ML20246P844
| ML20246P844 | |
| Person / Time | |
|---|---|
| Site: | South Texas  |
| Issue date: | 05/05/1989 |
| From: | HOUSTON LIGHTING & POWER CO. |
| To: | |
| Shared Package | |
| ML20246P849 | List: |
| References | |
| NUDOCS 8905220293 | |
| Download: ML20246P844 (56) | |
Text
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ERRATA Issued 05/05/89 Houston Lighting & Power correspondence ST-HL-AE-3021, dated March 30, 1989 was inadvertently distributed with typographical errors. Pages affected by this change have been annotated " Errata, 05/05/89". Please replace your copy with the attached revised ST-HL-AE-3021.
8905220293 890330 PDR ADOCK 05000498 F2 PNV n
L3/ errata.txt/nrc
- ATTACHMENT-
?-
~ ST Hi. AE 303 l
'PAGE I OF SS-i .
t ATTACHMENT Annotated Revisions-to FSAR and Technical Specification bases
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ATTACHMENT ST.HL. AE- J@
, reactivity compensation. The core is also designed to have an overall negative moderator temperature coefficient of reactivity so that average coolant temperature or void content provides another, slower compensatory effect. Nominal power operation is permitted only in a range of overall negative moderator temperature coefficient. The negative moderator tem-perature coefficient can be achieved through use of fixed burnable poison and/or control rods by limiting the reactivity held down by soluble boron.
Burnable poison content (quantity and distribution) is not stated as a design basis other than as it relates to accomplishment of a non positive moderator camperature coefficient at power operating conditions discussed above.
4.3.1.3 Control of power Distribution, Basis The nuclear design basis is that, with at least a,95 percent confidence level:
- 1. The fuel will not be operated at greater than 13.3 kW/ft under normal operating conditions including an allowance of 2 percent for calori-stetric arrer and not including power spike factor due to densification.
- 2. Ur der abnormal conditions including the maxirm overpower condition, the fuel peak power will not cause melting as defined in section 4.4.1.2.
. 3. The fuel vill not operate with a power distribution that violates the departure from nucleate boiling (DNB) design basis (i.e., the DNBR shall not be less than as discussed in Section 4.4.1) under h3 Condition I and II even including the maxiarm overpower condition.
- 4. k dL94% 24ct DM Fuel management will be such as tb produce rod powers and burnups consistent with the assumptions in the fuel rod mechanical integrity analysis of Section 4.2.
The above basis meets CDC 10.
Discussion Calculation of extreme power shapes which affect fuel design limits is performed with proven methods and verified frequently with measurements from operating reactors. The conditions under which limiting power shapes are assumed to occur are chosen conservatively with regard to any permissible operating state.
Even though there is good agreement between measured peak power cateulations and measurements, a nuclear uncertainty margin (section 4.3.2.2.7) is applied to calculated peak local power. Such a margin is provided both for the anal.
ysis for normal operating states and for anticipated transients.
4.3 3 Amendment 53
ATTACHMENT ST HL AE. 302 8 I STP FSAR PAGE 4 0F 66" 4.4 THERMA 1. AND HYDRAULIC DESIGN ,
4.4.1 Design Bases The overall objective of the thermal and hydraulic design of the reactor core is to provide adequate heat transfer which is compatible with the heat genera-tion distribution in the core such that ' heat removal by the Re* actor Coolant System (RCS) or the Emergency Core Cooling System (ECCS) (when applicable) assures that the following performances and safety criteria requirements are set:
- 1. Fuel damage (defined as penetration of the fission product barrier, i.e.
the fuel rod clad) is not expected during normal operation and operational transients (Condition 1) or any transient conditions arising from f aults of moderate frequency (Condition 11). It is not possible, however, to preclude a very sea 11 number of rod failures. 'Ihese will be within the capability of the plant cleanup system and are consistent with the plant design bases. -
- 2. The reactor can be brought to a safe state following a Condition III event with only a small fraction of fuel rods damaged (see above definition) although sufficient fuel damage might occur to preclude in::nediate resurp-tion of operation.
- 3. The reactor can be brought to a safe state and the core can be kept sub-critical with acceptable heat transfer geometry following transients 9 arising free Condition IV events.
I In order to satisfy the above criteria, the following design bases have been established for the thermal and hydraulic design of the reactor
- core.
4.4.1.1 Departure free Nucleate Boiling Design Basis. ,
Basis There will be at least a 95 percent probability that departure from nucleate' boiling (DNB) will not occur on the limiting fuel rods during normal operation and operational transients and any transient conditions arising from faults of
. soderate frequency (condition I and 11 events), at a 95 percent confidence level. E s or ca y hi haf be e se ati 'ly a by inifip he muy ur to n el te of ing at ( BR) e 1. , an fof Th a 1 atdon[ 1[
u BR of 1. w 11 nt nue o us
. I'mSer*
. k Diseussien By preventing DNB, adequate heat transfer is assured between the fuel clad and the reactor coolant, thereby preventing clad das. age as a result of inadequate cooling. Maximum fuel rod surf ace temperature is not a design basis as it i.~ .:. N 0 4.4-1 Amendment 18, $/1/81
ATTACHMENT ST HL AE. 30Al~
PAGE S OF St>
1 I
- Insert A
........ l
{
This criterion has been conservatively met by adhering to the !
following thennal design basis: there must be at least a 95%
robability that the minimum departure from nucleate boiling ratio p(DNBR) of the limiting power rod during Condition 5I and II events is greater than or equal to the DNBR limit of the DNB correlation being used. The DNBR limit for the correlation is established based on the variance of the correlation such that there is a 95% probability with 95% confidence that DNB will not occur when the calculated DNBR is at the DNBR limit.
Insert B *
, Historically, the DNBR li .it has been 1.30 for Westingl.cuse a;plicatiens, Jn this application, the WRB 1 correlation (Reference A) is used. With the significant improvement in the accuracy of thecorrelation, revious critical he61 flux prediction a DNBR N. c k 9byM. p using this correlation instead of,the i t f) ap a ad,Lusq.5 .s% nard spra.dM tw'eQ A plant. specific margin has been considered in the analyses.1heAplant safet) analysis DNBR limit of 3.27 was used in the safety analyses.
alloaance available between the DNBR limit used in the safety analyses and the design limit DNBR (7.8%) will be used to offset the effects of the ?CS flon anomaly and fuel rod bowing on DNBR and to provide for flexibilfy in the design, operation and analyses for the South Texas plants.
For conditions outside the range of parameters f or the WRE-1 correlation (ref er to Section 4.4.2.0.1), the W-3 correlation is used and a DNBR design limit of 1.30 appli es f or pressures equal to or greater than 1000 psia. For low pressure (500-1000 psi a) applications of the W-3 correlation, a design limit DNBR of 1.45 applies (Ref .C.be6ow ) .
's k
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ATTACHMENT
' 'STP FSAR . ST HL AE.3021 PAGE 6 OF S'5" 4
g
'4.4.2.2 Critical Heat Flux Rati r Departure from Noeleste Boilina Ratio
)
-and Mixina'Technolony. The minimus DNBR% for the rated power, design overpower and anticipated ~ transient conditions are given in Table 4.4-1.
The'sinfaus DNBR in the. limiting flow channel will be downstream of.the peak . heat flux location (hot spot) due to the increased downstream enthalpy
' rise. - i DNBR's are calculated by esing the correlation and definitions described j in the~fo11owing Subsections 4.4.2.2.1 and 4.4.2.2.2.. The THINC-IV computer
- code - (4.3-18, 4.3-49) is ased to determine the flow distribution in the core and the -local conditions in the hot channel for use in the DNB cor-i relation. The use of hot channel factors is discussed in Subsection 4.4.4.3.1 (nuclear hot channel factors) and in Subsection 4.4.2.2.4 (engineering hot channel factors).
4.4.2.2.1 DNB Technolonyt Early experimental studies of DNB were conducted with fluid flowing inside single heated tubes or channels and .
with single annulus configurations with one or both walls heated. The results of the experiments were analyzed using many different physical models for describing the DNB phenomenon, but all resultant correlations are highly empirical in nature. The evolution of these correlations is given by Tong [4.4-2, 4.4-3), including.the W-3 correlation which is in wide use in the pressurized water reactor (PWR) industry.
As testing methods progressed.to the use of rod bundles, instead of single g ' channels, it became apparent that the bundle average flow conditions cannot be used in DNB correlations.- As outlined by Tong [4.4-4) test results showed'that. correlations based on average conditions were not accurate spredictors of DNB heat flux. This indicated that a knowledee of the local subchannel conditions within the bundle is necessary.
r- great te- f44n ot~ tfst l afo &t cleS h M N M'$ N 8A' It is als r n6ted that in this power capability evaluation, there has not been' any$ change in the design basis. The reactor is designed to a minimus DNBR M as well as no fuel centerline melting during normal operation, ( 13 operational transients and faults of moderate frequency.
NB nalys s pe orme for his pliyttio ha in ud afNB mu 1 I[All y er .8 in a orda ew t resyfts 1 x1 se etr/D t t j
' 1. '
de/ete.
