ML19256G157
| ML19256G157 | |
| Person / Time | |
|---|---|
| Site: | Sequoyah  |
| Issue date: | 12/19/1979 |
| From: | Mills L TENNESSEE VALLEY AUTHORITY |
| To: | Rubenstein L Office of Nuclear Reactor Regulation |
| References | |
| NUDOCS 7912270533 | |
| Download: ML19256G157 (12) | |
Text
TENNESSEE VALLEY AUTHORITY CH ATTANOOGA. TENNESSEE 37401 400 Chestnut Street Tower II Dece=ber 19, 1979 Director of Nuclear Reactor Regulation Attention:
Mr. L. S. Rubenstein, Acting Chief Light Water Reactors Branch No. 4 Division of Proj ect Manage =ent U.S. Nuclear Regulatory Co==ission Washington, DC 20555
Dear Mr. Rubenstein:
In the hatter of the Application of
)
Docket Sos. 50-327 Tennessee Valley Authority
)50-32S Enclosed are proposed revisions to the Sequoyah Nuclear Plant Final Safety Analysis Report (FSAR).
These revisions are required to res.lve inconsistencies between the draft technical spec 2Neatieno and the FSAR.
Anend ent c4 will incorporate these revisions in the Sequoyah FSAR.
Very truly yours, TEhTESSEE VALLEY AUTHORITY
\\
L. M. Mills, Olanager Nuclear Regulation and Safety Enclosure 1645 043
@s 7912'270533 i
i 4, z u c: = <s na, e :.c m
SNP-23 i
.%1 design basis will be met for transients associated with Condition II events including overpower transients. There is an additional 5
large DNER =argin at rated power operation and during normal operating f
4.4.1.2 Fuel Temeerature Design Basis Basis During modes of operation associated with Conditien I and Condition II events, the maximum fuel temperature shall be less than the melting temperature 4
of UO,.
The UO, =elting temperature for at leas t 95% of the peak kW/f t I
fuel rods will not be exceeded at the 95% confidence level. The melting temperature of U0, is taken as 3080*F (Reference 1) unirradiated and reducing 58'F per 10,000 MWD /MTU. By precluding UO melting, the fuel geometry is preserved 3
and possible adverse effects of molte5 UO, on the cladding are eliminated.
To preclude center melting and as a basis'for overpower protection system setpoints, a ca.lculated centerline fuel temperature of 4700*F has been selected as the overpower limit. This provides suf ficient margin for uncertainties in the thermal evaluations as described in Subparagraph 4.4.2.10.1.
t Discussion Fuel rod thermal evaluations are performed at rated power, maximum overpower and during transients at various burnups. These analyses ascure that this design basis as well as the fuel integrity design bases given in Section 4.2 are met.
They also provide input for the evaluation of Condition III and IV. faults given in Chapter 15.
4.4.1.3 Core Flow Desien Basis Basis 42.6 70 g
A minimum of 95T5'% of the Thermal Flow Rate will pass through the fuel rod region of the core and be ef fective for fuel rod cooling. Coolant
(
flow through the thimble tubes as well as the leakage from the core barrel-a baffle region into the core are not considered effective for heat removal.
Discussion
- 9e
/ }A Core cooling evaluations are based on the Thernal Flow Rate (minimum flow) entering the reactor vessel. A maximum of AAT of this value is allotted as bypass flow. This includes RCC guide thimble cooling flow, head cooling flow, baffle leakage, and leakage to the vessel outlet nozzle.
4.4.1.4 Hydrodynamic Stability Design Bases Basis f
Modes of operation associa*.ed with Condition I and II events shall not lead to hydrodynamic instability.
5 b
4.4-2 July 1,19 75
.D"*D
"]D W Q SNP-29 o a J1 AlnLm 9
oo distributions are shown for the 4 foot elevation (1/3 of core height) in Figure 4.4-13, and 8 foot elevation (2/3 of core height) in Figure 4.4-14 and at the core exit in Figure 4.4-15.
These distributions are for the full power conditions as given in Table 4.4-1 and for the radial power density distribution shown in Figure 4.3-8.
