposed temperature change.
6.6' PCRV Rainforcina Rods Bonded reinforcement is provided in the PCRV to resist the computed forces and to distribute concrete cracks _ (Ref. 8.1, Section 5.5.1). The rebar is
' generally in a state of compression due to prestress applied by the PCRV tendons. This state of compression exists at the reactor reference pressure
'(845 PSIG).
15
The: proposed temperature change is enveloped by the PCRV analysis.
s previously performed. The results of this analysis are reported in Ref. 8.1, Table 5.3 2. It is concluded that the PCRV reinforcing rods are acceptable for the proposed temperature change.-
6.7 Core Support Floor 1.iner The Core Support Floor (CST)is an insulated and water cooled composite concrete and steel structure. The CSF is encased in a 3/4 inch thick steel liner (see Figure 6.1). The top and bottom surfaces of the liner have welded studs which are embedded in the CSF concrete and which enable the liner and concrete to act as a composite structure. The top, bottom, and sides of the liner have cooling tubes welded to them for which System 46 provides cooling water.
During the proposed System 46 operating temperature change it is possible
- th~ t the liner would be subjected to an increase in stress level. This possible a
stress would be due to differential thermal expansion between the liner and the concrete due to the fact that the thin liner would respond more quickly than the concrete to System 46 temperature variations.
The limiting case for this thermal stress can be calculated by assumin, that the liner instantly experiences the entire 15'F temperature drop of the proposed: scenario and, further, that the thermal strains due to this temperature dr_op are totally restrained. In this case the thermal stress can be calculated as follows:
f o = AT a E/(1 - p)
(Ref. 8.3, Pg. 374, Article 88, Case 2) -
where AT = 15'F (assumed) o = 6.5 x 10-8 Jdin
'F (Ref. B.4, Pg. 4) 8 E = 29 x 10 PSI p = 0.3 a=
15 6.5 X 10-8 29 X 10s/ 1 - 0.3
= 4,039 PSI 16
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Add this thermal stress to previously calculated stresses (from Ref. 8.1, Table 3.3 3) and compare to the allowable.
Area Stress (PSI)
Allowable (PSI)
Liner - maximum 2,600 + 4,039 = 6,639 23,100 tensile stress Liner - support 58,500 + 4,039 = 62,539 69,300 column junction The CSF liner stresses are acceptable with the addition of the conservatively calculated thermal stress. It is anticipated that fewer than 10 cycles of the proposed temperature variation will occur. The fatigue damage due to even 10 times this amount (100 cycles) will be insignificant. It is concluded that the CSF liner is st*ucturally adequate for the proposed System 46 operating temperature sange.
6.8 Core Support Floor Concrete The CSF concrete would possibly be subjected to some thermally induced stresses due to the initiation of the proposed temperature change. This stress would arise due to a non-uniform temperature distribution in the CSF concrete resulting from the fact that the surface of the concrete adjacent to the cooling tubes would respond more quickly to System 46 temperature variations than the concrete some distance from the surface concrete due to a non-uniform temperature distribution. The limiting case for this stress would occur if the PCRV concrete adjacent to the System 46 cooling tubes was assumed to instantly cool off the 15*F proposed for the event, in this case the resulting maximum tensile stress at the concrete surface due to the restraint of the adjr.<..t concrete would be:
(Ref. 8.3, Pg. 375, Article 88, Case 9)
,_g,g, where AT = 15*F Assumed a = 4.4 x 10 Inlin
- F 6
(Ref. 8.1, Section E.10.3)
E = 4.82 x 10 PSI p = 0.1667 6
o=
15 4.4 x 10 4.82 x 10 / 1 - 0.1667 = 382 PSI 4
l 17
1 4
This stress is somewhat higher than the allowable tensile stress of 232 PSI for unreinforced concrete (Ref. 8.1, Sect on 5.4.4.3) but is of no consequence due to the presence of bonded reinforcing steel.
6.9 Core Support Floor Reinforcino Bars The reinforcing bars of the CSF are subject to low levels of stress. The calculated rebar stress including the effect of the liner is 2300 PSI which is an order of magnitude lower than the 24,000 PSI allowable stress (Ref. 8.1, Table 3.3 3).
The rebar is in intimate contact with the CSF concrete with the result that the rebar and concrete wil! change temperature at the same rate. The temperatures anticipated for the proposed scenario are well within the bounds of ordinary reinforced concrete and no deleterious effects upon the rebar are anticipated. It is concluded that the CSF reinforcing rods are adequate for the proposed System 46 operating temperature change.