Fuel densification has been considbred in the DNB and fuel temperature L evalusi:fons utilising the methods and models described in detail in Reference 4.4-6.
In order to determine the local subchannel conditions, the TRINC [4.4-7) computer' code was developed. In the TRINC Code, a rod bundle is considered to be an array of subchannels asch of which includes the flow area formed by four adjacent rods. 'The subchannels are also divided into axial steps such that each any be treated as a control volume. By solving simultaneously the mass, energy, and acaente equations, the local fluid conditions in each cortrol volume e.re calculated. The W-3 correlation, devaleped from
% ainch ~~ w.1 ' ~ .vi pp1..".e4 to rod bundles by using the subchannel 6 locaA fau.a co.sa cica- .:ab.u atdd by the THINC code.
l 4.4-4 Amendment 18,5/1/81 w
l q
1 STp FSAR ATTACHMENT ST HL AE. 3021 (AGE 7 0F SS- ~
i L erE D: l
' )
) It was shown by Tong [4.4-4) that the above approach yielded conservative -
%Leredictions particularly in red bun'dles with mixing vane grid spacers.[ence a a correcuan sacsor was ceveloped to adapt sne W-3 correastaan, twnach a j veloped based on single channel data), to rod bundles with spacer ds.
< Th s correction factor, termed the " modified spacer factor", was dev oped as a tiplier on the W-3 correlation.
The no fied spacer factor was developed from cod bundle DNB tes results con-ducted the Westinghouse high pressure water loop at Columbi University.
These test were conducted on non-uniform axial heat flux tes sections to ;
determine t DNB performance of a low parasitic, top-split xing vane grid design, heres ter referred. to as the "R" grid. A descrip n of this test i program and a ry of the results are given below. e grid to be used in- l the 17 x 17 fue assembly will be similar in design to he "R" grid.
"R" grid red bundle DNB tests, References (4.4-8) d [4.4-9], were conducted over a wide range of inulated reactor conditions pplicable to this plant.
These conditions were:
Axial grid spacing 20 1 26 in. and 32 in. __.
Incal DNB quality -15 ree Local mass velocity a10gtto+15pergent to 3.7 x 10 lb,/hr-ft 2
local inlet temperature 440'F to 620'F Pressure 1490 to V local heat flux 0.3 x 10g440 psia t 1.1 6 x 10 BTU /hr-ft 2 Axial heat flux distributi Non-uniform (Cos u and u Sin u)
Reated length 8 ft. and 14 ft.
Heater rod 0.D. 422 in.
~
The experimental 1'progra consisted of a ENB est series for both an all and/or partial channel surfa heated condition in a rod bundle arranged in a 4 x 4 array. A radial p er profile was simulated erating the cantral 4 rods of the bundle at 1 percent hip,her power than the o r rods. Two test series were conducted 26 in, axial grid spacing: 1) all el surface baated condition (typ al call), 2) partial channel surface heat condition (thimble cold well cal .
For the t .able cold us11 test series, one of the central four heat rods was replace by an unheated red. The simulated unheated thimble was made of a thin eel rod over which were placed ceramic cylinders with an outer di ter equ to the thimble outer diameter. These thimbles are attached to the gr the same manner as in the reactor core using a sleeve which ta brased into e grid and then bulged out above and below the grid to connect to the thimble.
i L.de( &
% 4.4-5
hed C =
% g ort 6.s b en AMd
- ATTACHMENT '
ST HL AE 302 I I. PAGE 8 0F SF j
l 1
i i
i
[MSW The WRB-1 (Reference Of correlation was developed based exclusively CBF data (over on the large 1100 bank of mixing vane grid rod bundle points) that Westinghouse has collected. ;
The WRB-1 correlation, based on local fluid conditions repre-sents the rod bundle data with better accuracy over a w,ide ,
range of variables than the previous correlation used in
'derign.
This correlation accounts directly for both typical a,6 thimble cold wall cell effects, uniform cnd nonuniform heat flux profiles, and variations ,in rod heated length and in grid spacing.
1 The ' applicable range of variables is:
Pressure : 1440 < P 12490 psia Local Mass Velocity : 0. 9 < G Local Quality 3.7 lb/floE/106 t -br < -
Neated Length, Inlet to :
-0. 2 < X1oe 5 0.3 CHF Location Lh1 14 feet Grid Spacing :
Equivalent ~ Hydraulic Diameter 13 < 9<sp i 32 inches 0.37
- 5 < 0.G0 inches Equivalent Heated Hydraulic : 0.46 5 d,h 5 0.58 inches Diameter fd/
Figure 4.4-/ shows measured critical heat flux plotted against predicted critical heat flux using the WRB-1 correlation.
The WAB-I correhRon is e y //cs h/e .poy yg h af4 M< s 17xt7 w;t4' %"
- Pae/ assem/f NXI!!( we y;d ofegy,,,
th,E m at.tP - " h Q u h " appr ,. m, % a f .;
i
{
pe /efe EWNe ; A $a"[7073 ((
L 4
f&
These rod bundle tKB data have been analyzed ad a modified spacer factor, N Es ences 4.4-8 and 4.4-9, has been developed to conservatively incorpor e ;
the " ' mixing vane grid banefit for both typical and cold wall cells. is
. modifie acer factor is:
- .2)2 - o 73]
'Fj=( .895) (1.445 - 0.0371 L) [a .
0.3s -
+ IO I' ' '~ N 6 9}'
Ware:
P = the primary stem pressure Psia L = the total heat core length, ft.
x '= the local quali expressed in fragtional orm G = the local mass we city, Ib /hr-f t .
TDC = the thermal diffus n coefficient '
Kg = the axial grid spac coefficient ch has the following values:
Crid Spacing, in. h 32 0.027 26 0.046 g 20 0.066 c.
Figure 4.4-1 shows all the "R" gri typical 11 data. Figure 4.4-2 shows all the "R" grid thimble cell da ta. .a predicted eat flux in Figures 4.4-1 and 4.4-2 incorporates the modified pacer factor pe Equation (4.4-2) for typical calls and Equation (4.4-3) for thimble cold wall lis:
9"PRED
- 9"E'B N* h (4 4-2)
W-3 e ere:
is predicted non-uniform DNB heat flux us a the W-3 q"",),N y co elation as dest.ribed in kference 4.4-10.
9"PRED
- DNB.N.WxFj (4.4-3)
W-3 dere:
is the predicted ace-uniform DNB heat flui tse cell
",)
y 'I ' # having a cold (unbeated) vall evaluated with the W-3 cold well correlation described in Subsection 4.4.2.2.2 and h ference 4.4-3.
4.4-6
ATTACETENT
[5 1 0 6 I
- f. as defined in Equation (4.4-1) is the same in both Equations (4.4-2)
( a (4.4-3) for both typical and thimble cold vall cells. A Effect 17 x 17 Geometry on DNB A test progr imilar to the one described above was conduct at the Westing-house pressure va r loop at Columbia University. In thir. est program, DNB data was obtained fo 17 x 17 fuel assembly geometry, erence 4.4-5, in s 5 x 5 rod bundle array. -
Test results were obtained fo pical cells ( walls heated) in 8 ft and 14 f t btmales with unifort axial at flux for 14 f t typical and thimble cold wall cells with non-uniform t flux. All bundles were for mixing vane spacings of 22 in. (greater spac than this design). l 18 The data obtained were analyzed th the ez ing "1" grid DNB correlation described above to determine e effect on DNB the 17 x 17 fuel assembly ge ometry. Plots of ratio measured to predicte 1 heat flux versus flow parameters show that t "1" grid correlation proper 1 ccounto for local fluid parameters. ever, the *R" grid correlation cons tently overpredicts peDNBheatflux Bence, a multiplier of .88 on the modif spacer factor S is required o correctly predict the magnitude of the DNB hem lux for 17 .
x 17 geomet .
Figur 4-3 shows the 17 x 17 data obtained in this test program. The pr eted heat f2ux includes the 0.88 multiplier on the modified spacer factor.
untion (4.4-1), as noted above.
(
As stated in Subsection 4.4.1.1 Westinghouse has chosen the design criterion that DNB will not occur at a 95 percent probability with a V5 percent confi-dance level.
In order to meet this criterion, a liatting value of DNBR is determined by the method of Owen (4.4-11). Owen has prepared tables which give values of K such that *at least a proportion P of the population is greater than R7F-Kys
.with confidence 7,* where HTT and e are the sample mean and standard devis-tion, respectively. When this method was carried out using the data on Figure
-^ 2 ;,: o - may I 4.4-J, operatethewithresults indicated a minimus DNBR of that aandreactor satisfycore the deusingfign criterion.":-..;.I, n gg ,
_ - -..e t. e.... - t1 _. 4 u g=n_ -.. -
-- y 7 L- u y 4.4.2.2.2 Definition of Departure from Wueleste ling Ratio (DNBR):
The DNB heat flux ratio (DNBR) as applied to this designyn all flow cell walls are heated is: -Q g 'g
- ) Mfe, ,
where q"DNB. N is the heat flux predicted by the applicable DNB correlation For the W-3 correlation, ,
f 4,4 7 A ende.nt 16. SD /F1
x.invni..u i j ST HL.AE 302.1 SIP FSM PAGE // OF S*$"
I 9bNB, EU (4.4-5) .. ,
gjg3,p T
cnd 9' is the uniforc' DNB heat flux as predicted by the V 3 DNB correla-tien, uni [yrence 4.4-10 all flow cell valls are heated.