The THINC code analysis for this case utilized a uniform core inlet enthalpy and inlet flow distribution.
i 4.4.2.7 Core Pressure Droos and Hydraulic Loads j
f1 s g G 4.4.2.7.1 Core Pressure Drops
/ y d $ ' *f
/
The analytical model and experimental data used to calculate the pressure drops shown in Table 4.4-T afelscribed in Paragraph 4.4.2. 5.
The core pressure drop h
includes theN0el assembly, core support plate, and holddown plate pressure drops.
29,
The full power operation pressure drop values shown in the Table are the unrecover-able pressure drops across the vessel, including the inlet and outlet nozzles, and across the core. These pressure drops are based en the Best Estimate "Icw (most likely value for actual plant cperating conditions) as described in Subsectica 5.1.1.
Subsection 5.1.1 also defines and describes the Thermal Design ? lev (minimum flow) which is the basis for reactor core thermal perfor=ance and the Mechanical Design Flow (maximum flev) dich is used in the mechanical design of the reactor vessel internals and fuel assemolies.
Since the Best Estimate Flcw is that flew which is most likely to exist in an cperating plant, the calculated core i
nrcssure d: cps in Table 4.i-1 are based on this best estimate ficw rather than the Thermal Design Flcw.
Uncertainties associated with the core pressure drop values are discussed in Subparagraph 4.4.2.10.2.
T'h essure drops quoted in Table 4.4-1 are bas'd on seven grids and conserra-tivelv'esfinated crid pressure loss coefficients Phase 1 of the D-iccp tests l
(Reference'51"resulted in a measured core pressure drop of a magnitude sufficient-e- \\[
i ly lewer than the predicted pressure drgp that the pressura drops quoted in Table
~
\\;
s.4-1 eill be conservative even with the addition of an eighth grid.
The estimated h
pressure drcp compaied to the measured"%Msure dropdn ~ Reference 5 uses the s
sa e conseriativelv estideed grfdepressure-16ss coef ficients used for the
~
Table 4.4-1pressuredropfaNulahi'dnsIThus, it was expected that the calculated
[
f pressure drop would be pense $Mivelldrger)'7slative to the measured value.
The Sequoyah fuel asymbfl[ grids, top nozzle, and bottom nozzle designs are the 29 same as in thepatotype assembly tests and the hydraulib resis tances measured f
during the test are therefore directly applicable to the Sequoygh analysis.
Further verification of the 17x17 core pressure drops including u'acertainties Q,i.ll-be cbtained from Phase 2 of the D-loop tests.
w
/
==._-- =,_
t 4.4.2.7.2 Hydraulic Loads p
~ ~
The fuel assembly hold down springs, Figure 4.2-2, are designed to keep the fuel assemblies resting on the lower core plate under transients associated l
with Condition II and II events.
Maximum flow conditions are limiting because hydraulic loads are a maximum.
The most adverse flow conditions occur during a LOCA.
These conditions are presented in Subsection 15.4.1.
l29 4.4-19 November 27, 1974 1645 045 r
SNP-29 Hydraulic loads at normal opera *ing conditions are calculated based on the Mechanical Design Flow which is described in Section 5.1 and accounting for the minimum core O
bypass flow based on manufacturing tolerances. Core hydraulic loads at cold plant startup conditions are also based on this flow but are adjusted to account i
for the coolant density difference. Conservative core hydraulic loads for a pump
.f overspeed transient, which create flow rates 20" greater than the Mechanical Design Flow, are evaluated to be greater than twice the fuel assembly weight.
Core hydralic loads were measured during the prototype assembly tests described in Section 1.'.
Reference 5 contains a detailed discussion of the results.
[
Lift forces arc directly proportional to the pressure drop. The lift force on an eight grid assembly is thus less than 5 percent greater than the lif t j
force on a seven grid assembly for the same flow rate.
Reference 5 shows that lift off of an eight grid fuel assembly (5 percent greater than the seven grid assembly shown) is not predicted during a postulated pump overspeed transient even though it is not necessary to preclude lif t of f.