6.10 Fracture Mechanics Evaluation of the PCRV Liner Materials FSAR Sections 5.7.2.2 and E.24.5 discuss the experimentally determined initial and final Nil Ductility Transition (NDT) temperatures following exposure of each heat of liner material and a weldment of the liner material to an integrated neutron dose of 2.3 E18 n/cm2 (E M 1 MeV).
This integrated neutron dose utilized in the materials testing was the dose calculated for the most highly irradiated portion of the liner, at the top head, assuming a 30 year operational life at an 80% capacity factor (24 effective full power years, or 8760 effective full power days - EFPD). Each of the four heats of liner material had an initial NDT temperature below minus 60*F and experienced an increase in NDT temperature of less than 100*F 2
(FSAR Table E.24-16) following exposure to the 2.3 E18 n/cm integrated neutron flux. The weld metal had an initial NDT temperature of minus 75 F and experienced an increase in NDT temperature of 125 F (FSAR Table E.24-
- 16) following exposure to this same integrated neutron flux.
The reactor was permanently shut down on August 18, 1989, having accumulated 890 EFPD, which represents approximately one-tenth of the design lifetime and which corresponds to a maximum integrated neutron x
dose at the top head liner of 2.4 E17 n/cm2 (based on the integrated neutron flux equation in FSAR Section 5.7.2.2).
Assuming a linear correlation between neutron exposure and increase in the NDT temperatures, this neutron exposure would cause a shift in the NDT temperatures of approximately one-tenth (890/8760) of the experimentally determined NDT remperature shifts. The NDT temperatures are calculated to shift from minus 60*F to minus 50*F for the liner material and from minus 75 F to minus 62*F for the weldment material over the actual operating life of the reactor. The Fracture Transition Elastic (FTE) temperature is 18
4 approximately equal to the NDT + 60*F. The end of-life FTE temperatures are therefore calculated to be 10*F for the liner material and minus 2*F for the weldment material.
Maintaining the liner temperature above these FTE temperatures ensures that crack propagation in the liner at any tensile membrane stress up to yield stress would be incredible, and in this respect the liner meets the same criteria as are prescribed for steel nuclear pressure vessels, but is more conservative since the liner is in general compression during shutdown conditions, as it also was for all normal operating modes. Since the new 85'F minimum operating temperature of the PCRV liner is above the calculated end-of-life FTE temperatures of the liner and weldment materials, it is acceptable.
7.0 CONCLUSION
S The PCRV and CSF areas judged to be the most critical were analyzed for the conditions induced by the proposed 15*F System 46 temperature change. These areas were the PCRV liner, the PCRV liner anchor studs, the PCRV concrete, the PCRV penetration, the PCRV tendons, the PCRV reinforcing rods, the CSF liner, the CSF concrete, and the CSF reinforcing rods. The stresses due to the proposed 15'F were added in a conservative manner to stresses due to other loading conditions and were found to be within the allowab!a stresses as summarized in Section 2.0 of this EE.
It is concluded that the PCRV and CSF structures are structurally adequate for stresses due to the decrease of the average of the inlet and outlet temperatures from the 100*F specified in LCO 4.2.15 (e) to 85*F. This conclusion is based upon the premise that the maximum water outlet temperature is lowered from 120*F to 105*F (LCO 4.2.15 condition "b") and that no changes are made to conditions a, c, or d of LCO 4.2.15 (see Section 4.0 of this EE).
It is concluded that an adequate margin exists between the Nil Ductility Transition temperatures of the PCRV liner materials and the proposed 85'F temperature and that the liner materials will remain ductile at this temperature.
It is further concluded that adequate margins exist in the stresses and above the NDT temperature to allow an additional decrease in the average of the inlet and outlet temperatures should this become desirable. Any additional temperature decrease would require further analysis.
19
8.0 REFERENCES
8.1 Fort St. Vrain Nuclear Generating Station, Updated Final Safety Analysis Report, Revision 9.
8.2 Technical Specification LCO 4.7.1.
8.3 Formulas for Stress and Strain,4th Edition, Raymond J. Roark.
8.4 Stress Analysis for PSC/PCRV, G ADR 12 (Vol.1), dated 4/23/71.
8.5 Mechanicsof Materials,2nd Edition,Higdow I
I 20
FORT ST YRAiN NUCLEAR GENERATING STATION OPublic N.oc c5 I
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RES0i.UTION OF DESIGN DFFICIENCiES UNCOVERED DURING THE DESIGN VERIFICATION PROCESS l
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