T is the flux shape factor to account for nonuniform axial heat flux distribu-tions, Reference 4.410, with the "C" term modified as in Reference 4.4 3.
9 r
ds a if d sp rfagor efine 'by quat n (4 -) n Section
.0 f
i d
. .1 nd ing g spa ng eff ient, , t r-o if sto oeff c n of .059 ase on t l . ri sp ci 16 d a . evi sly 4(s ibe . ince e a tual rid s ac g ,s 9. i., e di ed pacer f ter s nse att sine the !). B er o anc s u t ir ove nd T cr sep as a al id a cingJ4 cr d, e re ces .
8/
d4 12. TD vpiue 20
.061./.grfdspacrng (a fa e pe ,
/
i J.
spacJngas hi d ign) is g gj,,istheactuallocalheatflux. g/
The DFB heat flux ratio as applied to this design when a cold vall is present is-
/, O DNBR=9bNB.NCVM5h) (4.4 6) where:
95cc Yllll Q gjNB,N.CV~ NB'EU'Dh C4'4 7)
F where:
gbNBEU.Dh is the unifor kNB heat flux as predicted by the V 3 cold voll DNB correlation Reference 4.4-3, when not all flow cell valls are heated (thimble cold vall cell).
I CWF [4.4-3) - 1.0 Ru [13.76 1.372e - 4.732 ( C ) 0.0535 gg,4,g) 108 {
0 0.0619 ( P ) 0.14 - 8.509Dh .107) cnd Ru = 1 - De/Dh fnse(? [ f O C lt?' f D d y q tic . 1) in to 4. .2. 1 s es efa use ' orf f
/
yu-
)fy p y g in D- ov de try/T le e ab a tv d in) tie'ns f.
4 R
at t a
1 it ro rfate 'e) va e c'i a
iI at d
/
6 I
I al cel vitti a c 1 i al s e e ,erfa[himbecold i v21 e ndi i n).
The procedures used in the evaluation of DNB sargin for this application show ,
that the calculated minimum DNBR for the peak red or rods in the cort vill be If cboveg'during Class I and 11 incidents, even when all the c gir,* c rit.t. h:,t ,
4.4 8 c co m .a .%
tAe des y n n i./ M a
- . ATTACHMENT
. ST.HL AE. 302 i PAGE IAOF 66' n Insert E For the WRB-1 correlation, 4"WRB-1 9,DNB, N = ........-
y ,
(4.4.Ba) where F is the same flux shape factor that is used with the W-3 correlation.
Insert F-The safety analysis for South Texas cores maintained sufficient margin between thesafetyanalysisDNBRlimitandthedesignDNBRlimittoaccd.datefulland low flow 20NBR penalties identified in Reference 4.4 85 with the f6 corporation of the L /1 scaling factor-(1 = fuel rod bending moment of inertia, L = span.
length) to account for.17x17 XI span lengths.
l
- e
AT1 ACHMENT STP FSAR ST HL AE- 3021 PAGE O OF 5'5-N 1. Pellet diameter, density and enrichment l18 Design values employed in the THINC analysis related to the above fabrica-tion variations are based on applicable limiting tolerances such that these design values are met for 95 percent of the limiting channels at a 95 per-cent confidence level. Measured manufacturing data on Westinghouse 17 x 17 fuel show the tolerances used in this evaluation are conservative. The l18 effect of variations in pellet diameter, enrichment and density is employed in the THINC analysis as a direct multiplier on the hot channel enthalpy gig rise.
- 2. Inlet Flow Ma1 distribution The consideration of inlet flow maldistribution in core thermal performances is discussed in Section 4.4.4.2.2. A design basis of 5 percent reduction in coolant flow to the hot assembly is used in the THINC-IV analysis.
- 3. Flow Redistribution -
The flow redistribution accounts for the reduction in flow in the hot channel resulting from the high flow resistance in the channel due to the local or bulk boiling. The effect. of the non uniform power distribution is inherently considered in the THINC analysis for every operating condition which is evaluated.
4 Flow Mixing The subchannel mixing model incorporated in the THINC Code and used in reactor design is based on experimental data (4.4-17] discussed in Section 4.4.4.5.1. The mixing vanes incorporated in the spacer grid design induce additional flow mixing between the various flow channels in a fuel assemb'.y as well as between adjacent assemblies. This mixing reduces the enths.1py rise in the hot channel resulting from local power peaking or unfavorable mechanical tolerances.
4.4.2.2.5 Effects of Rod Bow on DNBR: The phenomenon of fuel rod bowing, as described in Reference 4.4 84, must be accounted for in the DNBR safety analysis of Condition I and Condition II events for each plant application. Applicable generic credits for margin resulting from retained conservatism in the evaluation of DNB plantoperatintparameters(suchasF'Qand/ormarginobtainedfromacasured or core flow), which are less limiting thanthoserecuiredbytheplantsafehanalysis,canbeusedteoffsetthe effect of rod bow.
c-The s fety a lysis f South Je'xas cor f maintain per ent) : accommo te fu11 4nd low sufficentmafin( .3 ow DNBR p alties ' entif d in 0,
\0LE. R erence .4-85 w h the corpora on of the /I sea ng fac or ( = fue g'e d ben ng mome of ine .ia, L - span lengt length . A des gn limi DNBR of to ac unt fo 17
.30 vs. 1 8,agr/dspac g co fici 7 XL span t
Q'M (Kg ) f .059
.05 (used 066 and a tb real diff ion coe ficien (TDC of .
r modif ed space / factor F; only) a exampjes o conse.
9 vs ati l5 u lized i the saf ry analysis. / /
/ -
The .aaximum rod bow penalties accounted for in the design aafety analysis are based on an assembly average burnup of 33,000 MiJo/MTU. At burnups 4.4-11 Amendment 53
m % nmem
. ST HL.AE. 302 i STP FSAR PAGE /4 OF E6" greater than 33,000 mwd /MTU, credit is taken for the effect of F" burndown, duetothedecreaseinfissionableisotopesandthebuildupoffNsionproduct (~
inventory, and no additional rod bow penalty is requiredg ( g gg
~4.4.2.3 Linear Heat Generation Rate. The core average and maximum Linear Powers are given in Table 4.4-1. The method of determining the maximum Linear Powers is given ir,section 4.3.2.2.
4.4.2.4 Void Fraction Distribution. The calculated core average and the hot subchannel maximum and average void fractions are presented in Table 4.4-3 for operation at full power with design hot channel factors. The void frac-tion distribution in the core at various radial and axial locations is present-ed in Reference 4.4 18. The void models used in the THINC IV computer code are described in Section 4.4.2.7.3. Normalized core flow and enthalpy rise distributions are shown on Figures 4.4-5 through 4.4 7.
4.4.2.5 Core Coolant Flow Distribution. Assembly average coolant mass velocity and enthalpy at various radial and axial core locations are given below. Coolant enthalpy rise and flow distributions are shown for the 1/3 core height elevation on Figure 4.4-5, and 2/3 core height elevation on Figure 4.4-6 and at the core exit on Figure 4.4-7. These distributions are for the full power conditions as given in Table 4.4 1 and for the radial power density distribution shown on Figure 4.3-7. The THINC Code analysis for this case utilized a uniform core inlet enthalpy and inlet flow distribution. No orificing is employed in the reactor design.
4.4.2.6 C_ ore Pressure Drops and Hydraulic Loads.
4.4.2.6.1 Core Pressure Drops: The analytical model and experimental data used to calculate the pressure drops shown in Table 4.4-1 are described in Section 4.4.2.7. The core pressure drop includes the fuel assembly, lower core plate, and upper core plate pressure drops. The full power operation pressure drop values shovn in Table 4.4-1 are the unrecoverable pressure drops across the vessel, including the inlet and outlet nozzles, and across the core. These pressure drops are based on the best estimate flow for actual plant operating conditions as described in Section 5.1.1. This section also l53 defines and describes the thermal design flow (minimum flow) which is the basis for raaetor core thermal performance and the mechanical design flow (maximum flsv) which is used in the mechanical design of the reactor vessel internals and fuel assemblies. Since the best estimate flow is that flow which is most likely to exist in an operating plant, the calculated core pressure drops in Table 4.4-1 are based on this best estimate flow rather than the thermal design flow.
Uncertainties associated with the core pressure drop values are discussed in Section 4.4.2.9.2.
4.4.2.6.2 Hydraulic Loads: The fuel assembly hold down springs, l18 Figure 4.2-2, are designed to keep the fuel assemblies in contact with the lower core plate under all Condition I and Il events with the exception of the turbine overspeed transient associated with a loss of external load. The hold down springs are designed to tolerate the possibility of an over deflection associated with fuel assembly liftoff for this case and provide i
contact between the fuel assembly and the lower core plate following this f .