Additionally, the flow 29 rates used in the test are based on a plant with flow ratese
,.' 3 mme-greater than Sequoyah. This results in additional margin for Sequoyah since the lift force is reduced.
The hydraulic loads during normal operation can be obtained from Reference 5 by I
adjustin: the loads fcr the Sequcyah pressure drop and flow rate. The effect I
of startup and shutdown transients are shewn to be inconsequential in Ref erence 5.
4.4.2.8 Correlation and Physical Data
.m 4.4.2.8.1 Surface Heat Transfer Coefficients Forced convection heat transfer coefficients are obtained from the familiar Dittus-Boelter correlation (Reference 53), with the properties evaluated at bulk fluid conditions:
hD DG 0.8 Cu 0.4
- = 0.023
(
)
( f, )
(4. 4-15 )
K u
A where:
9 h
= heat transfer coefficient, BTU /hr-f t
- F
~
De = equivalent diameter, ft K
thermal conductivity, BTU /hr-ft *F
=
2 G
= mass velocity, lb/hr-ft u
= dynamic viscosity, lb/ft-hr C
= heat capacity, BTU /lb *F F
This correlation has been shown to be conservative (Reference 54) for rod bundle geometries with pitch to diameter ratioe in the range used by P'a~ds.
6 The onset of nucleate boiling occurs when the clad wall temperature reaches f
the amount of superheat predicted by Thom's (Reference 55) correlation. After this occurrence the outer clad wall temperature is determined by:
O 4.4-20 November 27, 1974 2
1645 046
f i
SNP-23 through 27), and by measurements of the fuel and clad dimensions during fabrication.
The resulting uncertainties are then used in all evaluations t
involving the fuel temperature. The ef fect of densification on fuel temperature f
uncertainties is presented in Reference 6.
In additica to the temperature uncertainty described above, the measure-ment uncertainty in determining the local power, and the effect of density and enrichment variations on the local power are considered in establishing the heat flux hot channel factor. These uncertainties are described in Subparagraph 4.3.2.2.1.
Reacter trip setpoints as specified in the Technical Specifications Subsection i
16.2.3 include allowance for instrument and measurement uncertainties such b
as calorimetric error, instrument drif t and channel reproducibility, temperature measurement uncertainties, noise, and heat capacity variations.
Uncertainty in determining the cladding temperature results f rem uncertainties in the crud and oxide thicknesses.
Because of the excellent heat transfer between the surf ace of the rod and the coolant, the film temperature drop does not appreciably centribute to the uncertainty.
F 4.4.2.10.2 Uncertainties in Pressure Drops Core and vessel pressure drops based on the Best Estimate Flow, as described in Section 3.1, are quoted in Table 4.4-1.
The uncertainties quoted are based on the uncertainties in both the test results and the analytical extension The-cagnitude-ofC hel-uncertaintics_.
of these values to the reactor applicatien.
t Will-be-cenfirmed-when-the c.g. i.. ental-datarenethe prototype-fuel asser.blA._
Bett.ica--triFinobtained.
1 A majnr ase of the core and vessel pressure drops is to determine the primary system coolant flow rates as discussed in Section 5.1.
In addition, as discussed in Paragraph 4.4.4.1, tests on the primarv system prior to initial r
criticality will be made to verify that a conservative primary system coolant flow rate has been used in the design and analyses of the plant.
p 4.4.2.10.3 Uncertainties Due to Inlet Flow Ma1 distribution The ef f ects of uncertainties in the inlet flow maldistributica criteria used in the core thermal analyses is discussed in Subparagraph 4.4.3.1.2.
4.4.2.10.4 Uncertainty in DNB Correlatica The uncertainty in the DN3 correlation (Paragraph 4.4.2.3) can be written
.as a statement on the probability of not being in DSB based on the statistics of the DNB data. This is discussed in Subparagraph 4.4.2.3.2.
g 4.4.2.10.5 Uncertainties in DNER Calculations
[
[~
The uncertainties in the DNBR's calculated by THINC analysis (seeSubparagraph4.4.3.4.1)f due to uncertainties in the nuclear peaking factors are accounted for by applying 4.4-23 July 1 1974 1
- i 6-4-5-041
i SNP-23 conservatively high values of the nuclear peaking factors and including measurement error allowances.