1 i
4.4-12 Amendment 53
. ST HL AE. 302.1 STP HAR PAGE /6OF 56' Tests of the primary coolant loop flow rates will be made (see Subsection 4.4.5.1) prior to initial criticality to verify that the flow rates used in
.( '
f the design, which were determined in part from the pressure losses calculated I Qtr by the method described here, are conservative.
% 4.4.2.7.3 Void Fraction Correlation: There are three separate void y regions considered in flow boiling in a PVR as illustrated on Tigure 4.4-8.
They are the wall void region (no bubbl.e detachment), the subcooled boiling region (bubble detachment) and the bulk boiling region.
In the wall void region, the point where local boiling begins is determined
. when the clad temperature reaches the amount of superheat predicted by Thor's g [4.4-22) correlation (discussed in Subsection 4.4.2.7.1).
The void fraction gw in this region is calculated using Maurer's [4.4-27] relationship. The bubble w detachment point, where the superheated bubbles break away from the wall, is determined by using Griffith's (4.4-28) relationship.
h The void fraction in the subcooled boiling region (that is, after,the detachment
) is calculated from the Bowring [4.4-29) correlation. This correlation 0 pointpredicts the void fraction from the detachment point to the bulk boiling rejion.
g The void fraction in the bulk boiling region is predicted by using homogeneous flow theory and assuming no slip. The void fraction in this region is therefore g a function only of the thermodynamic quality. 84 gg //pjf pg, 4.4.2.8 Thermal Effects of Operational Tr nsients. DNB core safety limits
( are generated as a function of coolant tempera re, pressure, core power and axial power imbalance. Steady-stat cperatio within these safety limits
._ Figure 15.0-1 shows M l insures tha_t the minimus DNBR is not less tha 1 a , mmr4MM limit linegand the resulting overtesperature AT trip lines (which become pan d the Technical Specifications), plotted as AT, versus Tavg for various pressures. This system provides adequate protection against anticipated operational transients that are slow with respect to fluid trans-port delays in the primary system. In addition, for fast transients (e.g.,
uncontrolled rod bank withdrawal at power incident (Section 15.4.2)) specific protection functions are provided as described in Section 7.2 and the use of these protection functions are described in Chapter 15.
l 4.4.2.9 Uncertainties in Estimates.
4.4.2.9.1 Uncertainties in Fuel and Clad Temperature: As discussed in Subsection 4.4.2.11, the fuel temperature is a function of crud, oxide, clad, gap, and pellet conductances. Uncertainties in the fuel temperature calcula-tion are essentially of two typast fabrication uncertainties such as variations in the pellet and clad dimensions and the pellet density; and model uncertain-ties such as variations in the pellet conductivity and the gap conductance.
These uncertainties have been quantified by comparison of the thermal model to the impile thermocouple sensurements, References 4.4-30 through 4.4-36, by out-of-pile measurements of the fuel and clad properties Reference 4.4-37 through 4.4-48 and by measurements of the fuel and clad dimensions during fabrication. The resulting uncertainties are then used in all evaluations involving the fuel temperature. The effect of densification on fuel temperature
[ uncert ainties is pa esented in Ref erence 4.4-6.
4.4-15 Amendment IE. 5/1/S1
. ST HL At.- ov*
- PAGE /6 OF ES" SU T .c p,r, flow blockages in terms of maintaining rated core performance are detere.ined
'(- both by analytical and experimental methods. S -periments) dats are a' to usually used to augment analytical tools such as corputer progra .s sirilst the THINC IV program. Inspection of the DNE correlation (Subsection 4.4.1.2 and Reference 4.4-10 shows that the predicted DNBR is dependest upon the local values of quality and mass velocity.
The THINC IV Code is capable of predic' ting the effects of local flow blockages on DNSR within the fuel assembly on subchannel basis. regardless of where the flow blockage occurs. In Reference 4.4 49 it is shown that for a fuel assembly similar to the design THINC.IV accurately predicts the flow distribution within the fuel assembly when the inlet nor.sle is co=pletely blocked. yull recovery of the flow was found to occur about 30 in dovnstrean of the blockage. With the reactor operating at the nominal full power conditions specified in Table 4.41, the effects of an increase in enthalpy and decrease in mass velocity in the lover portion of the fuel assembly would not result in the reactor reaching a minimus DNBRM/cys g4c m g egg, f afe g;;gs. n /,wrl r M M o From a review of the open literature it is conclu5ed that flow blockage in
'open lattice cores' aimilar to the cores cause flow perturbations which are local to the blockage. For instance, Ohtsubol [4.4 81] at al., show that the
~
mean bundle velocity is approached asymptotically about 4 in, downstream from a flow blockage in a single flow cell. Similar results were also found for 2 and 3 cells complet'ely blocked. Basser [4.4-82), at al.. tested an open lattice fuel assembly in which 41 parcent of the subchannels were completely blocked in the center of the test bundle between spacer grids. Their results show the stagnant sone behind the flow blockage essentially disappears after 1.45 L/De or about 5 in, for their test bundle. They also found that leakage
[ flow through the blockage tended to shorten the stagnant sons or, in essence, A.
the complete recovery length. Thus, local flow blockages within a fuel in assembly have little effect on subchannel anthalpy rise. The reduction If the local mass velocity is then the main parameter which affects the DNBR.
plants were operating at full power and nominal steady state cenditions as specified in Table 4.4 1. a reduction in local mass velocity a"_ eater than 58 l percent would be required to reduce the DNBRIfter LM t#1.30/1 The above mass velocity effect on the DN8 correlation was basedIn onsality the assumption of l16 a local flow f fully developed flow along the full channel length.
blockage is expected to promote turbulence and thus v d likely not effect DNBR at all. ggg /4,f.
Coolant flow blockages induce local crossflows as well as promote turbulence.
Fuel rod behavior is changed under the influence of a sufficiently high cross-flow component. Fuel rod vibration could occur, taused by this If the crossflow component, through vortes shedding or turbulent mechanisms.
crossflow velocity saceeds the limit established for fluid elastic stability.
1arge amplitude whirling results. The liatts for a controlled vibration sechanism are established from studies of worten shedding and turbulent pressure fivetuations. The crossflow velocity required to saceed fluid elastic stability liatts is dependent on the asial location of the blockage These limits are and the characterization of the crossflov Ost flow or not).velocity above the Crossflow greater than those for vibratory fuel rod wear.
established limits can lead to mechanical wear of the fuel rods at the grid support locations. Fuel rod wear due to flow induced vibration is considered in the fuel rod fretting evaluation (Section 4.2).
a.4 M /c < :.i. . :. , !!
3 STP FSAR . ST HL AE. So2.1 PAGE / 7 OF 5"S~
~
RZfT1ENCIS SECTION 4.4.71 Christensen, J. A., Allie, R. J. and Biancheria, A., " Melting Point 4.4-1.
of Irradiated UO ," WCAF-6065, February,1965.
i 2 4.4-1. Tong, L. S., "Sciling East Transfer and Tw-Phase Flev," John Wiley 4 Sons, New York,1965. .
4.4=3. Tong, L. S., " Soiling Crisis and Critical Beat Flux," AEC Critical Saview Series TID-25867,1972.
4.4-4 Tong, L. 8., " Critical East Fluzes on led Dundles," in "Tw-Phase _
Flow ad Heat Transfer in mod Sadles," pp. 31-41, American ,
T Society of Hachanical Engineers, New York,1969.
? 7., " sacal N F1ux
. t1ey, F. Qvental A. R. and CAS Aassably vetry vit - 2 Inch .
st N .
cing,' [7' .C"' u . (7r istary), , 1975 an ' CAP-8537,\
[ # Ney, 975.
' * * - - ' ~ - - -- - - 1 4.4=6. Es11aan, J. M. (Ed.), " Fuel Densification Experimental Results and Model for Mctor Application," WCAP-4218-7-A (Proprietary) i March,1975 and WCAP-8219-A, March,1975. .
4.4-7. Cheleser E., Weisman, J. and Tong, L. 5., "Subchannel Nrmal Analysis of Rod Bundia ceras " WCAP-7015. Revision 1. January, 1969.
1 I and Cada . "W5 Test suite for Mixing 8 Motley, .
-76 P-A (Fro e ry),J ry, 75 d Vae s (R) 1
-795 anuary, 1975.
i 4 -9. tal F. . and k, F. F "W3 st uits R d lle, WCAP-76 dondum 1=? Froprietar -
Th e Cold Wall i A. January, 75.
J us 1975 and -7958-Addendus j
4.d=10. Teng, L. 8., "Frediction of Departure froa Fucisate Sciling for '
an Axially Non-Dniform Meat Flux Distribution," J. Fue.l. Eneray. ~
g
& 241-248 (1967). '
4.4-11. Owen, D. E., 'Fasters for One-81 dad Tolerasce Limits and for Variable Sampling Plans," 3C1407, Marsh,1963. ;
4.d=12. Cadek, F. F., MotleF, F. 5. and Dominicia, D. P., "Iffect of Axisi i Spacing en Interchanasi Ntiaal Mixing with the I Mixing Vane 4 Crid," WCAf=7941a?-A Proprietary), January,1975 and WCAF-7959-A, q January, 1975.
~j 4.4-13. nove, D. S., Angle, C. W., "" Crossflow MLxing netween Parallet Flow Channels During Soiling, Part 12 Measursaants of Flev and
'Mthalpy in two Parallel Channels," 3NWi-371, part 2, December. .f
,* ~
sni.
4.4=36
ST HL AE.3021 . _ _ _ _ _ _ _ _ _ _ _
PAGE /6 0F FC hlN References A. Motley, f. E. , Hill, K. W. , Cadek, F. F. and Shefchek, J. , *New Westinghouse Correlation WRB.1 for Predicting Critical Heat flux in Rod Bundles with Mixin Vane Grids," #Cr 3752 Jdy, 10M (Proprietary) and WCAP-8763, July 76 (Non. Proprietary)4 l Withf-t%2.rP-A , Jul y@ l B. Request for Reduction in Fuel Assw.bly Burnup Limit for Calculation of Maximum Rod Bow Penalty." Letter, C. Berlinger (USNRC) to E. P. Rahe, Jr.
(Westinghouse), June 18, 1986.
M % 4. C, p Le t t er from A. C. Thandani (NRC) to W. J. Johnson (Westinghouse), January 31, 196V,
Subject:
Acceptance for Referencing of Licensing Topical Report, WCAP-9226 a/9227-NP, " Reactor Core Response to Excessive Seedndary Steam Releases" m -
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Figure 4.4 3.
_ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ - - - - - _ l
S$lQL; sp/g- .
" ~ ' ' ~
up to 2,000 spatial poIn'ts, and performs its own steady Jtat.: *aitialization. .
Aside from basic cross section data and thermal-hydraulic pa,tameters, the code accepts as input basic driving functions such as inlet temperature, pressure, flow, boron concentration, control rod motion, and others. Various edits are provided (e.g., channelwise power, axial offset, enthalpy, volumetric surge, pointvise power, and fuel temperatures).
The TWINKLE code is used to predict the kinetic behavior of a reactor for transients which cause a major perturbation in the spatial neutron flux dis.
tribution.
TVINKLE is further described in Referenca 15.0 8.
15.0.10.6 THINC. The THINC code is described in Sution 4.4. 2 15.0.11 Summary of Accident Re uits Q211.7 For all Condition II ransie MiJ b.M YM h A nalyzecr in the FSAR, the calculated miniaram DNBR is greater than For each of chese transients, the peak RCS pres-sure is less than the safety limit of 110 percent of design pressur's (2750 psia) and there is no failed fuel as a result of the transient. For all of 60 the applicable Condition III transients, the minimum DNBR is greater than and there is no failed fuel, except for a single RCCA withdrawal at full power. For this transient, the upper bound of the number of fuel rods 3 experiencing DNBR is 5 percent of the total rods in the core. All of %
Q theapplicableConitionIIItransieny*exerenceapeakRCSpressureless than 2750 psia. * [ l
%gg y g
All the applicable Condition IV transients analyzed in the FSAR have a minimum 60 DNrs(" except LOCA, locked reactor coolant pump rotor, and rod ejection.
For these three transients, the amount of failed fuel is $100 percent, $10 l60 percent and $10 percent, respectively. Major rupture of a steam line experi-J ences $5 percent failed fuel. All other applicable transients have no failed fuel. All of the applicable Condition IV transients experience a peak RCS l60 pressure less than 2750 psia.
Block diagrams identifying protection sequences for Chapter 15 events are 2 prov;ded in Figures 15.0 7 through 15.0 31. Q211.(
l W WlU.l l
l 15.0 11 Amendment 60
PAGE 260F FS" 15,ti31 -1
.. .110 -- ---
LDCUS OF CONDl110NS WHERE POWER 3 llM OF p0as =AL. toa THIRMAL Dt51Ga FL0w 100 -
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STp 75gp 15.1 7FC". ". 7' HEAT R.EMOVAL SY THE SECONDARY SYSTEM - -
A number of events has been 'pos'tulated which could result in an increase in heat removal from.the Reactor Coolant System (RCS) by the secondary s/ stem.
Analyses are presented for several such events which have been identified as limiting cases.
Discussions of the following RCS cooldown events are ' presented:
. (Section 15.1.1).
- 3. Excessive increase in secondary steam flow (Section 15.1.3).
- 4. Inadvertent opening of a steam generator relief or safety valve causing a depressurization of the Main Steam System (Section 15.3.4). *
- 5. Spectrum of steam system piping failures inside and outside Containment (Section 15.1.5).
The above are considered to be American Nuclear Society (ANS) Condition II events, with the exception of a major steam system pipe break, which is con.
sidered to be ANS Condition IV event (Section 15.0.1).
15.1.1 Feedwater System Malfunctions Causing a Reduction in Feedwater Temperature 15.1.1.1 Identification of Causes and Accident Description. Reductions in feedwater temperature will cause an increase in core power by decreasing reactor coolant temperature. Such transients are attenuated by the therms 1 capacity of the secondary system of the RCS. The overpower overtemperature protection (neutron overpower, and overtemperature and overpower AT tripa) k3 prevents any power increase which could les o a departure from nucleate boiling ratio (DNBR) less than P99. & S '
g2 M Wlut. 10 A reduction in feedwater temperature could be caused by the accidental opening of a feedvater bypass valve which diverts flow around one of the high presaure feedvater heaters, or by the accidental closing of the extraction steam block valves or nonreturn valves to the hi h5 pressure feedwater heaters. (The deserator will attenuate any upstream disturbances, i.e., low pressure heaters out of service, loss of extraction, steam, etc.) In the event of an acci- 43
. dental opening of a bypass valve, there is a sudden reduction in feedwater 57 inlet temperature to the steam generators. If the extraction steam valves are accidentally closed, a more gradual, thou8h greater, reduction in feedwater inlet temperature to the steam generators will occur. At power, this in-creased subcooling will create a greater load demand on the RCS.
With the plant at no load conditions, the addition of ccid feedwater may cause a decrease in RCS temperature and, thus, a reactivity insertion due to the effects of the negative moderator temperature coefficient of reactivity.
However, the rate of energy change is reduced as load and feedwater flow decrease so the transient is less severe than the full power case.
15.1 1 Amendment 57
- ' . ST HL AE. 302 I PAGE.2 8 0F SS*
)
15.1.1.4 conclusions. The decrease in feedwater temperature transient is less severe than the increase in feedwater flow event (Section 15.1.2) and the increase in secondary steam flow event (Section 15.1.3). Based on results presented in Section 15.1.1 and 15.1.3, the applicable acceptance criteria for the decrease in feedwater temperature event have been met. There are no radi-ological consequences of this event.
15.1.2 Feedwater System Malfunctions causing an increase in Feedwater Flow 15.1.2.1 Identification of causes and Accident Description. Additions of excessive feedwater will cause an increase in core power by decreasing reactor co-Jant temperature. Such transients are attenuated by the thernal capacity of the secondary system and of the RCS. The overpower overtempera- 43 ture protection (neutron overpower, and overtemperature and overpower AT trips) prevents any pver. increase which could lead to a DNBR less than.2::$&. l18 IN JKTLMy .54J lA d ValM4.
An example of excessive feedvater flow would be a full opening of a feedwater control valve due to a feedwater control system malfunction or an operator error. At power, this excess flow causes a greater load demand on the RCS due to increased subcooling in the steam generator. With the plant at no-load conditions, the addition of cold feedwater may cause a decrease in RCS temper-ature and, thus, a reactivity insertion due to the effects of the negative isoderator temperature coefficient of reactivity. 43 Continuous addition of excessive feedwater is prevented by the steam generator high high water level signal, which initiates feedwater isolation. The high-high steam generator water level signal initiates a turbine trip which 43 57 then initiates a reactor trip.
An increase in normal feedwater flow is classified as an ANS Condition 11 event, a fault of moderate frequency (See Section 15.0.1).
Plant systems and equipment, which are available to mitigate the effects of the accident, are discussed in Section 15.0.8 and listed in Table 15.0 6.
A block diagram summarizing various protection sequences for safety actions 2 required to mitigate the consequences of this event is provided in Figure Q211.6 15.0-7.
15.1.2.2 Analysis of Effects and Consequences.
Method of Analysis The excessive heat removal due to a feedwater system malfunction transient is
~
analyzed by using the detailed digital computer code 14FTRAN (Ref.15.1-1).
This code simulates a multiloop system with neutron kinetics, pressurizer, pressurizer relief and safety valves,' pressurizer spray, steam generator, and 43 steam generator safety valves, The code computes pertinent plant variables including temperatures, pressures, and power level.
A control system malfunction or operator error is assumed to cause a feedwater control valve to open fullf. Two cases are analyzed as follows:
15.1 3 Amendment 57
M 29..OF SF STP FSAR automatic rod control mode results in a similer, ~2ut less limiting. . transient. 57 The rod control system is not required to functian for an excessive feedwater flow event.