In addition, conservative values for the engineering hot channel factors are used as discussed in Subparagraph 4.4.2.3.4.
The results of a sensitivity study (Reference 52) with THINC-IV show that the minimum DNBR in the hot channel is relatively insensitive to variations in the core-wide radial power distribution (for thesamevalueofFjg).
The ability of the THIMC-IV computer code to accurately predict flow and L
enthalpy distributinns in rod bmidlea is discussed in Subparagraph 4.4.3.4.1 I
and in Reference 63.
Studies have been performed (Reference 52) to determine
[
the sensitivity of the minimum DNBR in the hot channel to the void fraction correlation (see also Subparagraph 4.4.2.8.3); toe inlet velocity and exit pressure distributions assumed as boundary conditions for the analysis; and the grid pressure loss coef ficients. The results of these studies show that the minimum DN3R in the hot channel is relatively insensitive to variations in these parameters. The range of variations considered in these studies covered the range of possible variations in these parameters.
4.4.2.10.6 Uncertainties in Flow Rates The uncertainties associated with loop flew rates are discussed in Section s
5.1.
For core thermal performance evaluations, a Thermal Desian Loop Flow is used which is less than the Best Estimate Loop Flow (approximately 4% for g,the four-loop plant and 57: for the three-loop plant).
In addition another
,g v5Y of the Thermal Design Flow is assumed to be ineffective for core heat i
(,
removal capability because it bypasses the core through the various available vessel flow paths described in Subparagraph 4.4.3.1.1.
E 4.4.2.10.7 Uncertainties in Hydraulic Loads I
As discussed in Subparagraph 4.4.2.7.2, hydraulic loads on the fuel assembly are evaluated for a pump overspeed transient whidi create flow rates 207 greater than the Mechanical Design Flow. The Mechanical Design Flew as stated in i
Section 5.1 is greater than the Best Estimate or most likely flow rate value F
for the actual plant operating condition (by approximately 4.5%).
4.4.2.10.8 Uncertainty in Mixing Coefficient The value of the mixing coef ficient, TDC, used in THINC analyses for this application is 0.033.
The mean value of TDC obtained in the "R" grid mixing p
tests described in Subparagraph 4.4.2.3.1 was 0.042 (for 26 inch grid spacing).
The value of 0.038 is one standard deviation below the mean value; and s90% of the data gives values of TDC greater than 0.038 (Reference 46).
s The results of the mixing tests done on 17 x 17 geometry, as discussed in k
Subparagraph 4.4.2.3.3, had a mean value of TDC of 0.059 and standard deviation of a = 0.007.
Hence the current desig.n value of TDC is almos t 3 standard
- "h deviations below the mean for 26 inch grid spacing.
1645 048 4.4-24 July 1, 1974 t
~ *
- e
.m---
SNP-29 3.
Leakage flow from the vessel inlet nozzle directly to the vessel outlet nozzle through the gap between the vessel and the barrel.
Q<
%4 Flow entering into the core f rom.the baffle-barrel region through the gaps between the baffle plates, y,
,Vhe above contributions are evaluated to confirm that the desi,gn value of. core b'
ass flow is cet.
The design value of core bypass flow fo Sequoyah is equal t
9 f the total vessel flow. Of the total allowance,.2 M is associated 29 with the internals (Items 1, 3, and 4 above) and 2.0% for the core. Calculati ons have been performed using drawing tolerances on a worst case basis and accounting for uncertainties in pressure losses. Based on these calculations, the core bypass
[
flow for Sequovah is < s.K. This design bypass value is also used in the evaluation of the core pressure drops uoted in Table. 4.4-1, and the determinatica of reactor l
flow rates in Section 5.1.
7 5~ %
7 Flow model test results for the flow path through the reactor are discussed in Section 4.4.2.8.2.
4.4.3.1.2 Inlet Flow Distributions s
Data has been considered from several 1/7 scale hydraulic reactor codel tests (References 56, 57, and 64) in arriving at the core inlet ficw maldistribution criteria to be used in the THINC analyses (See Subparagraph I. 4. 3.4.1). THINC I ana1 3es :ade tuin;; this data ha te indicated that a conservative design basis 3
is to consider a 5 percent reduction in tae flew to the hot as s emb ly.