The calculated sequence of events for this accident is shown in Table 15.1 1. 43 When the. steam generator water level in the faulted loop reaches the high high level setpoint, all feedwater isolation valves and feedwater control valves h3 are automatically closed and the SG feed pumps are tripped. This prevents continued addition of feedvater. In addition, a turbine trip is initiated.
Following turbine trip, the reactor will be automati..a12y tripped directly due to turbine trip. If in manual rod control, the ensuing transient would then be similar to a turbine trip event as analyzed in Section 15.2.3 resulting in an overtemperature AT signal. If the reactor were in the automatic control 57 mode, the control rods would be inserted at the maximum rate following turbine trip.
Transient results, see Figures 15.1-1 and 15.1-2, show the core heat flux, pressurizer pressure. Tavg and DNBR as well as the increase in nuclear power and loop AT associated with the increased thermal load on the reactor. The DNBR does not drop below Following the reactor trip, the plant approaches a stabilized h8 ndition; standard plant shutdown procedures may then be followed to further co down.the lant. .
.Jo f4 ddydJ /M. W/ut .
Since the power level rises during th excessive feedwater flow incident, the fuel temperatures will also rise until after reactor trip occurs. The core heat flux lags behind the neutron flux response due to the fuel rod thermal time constant, hence the peak value does not exceed 118 percent of its nominal value (i.e., the assumed high neutron flux trip setpoint). The peak fuel temperature will thus remain below the fuel melting temperature.
The transient results show that departure from nucleate boiling (DNB) does not occur at any time during the excessive feedwater flow incident; thus, the ability of the primary coolant to remove heat from the fuel rod is not reduced. The fuel cladding temperature, therefore, does not rise signif-icantly above its initial value during the transient.
15.1.2.3 Radiolerical Consequences. There are only minimal radiological consequences from this event. The turbine trip causes a reactor trip and heat is removed from the secondary system through the steam generator power oper- 43 ated relief valves (PORVs) or safety valves. Since no fuel damage is postu-lated to occur from this transient, the radiological consequences are less severe than the steam line break analyzed in Section 15.1.5.3.
15.1.2.4 conclusions. The results of the analysis show that the DNBRs encountered for an excessive feedwater addition at power are at all times above hence, the DNB design basis as described in Section 4.4 is set.
Additi nally, it has been shown that the reactivity insertion rate which lI6 occurs t no load conditions following excessive feedwater addition is less than t maximum value considered in the analysis of the rod withdrawal from a suberit cal condition analysis. The radiological consequences of this event are not limiting.
I h YE j MN N 15.1-5 Amendment 57
. ST HL AE 302 :
- 2. Esactor centrol in manual with maximum moderator reactivity feedback.
- 3. Reactor control in automatic with minimum moderator reactivity feedback.
- 4. . Reactor control in automatic with maximum moderator reactivity feedback.
For the minimum moderator feedback cases, the core has the least negative moderator temperature coefficient of reactivity and, therefore, the least inherent transient capability. For the maximum moderator feedback cases, the moderator temperature coefficient of reactivity has its highest absolute val-me.
This results in the largest amount of reactivity feedback due to changes in coolant temperature. For all cases, a small (absolute value) Doppler coef-ficient of reactivity is assumed (see Figure 15.0-2).
A conservative limit on the turbine valve opening is assumed, and all cases are studied without credit being taken for pressurizer beaters. Initial oper-sting conditions are assumed at extreme values consistent with the steady-atate full power operation allowing for calibration and instrument errors. This assumption results in minimum margin to core DNB at the start of the accident.
Plant 15.0.3.characteristics and initial conditions are further discussed '
in Section Normal reactor control systems at:d EST systems are not required to function.
The RTS is assumed not to be operable in order to show that DNBR criteria v111l43 be satisfied in the absence of reactor trip. No single active failure vill prevent the RTS from performing its intended function when required. 18 l43 The cases which assume automatic rod control are analyzed to ensure that the worst case is presented. The automatic function is not required as a safety feature. l18 Results The calculated sequence of events for the excessive lead increase incident is shown in Table 15.1-1.
Figures 15.1-3 through 15.1-6 illustrate the transient with the reactor in the manual control mode. For the minimum moderator feedback case, there is a slight power increase, the average core temperature decreases, and the DNBR increases slightly. For the maximum moderator feedback, manually controlled 18 case there is a much larger increase in reactor power due to the moderator feedback.
.14 N analy$bfinid h/as.A reduction in DNBR is experienced but DNBR remains above k:$h: k l 18 Figures 15.1-7 through 15.1-10 illustrate the transient assuming the reactor is in the autcastic control mode. Both the minimmm and maximum moderator feedback cases show that core power increases, treby reducing the rate of decrease in coolant average temperature and pressurizer pressure so that there are only small changes in these parameters, as is also the case for the maxi- 18 sua moderator feedback case with manual control. For both of these cases, the minimum DNBR remains above k:SL / * ,4f gg ,
~5 & % alyJil 15.1-7 Amendment 43 l _ _ _ - _ _ _ _ _ _ _ _ - _ _ _ _ -_ _
p ,
STr ySAR PAGE A 01 56 1 For all cases, the plant rapidly reaches a stabilized condition at the higher power level. Normal plant operating procedures would then be followed to 2 l
reduce power.
The excessive load increase incident is an overpower transient for which the fuel temperatures will rise. Reactor trip is not assumed for the cases ana- $8 lysed, and the plant reaches a new equilibrium condition at a higher power level corresponding to the increased steam flow. If reactor trip were assumed, the analysis results would be less severe. [8
, Since DNB does act occur at any time during the ascessive load increase tran-sients, the ability of the primary coolant to remove beat from the fuel rod is not reduced. Thus, the fuel cladding temperature does not rise significantly ;
above its initial value during the transient. g gf '
- 4. g y 15.1.3.3 Conclusions. The analysis presente above shows that for a ten percent step load increase, the DNBR remains above 3790F thusj the DNB design l18 l basis as described in Section 4.4 is set. The plant reaches a stabilized condition rapidly following the load increase. -
15.1.4 Inadvertent Opening of a Steam Generator Relief or Safety Valve ,
causing a Depressurization of the Main Steam Systes 15.1.4.1 Identifiestion of Causes and Accident Description. The most severe core conditions for an accidental depressurisation of the main steam system result from an inadvertent opening of a single steam dump, relief, or safety valve. 'The analyses performed assuming a rupture of a main steam line are given in Section 15.1.5.
The steam release as a consequence of this accident results in an. initial increase in steam flow which decreases during the accident as the steam pres-sure falls. The increased energy removal from the RCS causes a reduction in coolant temperature and pressure. In the presence of a negative moderator temperatttre coefficient, the cooldown results in an insertion of positive j reactivity.
l The analysis is performed to demonstrate that the following criterion is sat-isfied: assusing a stuck rod cluster control assembly (RCCA), with offsite 2
power available, and assuming a single failure in the ESy, there is no conse-quantial damage to the core or reactor coolant system after reactor trip for a steam release equivalent to the spurious opening, with failure to close, of the largest of any single steam dump, relief, or safety valve.
Accidental depressurisation of the secondary systes is classified as an AES Condition Il event (see Section 15.0.1).
The following systems provide the necessary protection for an accidental depressurization of the main steam system; 2
- 1. Safety injection actuation from either:
- s. Two out of four low pressurizer pressure signals 15.1-8 Amendment 43 l
l l
L_______________.-.----.---_----_---_-_-_-_--.-_.----------
.PAGE32 or s-s- -
- 5. In computing the steam flow, the Moody Curve (Ref.15.13) for f(1./D) - 0 is used. 43
- 6. Perfect moisture separation in the steam generator is assumed.
Results The calculated time sequence of events for' this accident is listed in Table 15.1-1.
The results presented are a conservative indication of the events which would occur assuming a secondary system steam release since it is postulated that all of the conditions described above occur simultaneously.
Figures 15.1-12 and 15.1-13 show the transient results for a steam flow of 292 lbs/sec at 1300 psia from one steam generator.
The assumed steam release is typical of the capacity of any single steam dump, relief, or safety valve.
Safety injection is initiated by low pressurizer pressure. Boron solution at l1 2,500 ppa enters the RCS providing sufficient negative reactivity to maintain 3 significant margin to core design limits. The cooldown for the case shown on Figures 15.1-12 and 15.1-13 as a result of the conservative analysis is more l43 rapid than the case of steam release from all steam generators through one steam dump, relief, or safety valve. The calculated transient is quite con- l2 servative with respect to cooldown, since no credit is taken for the energy stored in the system metal other than that of the fuel elements or the energy stored in the other steam generators. Since the transient occurs over a peri-od of about five minutes, the neglected stored energy will have a significant effect in slowing the cooldown.
15.1.4.3 Radiological Consequences. There are only minimal radiological consequences from this event. The inadvertent opening of a single steam dump.
relief or safety valve can result in steam release from the secondary system. '3 Since no fuel damage is postulated to occur from this transient, the radiolog-ical consequence are less severe than the steam line break analyzed in Section 15.1.5.3.
15.1.4.4 conclusions. The analysis shows that the criteria stated ear- ,
lier in this section are satisfied. For an accidental depressurization of the Main Steam System, the minimum DNBR remains well above the" limit value and 3 no system design limits are exceeded.