Reference 65.
The same design basis of 52 reduction to the hot assembly inlet is used in n
THINC IV analyses.
The experimental error esticated in the inlet velocity distribution has 1 een considered as outlined in Reference 52 where the sensitivity of changes in inlet velocity distributions to hot channel thermal performance is shown to e
be small. Studies (Reference 52) made with the improved THINC model (THINC-IV) show that it is adequate to use the 57. reduction ~ in inlet flow to the hot assembly for a loop out of service based on the experimental data in References 56 and 57.
r The effect of the total flow rate on the inlet velocity distribution was studied in the experiments of Reference 56.
As was expected, on the basis of the theo-F retical analysis, no significant variation could be found in inlet velocity distribution with reduced flow rate.
No relative ef fects on the core inlet velocity distribution caused by the change f rom a 15 x 15 to 17.17 fuel assembly array are expected since the icwer internals design will remain unchanged. The flow impedance of the lower core plate and fuel assembly noz:les is equal at all locations.
4.4.3.1.3 Empirical Friction Factor Correlations Two empirical f riction factor correlaticas are used in the THINC-IV computer code (described in Subparagraph 4.4.3.4.1).
I P
The f riction f actor in the axial direction, parallel to the fuel rod axis, is eval-usted using the Novendstern-Sandberg correlation (Reference 66). This correlation Q consists of the following:
4.4-26 November 27, 1974 e
1645-049
Added by Amendment 23, July 1, 1974 TABLE 4.4-1 (Continued)
REACTOR DESIGN COMPARISON TABLE L
Sequoyah Units Reference Plant 1 & 2 17 x 17 17 x 17 With Thermal and Hydraulic Design Parameters With Densification Densification fC1 Average in Core, *F
,5S18 585.9
~
Average in Vessel, 'F 578.2 584.7 r
Heat Transfer Active Heat Trancier, Surfacc Area, Ft 59,700 59,700 Average Heat Flux, BTU /hr-ft' 139,300 189,300 I
Maximum deat Flux, for normal bl l
~
'74,500 operation BTU /hr-ft
'74,500 k'i/ f t 5.44 5.44
~ f Average Thermal Cutput, Maximum Thermal Output, for E
I normal operation kW/f:
13.6 "l 13.6 ^l Peak Linear Pcwer for Determination of protection setpcints, kJ/f t 13.0[c]
13.0 Fuel Central Temperature Peak at 100" Power, 'F 3400 3400 Peak at Thermal Output Maximum for Maximum Overpower Trip Point, 'F 4150 4150 Pressure Drop
, z 3 g y, ej g,,j./ g, (,
J crf + dr0'
.2K O. %
2/
Across Core, psi
._SY.G
~L 424 6-i_E.4--
Across Vessel, including no::le psi h.GS y, j 3 g, g I
[a] This limit is associated with the value of F = 2.50 Q
[b] Based on best estimate reactor flow rate as discussed in Section 5.1.
[c] See Subparagraph 4.3.2.2.6.
1645 050 4.,-49 I
N Three reactor coolant flow rates are identified for the various plant O
design considerations.
The definitions of these flowc are presented in the following paragraphs, and the application of the definitions 3
is illustrated by the system and pu=p hydraulic characteristics on Figure 5.1-11.
Eest Estimate Flow The best estimate flow is the most likely value for the actual plant operating condition.
This flow is based on the best estinate of the reactor vessel, stean generator and piping flow resistance, and on the best estimate of the reactor coolant pu=p head, with no uncer-tainties assigned to either the system flow resistance or the pump head.
System pressure losses based on best estimate flow are pre-sented in Table 5.1-1.
Although the best estimate flow is the most likely value to be expected in operation, more conservative flow rates are applied in the thermal and =echanical designs.
Thernal Desi;n Flew Thermal design flow is the basis for the reactor core thermal perfor -
the steam generator thermal performance, and the naminal plant
- ance, parameters used throughout the design.