.54% NJI 15.1.5 Spectrum of Steam System Piping Failures Inside and outside Con-tainment 15.1.5.1 Identification of causes and Accident Description. The steam release arising from a rupture of a main steam line would result in an initial increase in steam flow which decreases during the accident as the steam pres-sure decreases. The increased energy removal from the RCS causes a reduction l2 of coolant temperature and pressure. In the presence of a negative moderator
.. ww -. , .
15.1-11 Amendment 57
t Plent 15.0.3.chnrectoristics cnd initici canditiens era furchar discussed in Section 57 Except as discussed above. normal reactor control systems and ESF systems are not required to function. Several cases are presented in which pressurizer spray and power operated relief valves are assumed, but the more limiting cases wh*re these functions are not assumed are also presented.
The RTS may be required to function following a turbine trip. Pressurizer safety valves and/or steam generator safety valves may be required to open to maintain system pressures below allowable limits. No single active failure will prevent operation of any system required to function. A discussion of ATWT considerations is presented in Reference 15.2-2.
Results n e transient responses for a turbine trip from 102 percent of full power operation are shown for four cases: two cases for minimum moderator feedback and two cases for maximum moderator feedback (Figures 15.2-1 through 15.2 8).
For the minimum moderator feedback cases, the core has the least negative coderator coefficient of reactivity. For the maximum moderator feedback cases, the moderator temperature coefficient has its highest absolute value.
The calculated sequence of events for the accident is shown in Table 15.2-1.
Figures 15.2 1 and 15.2-2 show the transient responses for the turbine trip with minimum moderator feedback, assuming full credit for the pressurizer epray and pressurizer power operated relief valves. No credit is taken for 57 the turbine bypass. The reactor is tripped on the high pressurizer pressure oignal. The minimum departure from nucleate boiling ratio (DNBR) remains well cbove 7:RL The pressurizer safety valves are actuated and maintain primary 3g system pressure below 110 percent of the design value. The steam generator cafety valve s epoint. alves limit the secondary steam conditions to saturation at the safety N$r.Saftfy 4 d]1Ls I m A & -
Figures 15.2 3 and I5.2-4 show the responses for the turbine trip with maximum coderator feedback. All plant parameters are the same as the above and the reactor is tripped on the high pressurizer pressure signal. The DNBR 62 increases throughout the transient and never drops below its initial value.
Pressurizer relief valves and steam generator safety valves prevent I cverpressurization in primary and secondary systems, respectively. The pres-( curizer safety valves are not actuated for this case.
The turbine trip accident was also studied assuming the plant to be initially cperating at 102 percent of full power with no credit taken for the pressuriz-or spray, pressurizer power-operated relief valves, or turbine bypass. The 43 reactor 1:: tripped on the high pressurizer pressure signal. Figures 15.2-5 cnd 15.2 6 show the transients with minhum moderator feedback. The neutron flux remains essentially constant at 102 percent of full power until the reac-tor is tripped. The DNBR increases throughout G transient.' In this case the pressurizer safety valves are actuated and maintain system pressure below 110 percent of the design value.
Figures 15.2 7 and 15.2 8 show the transients with maximum moderator feedback with the other assumptions being the same as in the preceding case. Again, the DNBR increases throughout the transient, and the pressurizer safety valves cre actuated to limit primary pressure. -
- '.26 ;. m -
c ~-
i NGQO g g-S7p 734g
- ! - 4teference.15.2-4 presents additional results of analysis for a complete loss of heat sink including loss of main feedwater. This analysis shows the over.
pressure protection that is afforded by the pressurizer and steam generator safety valves.
15.2.3.3 Itadiolozical Consequences. There are only minimal radiological consequences associated with this event, therefore, this event is not limiting. The radiological consequences resulting from atmospheric steam dump are less severe than the steam line break event discussed in Section 15.1.5.
15.2.3.4 Conclusions. Results of the analyses, including those in Refer-ence 15.2-4, show that the plant design is such that a turbine trip without a direct or immediate reactor trip presents no hazard to the integrity of the RCS or the main steam system. Pressure relieving devices incorporated in the-two systems are adequate to limit the maximum pressures to within the design
& fafy kslopij MVakst The DNBR remains above M for all cases analyzed; thus, the DNB sesign bacis p as described in Section 4.4 is set. The above analysis demonstrates the ability of the NSSS to safely withstand a full load rejection.
15.2.4 Inadvertent closure of Main Steam Isolation valves The inadvertent closure of main steam isolation valves would cause a turbine trip and other consequences as described in Section 15.2.5 below.
15.2.5 loss of Condenser vacuum and other Events causing a Turbine Trip loss 'of condenser vacuum is one of the events that can cause a turbine trip.
Turbine trip initiating events are described in Section 10.2. A loss of condenser vacuum would preclude the use of turbine bypass to the condenser; '
however, since turbine bypass is assumed not to be available in the turbine trip analysis, no additional adverse effects would result if the turbine trip were caused by loss of condenser vacuum. .Therefore, the analysis results and conclusions contained in Section 15.2.3 apply to loss of condenser vacuum. In addition, analyses for the other possible causes of a turbine trip, as listed in Section 10.2, are covered by Section 15.2.3. Possible overfrequency 4 effects due to a turbine overspeed condition are discussed in Section 15.2.2.1 and are not a concern for this type of event.
l_
15.2 6a Amendment 54
PAGE3 [o & g ~-'
sTP rsAR Emerganc, 3... ting procedures following a main feedwater line rupture require 32 g the oper ar ec, isolate feedwater flow spilling from the ruptured feedwater Q211.
line and to control the RCS temperature which also prevents the pressurizer 74 from filling. 54 1
A block diagram summarizing various protection sequences for safety actions required to mitigate the consequences of this event is provided in Figure Q211.
15.0-13. 06 Plant 15.0.3.characteristics and initial conditions are further discussed in Section No reactor control systems with the exception of the pressurizer PORVs, are assumed to function. The operation of the PORVs serves to worsen the ,
transient via minimizing the saturation temperature and therefore minimizing 54 the margin to subcooling. It also allows a greater discharge of mass from the primary system, thus maximizing the liquid volume in the pressurizer. The RTS is required to function following a feedwater line rupture as analyzed here.
No single active failure will prevent operation of this system. .
only one auxiliary feedwater pump is assumed to function following receipt of an initiating signal. Following initiation, the auxiliary pump is assumed to deliver 540 gal / min of auxiliary feedwater to one intact steam generator. l 32 Results 32 Q211.
74 Calculated plant parameters following a major feedwater line rupture are shown on Figures 15.2-11 through 15.2 24 Results for the case with offsite power available are presented on Figures 15.2-11 through 15.2-17. Results for the case where offsite power is lost are presented on Figures 15.218 through 15.2-24 The calculated sequence of events for both cases analyzed is listed in Table 15.2-1.
The system response following the feedwater line rupture is similar for both cases analyzed. Results presented on Figures 15.2 12 and 15.2-15 (with offsite power available) and Figures 15.2 19 and 15.2-22 (without offsite power) show that pressures in the RCS and main steam system remain below 110 percent of the respective design pressures. Pressurizer pressure increases until reactor trip on low lov steam generator water level. Pressure then !
decreases, due to the loss of heat input. Coolant expansion occurs due to 43 reduced heat transfer capability in the steam generators; the pressurizer power operated relief valves open to maintain RCS pressure at an acceptable value.
k ,Sithdy M QL.
DNBR remains above MOSesatduring all tim &O 1 O M the transients, as shown on Fig- l18 ures 15.2 17 and 15.2-24; thus, the DNB design basis as described in Section 4.4 is set.
The reactor core remains covered with water throughout the transient, as water relief due to thermal expansion is limited by the heat removal capability of the AFW system.
The major difference between the two cases analyzed can be seen in the plots of hot and cold leg temperatures, Figures 15.2-13 and 15.214 (with offsite 15.2 16 Amendment 54
. m a..n. ,3m ,
[
Figurns 15.31 through 15.3-4 show the transient response for the loss of one rm :t3r coolant pump with four loops in operation. Figure 15.3-4 shows the l18 31 to be always greater than 39ME b ,$aftg w/p My Q, 54 Since DNB does not occur, the ability of the primary coolant to remove heat from the fuel rod is not significantly reduced. Thus, the average fuel and clad temperatures do not increase significantly above their respective initial values.
De calculated sequence of events is shown in Table 15.31. The affected h4 ;
reactor coolant pump will continua to coast down, and.the core flow will reach a new equilibrium value corresponding to the manber of pumps still in operation. With the reactor tripped, a stable plant condition will eventually be attained. Normal plant shutdown may then proceed.
15.3.1.3 Radiological Consequences. A partiel loss of reactor coolant flow from full lead would result in a reactor and turbine trip. . Assuming, in addition, that the condenser is not available, atmospheric steam dump may be required.
here are only minimal radiological consequences associated with this event.
Therefore this event is not limiting. ne radiological consequences resulting from atmospheric steam dump are less severe than the steam line break event analyzed in Section 15.1.5 since fuel damage as a result of this transient is not postulated.