To provide the required cargin, the thermal design flow accounts for the uncertainties in reactor
~
vessel, steam generator and piping flow resistances.
The cc:binc. tion cf these uncertainties, which includes a conservative estimate of the punp discharge weir flow resistance, is equivalent to increasing q
the best esticate reactor coolant systen flow resistance by approxi-cately 18 percent. The intersection of this conservative flou resist-ance with the best estimate pump curve, as shown in Figure 5.1-11, establishes the thernal design flow.
This p margin for thermal design of approximately !.qocedure provides a ficw
.5/ percent. The ther:al design flow will be confirmed when the plant, s placed in operation.
Tabulations of i=portant design parameters based on the theraal design flow are provided in Table 5.1-1.
Mechanical Design Flow s.51 Mechanical design flow is the conservatively high flow used in the mechanical design of the reactor vessel internals and fuel assemblies.
To assure that a conservatively high flow is specified, the mechanical design flow is based on a reduced systes resistance (90 percent of best estinate) and on the maxi =un uncertainty on pung head capability (105.5 percent of best estimate for machined pump impellers). The intersection of this flow resistance with the higher pump curve shown on Figure 5.1-2, establishes the mechanical design flow.
, as The rasulting flow is approxi=ately 4 percent greater than the best estimate flow % 'psJ.'
/ 0 / 70 0 o p m Pu=p overspeed, due to a turbine generatcr overspeed at zu percent, rasults in a peak reactor coolant flow of 120 percent of the mechanical dasign flow.
The overspeed condition is applicable only to operating O
conditions when the reactor and turbine generator are at power.
1645 051 5.1-4
TABLE 5.1-1 SYSTDI DESIGN A!iD OPERATTiG PARA!ETERS Plant Design Life, years 40 tiominal Operating Pressure, psig 2235
. ~.
Total System Volume, including pressurizer and surge line, ft3 12,612 System Liquid Volume. including pressurizer water at maximum guaranteed power, ft3 11,892 04 tbet~ Powe, M We 34 //
r 5
NSSS Power,.T>tch:r-MW(
, Z
,.?
M23 System Thermal and Hydraulic Data Temperatures (Eased on Thermal Design Flow)
Thermal Design Flow, gpm/ loop
-4&,-500-9/ 400 Total Reactor Coolant Flow, lb/hr
/3g,/ h x 10 Reactor Vessel Inle Temperature, 'F
-MM C/[. '7 Reactor Vessel Outlet Tempera *ure. *F 4&~n+ dC4 7 Steam Generatar Out'.e Temperature, 'F l,.:. M 6M.. p Steam Pressure at Full Power, psia 857 Steam Generator Steam Temperature, ?
516 Steam Flow at Full Power, lb/hr (total)
/ t/,92
'.l M x 10 Feedwater Inlet Temperature, *F
<r39 4 3 Y. [
Pressurizer Spray Rate, max., gpm 800 Pressurizer lieat Capacity, kW 18C0 Pressurizer Relief Tank Vclume, ft 1800 Flows and Pressure Drops (Eased on Best Estimate Flow)
Best Estimate Flow, gpa/ loop
-Gh 500--- */ 7 80 0 Pump ilead ft.
26o Reactor Vessel AP, psi 46.2 Steam Generator AP, psi 34.6 h
Piping AP, psi 6.4 1645 052 5.1-9
. 5.5.7.3.3 Overpressurization Protection
' The inlet line to the Residual Heat Re= oval System is equipped with a pressure relief valve sized to relieve the combined flow of all e
the charging pumps at the relief valve set pressure.
s Each discharge line to the Reactor Coolant System is equipped with a pressure relief valve to relieve the maximum possible back-leakage through the valves separating the Residual Heat Removal Syste= froci the Reactor Coolant System. These relief valves are located in the Frergency Core Cooling System (see Figure 6.3-1).
The design of the Residual Heat ".enoval System in-ludes two isolation valves in series on the inlet line between the high pressure Reactor Coolant System and the lower pressure Residual Heat Removal System.
Each isolation valve is interlocked with one of the two independent Reactor Coolant System pressure signals. The interlocks prevent
@g the valves from being opened when Reactor Coolant System pressure p
is greater than approximately 425 psig.
If the valves are in the hf open position, the interlocks cause the valves to automatically
, \\\\[ f
/
close when the Reactor Coolant System pressure increases to ap,.
,)[,y p n ic;-; p sig.
These interlocks are described in more detail in 3ubsection 7.6.2.
5.5.7.3.4 Shared Function The safety function performed by the Residual Heat Removal System is not compromised by its normal function which is normal plant d,
cooldown.
The valves associated with the Residual Heat Removal Syste are nornally aligned to allow i=ediate use of this system in its safeguard code of operation.
The system has been designed in such a canner that two redundant flow circuits are available, assuring the availability of at least ene train for safety purposes.
The nor al plant cooldown function of the Residual Heat Removal System is accomplished through a suction line arrange =ent which is independent of any safeguards f uncticn.
The normal cooldown return lines are arranged in parallel redundant circuits and are utilized also as the low head safguards injection lines to the Reactor Coolant System. Utilization of the same return circuits for safeguards as well as for normal cooldown lends assurance to the proper functioning of these lines for cafeguards purposes.
5.5.7.3.5 Radiological Considerations The highest radiation levels experienced by the Residual Heat Removal System are those which would result f rom a loss of coolant accident.
Following a loss of coolant accident, the Residual Heat Removal System is used as part of the Emergency Core Cooling System.
During the recirculation phase of emergency core cooling, the Residual Heat Removal System is designed to operate for up to a year pumping water from the containment sump, cooling it, and returning it to the containment to cool the core.
1645 053 5.5-28
5.5.7.3.3 Overpressurization Protection The inlet line to the Residual Heat Removal System is equipped with a pressure relief valve sized to relieve the combined flow of all the charging pumps at the relief valve set pressure.
Each discharge line to the Reactor Coolant System is equipped with a pressure relief valve to relieve the maximum possible back-leakage through the valves separating the Residual Heat Removal System from the Reactor Coolant System. These relief valves are located in the Emergency Core Cooling System (see Figure 6.3-1).
The design of the Residual Heat Removal System includes two isolation valves in series on the inlet line between the high pressure Reactor Ccolant System and the lower pressure Residual Heat Removal System.
Each isolation valve is interlocked with one of the two independent Reactor Coolant System pressure signals. The interloc's prevent e
h[ y the valves from being opened when Reactor Coolant System pressure y
is greater than approximately 425 psig.
If the valves are in the
,, f j open position, the interlocks cause the valves to automatically V
/
close when the Reacter Coolant System pressure increases to apprc E-
))f
% eat ef-60 W g.
These interlocks are described in more detail p) -
in Subsectiot.,'. 6. 2.
5.5.7.3.1 Shared Function The safety function performed bv the Residual Heat Recoval System
~
is not ccepremised by its nor=al function *.hich is ner=al plhnt cooldown.
The valves associated with the Resi::ual Heat Renoval Syste= are normally aligned to allow innediate use of this systen in its safeguard code of operation.
The system has been designed in such a nanner that two redundant flow circuits are availabla, assuring the availability of at least one train for safety purposes.
The normal plant cooldown function of the Residual Heat Renoval Systen is accomplished through a suc'; ion line arrange =ent which is independent of any safeguards f unction.
The normal cooldown recurn lines are arranged in parallel redundant circuits and are utilized also as the low head safguards injection lines to tne Reactor Coolant System. Utilization of the same return circuits for safeguards
~
as well as for normal ecoldown lends assurance to the proper functioning of these lines for safeguards purposes.
5.5.7.3.5 Radiological Considerations The highest radiation levels experienced by the Residual Heat Remeval System are those which would result f rom a loss of coolant accident.
Following a loss of coolant accident, the Residual Heat Removal System is used as part of the Emergency Core Cooling System.
During the recirculation phase of emergency core cooling, the Residual Heat Removal System is designed to operate for up to a year pu= ping water from the containment sump, cooling it, and returning it to the containment to cool the core.
5 5-1645 054