15.3.1.4 Conclusions. The analysis shows that the DNBR will not decrease below _ at any time during the transient. Thus, the DNB design l1g The radiological c$ }*$l43iJAL h lL9albasis as descri ed in Sectio onsequenchs of this event are not limiting.
15.3.2 Complete Imss of Forced Reactor Coolant Flow 15.3.2.1 Identification of Causes and Accident Descriptio,n. A complete
' loss of forced reactor coolant flow may resvit from a simultaneous loss of -
electrical power to all reactor coolant pumps. If the reactor is at power at l43 the time of the. accident, the immediate effect of loss of coolant flow is a rapid increase in the coolant temperature. This increase could result in DNB with subsequent fuel damage if the reactor were not tripped promptly.
, Normal power for the reactor coolant pumps is supplied through buses from a
~
transformer connected to the generator. Yhen a generator, turbine, or reactor trip occurs, without an electrical fault, the generator circuit breaker auto- 43 matica11y opens and back-feed of off-site power occurs through the main trans-former and unit auxiliary transformer. Thus, the pumps will continue to sup-ply coolant flow to the core. l$4 ;
i 1
15.3-3 Amendment 54 i
I i
- undarvoltego or %derfraqu2ncy. Ons veriction batwaen this analysis end that of the previous sietion is that the RCCA insertion time to dashpot entry is 55 2.58 seconds. This is a conservative insertion time under the reduced flow conditions that exist when the RCCAs are inserted for this transient.
Results Figures 15.3 9 through 15.3-12 show the transient response for the loss of power to all raaetor coolant pumps with four loops in operation. The reactor is again assumed to be tripped on undervoltage sipal. Figure 15.3-12 shows theDNBRtobealwaysgreaterthanbuG&,H.g,_fk.g gpQ gg, l54 Since DNB does not occur, the ability of the primary coolant to remove heat from the fuel rod is not greatly reduced. Thus, the average fuel and clad temperature do not increase significantly above their respective initial values.
The calculate $ sequence of events is shown in Table 15.3-1. The reactor 54 coolant pumps will continue to coastdown, and natural circulation flow will eventually be established, as demonstrated in Section 15.2.6. With the reactor tripped, a stable plant condition will be attained. Normal plant shutdown may then proceed.
15.3.2.3 Radiological Consequences. A complete loss of reactor coolant flow from full load results in a reactor and turbine trip. Assuming, in addi-tion, that the condenser is not available, atmospheric steam dump would be required. The quantity of steam released would be the same as for a loss of offsite power.
There are only minimal radiological consequences associated with this event.
Therefore, this event is not Itaiting. Since fuel damage is not postulated, the radiological consequences resulting from atmospheric steam dump are less severe than the steam line break, discussed in Section 15.1.5.
15.3.2.4 conclusions. The analysis performed has demonstrated that for the complete less of forced reactor coolant flow, the DNBR does not decrease below at any time during the transient. Thus, the DNB design basis as di l18 descrie$nS.ection44issep.
Se hiky Wa$y.h'S AN W 15.3.3 Reactor' Coolant Pum)p Shaft Seizure (Locked Rotor) 15.3.3.1 Identification of causes and Accident Description. The acci-dent postulated is an instantaneous seizure of a reactor coolant pump rotor such as is discussed in Section 5.4 Flow through the affected reactor cool-ant loop is rapidly reduced, leading to an initiation of a reactor trip on a low reactor coolant flow signal. l'3 >
Following initiation of the reactor trip, heat stored in the fuel rods contin-ues to be transferred to the coolant causing the coolant to expand. At the same time, beat transfer to the shell side of the steam generators is reduced, first because the reduced flow results in a decreased tube side film coeffi-cient and then because the reactor coolant in the tubes cools down while the shell side temperature increases (turbine steam flow is reduced to zero upon turbine trip). The rapid expansion of the coolant in the reactor core, com-bined with reduced heat transfer in the steam generators causes an insurge 43
, 15.3 5 A m ndment 55
g( , l PAGEMOF ST Ssp rSAn
- u. .
Figures 15.41 shows the me,utron flux trar.ais a The neutron flux does not overthoot the nominal full power value.
The energy release and the fuel temperature increases are relatively small.
The thermal flux response, of interest for DNB considerations, is shown on Figure 15.4-2. The beneficial effect of the inherent thermal lag'in the fuel is evidenced by a heat flux such less than the full power nominal value.
There is a large margin to DNB during the transient since the rod surface heat ;
flux remains below the design value, and there is a high degree of subcooling at all times in the core. Figure 15.4-3 shows the response of the hot spot fuel and cladding temperature. The hot spot fuel average temperature increas-es to a value below the nominal full The min at all times remains above TpBE. &. power hot spot value. gig Q Vsdq The calculated sequence of events for this accident is shown in Table 15.41.
With the reactor tripped, the plant returns to a stable condition. The plant may subsequently be cooled down further by following normal plant shutdown )
procedures.
15.4.1.3 Radiologieel Consequences. There are no radiological conse-quences associated with an uncontrolled RCCA bank withdrawal from a sub- -
critical or low power startup condition event since radioactivity is contained 1 within the fuel rods and RCS within design limits.
15.4.1.4 conclusions. In the event of a RCCA withdrawal accident from the suberitical condition, the core and the RCS are not adversely affected, since the combination of thermal power and the coolant temperature result in a DNBR which is well above th limit value.oiatate. Thus, the DNB design 18 basis as described in Section .4 Js set.
15.4.2
.Jafg % %
- Uncontrolled Rod cluster control Afsembly Bank Withdrawal at Power 15.4.2.1 Identification of Causes and Accident Description. Uncen. I trolled RCCA bank withdrawal at power results in an increase in the core heat flux. Since the heat extraction from the steam generators legs behind the core power generation until the steam generator pressure reaches the relief or safety valve setpoint, there is a not increase in the reactor coolant tempera-ture. Unless terminated by manual or automatic action, the power mismatch and resultant coolant temperature rise could eventually result in DNS. Therefore, in order to avert damage to the fuel clad, the Reactor Trip System (RTS) is i e an 43 des gnedMkJ to terminajM
. @y such transient VA/444. . before the DNBR falls below EEE. 4(g, Thi e nt is flessified as an ANS Condition II incident (an incident of mod-erste frequency) as defined in Section 15.0.1.
The automatic features of the RTS which prevent core damage following the 1 43 postulated accident include the following: 1
- 1. Power range neutron flux instrumentation actuates a reactor trip if two out of four channels exceed an overpower setpoint. f l
1
- 2. Reactor trip is actuated if any two out of four AT channels exceed an overtemperature AT setpoint. This serpoint is automatically varied with axial power imbalance, coolant temperature and pressure to protect against DNS.
' 1 15.4 5 Amendment 57 i
't . 'wetoe trip is actuated if any two out of four AT channels exceed an
.,v 9.v.:.- AT se tpoint. This setpoint is automatically varied with axial power imbalance to ensure that the allowable heat generation rate (kW/f t) is not exceeded.
4 A high pressurizer pressure reactor trip is actuated if any two out of 43 four pressure channels exceed the setpoint. This set pressure is less than the set pressure for the pressurizer' safety valves.
- 5. A high pressurizer water level reactor trip is actuated if any two out of four level channels exceed the setpoint. l43 In addition to the above listed reactor trips, there are the following RCCA withdrawal blocks:
- 1. Eigh neutron flux (one out of four):
- 2. Overpower AT (two out of four); and,
- 3. Overtemperature AT (two out of four). -
The a anner in which the combination of overpower and overtemperature AT trips provide protection over the full range of RCS conditions is described in Chap-ter 7. Figure 15.0 1 presents allowable reactor coolant loop average tempera-tures and ATs for the design power distribution and flow as a function of pri-mary coolant pressure. The boundaries of operation defined by the overpower AT trip and the overtemperature AT trip are represented as " protection lines" cn this diagram. The protection lines are drawn to include all adverse instrumentation and setpoint errors so that under nominal conditions trip would occur well within the area bounded by these lines. The utility of this diagram is*in the fact that the limit imposed by any given DNBR can be repre-cented as a line. The DNB lines represent the locus of conditions for which the DNBR equals All points below and to the left of a DNB line for a given pressure ave a DNBR greater than e diagram shows that DNB is prevented for 1 cases if the area enclosed with t imum protection lines is not travers the a plicable Dh R .i ta cint. QQyg ,
The area of permissib1 op ration power, pressure, and temperature) is bounded by the combination of reactor trips: hi 5h neutron flux (fixed set-point); high pressurizer pressure (fixed setpoint); low pressurizer pressure '3 (fixed setpoint); overpower and overtemperature AT (variable setpoints).
15.4.2.2 Analysis of Effects and Consequences.
Method of Ar.alysis This transient is analyzed by the IDFTRAN code (Ref.15.4 3). This code simu-lates the neutron kinetics, RCS, pressurizer relief and safety valves, pres-curizer spray, steam generator, and steam generator safety valves. The code computes pertinent plant variables including temperatures. Tressures, and power level. The core limits as illustrated on Figure 15.0 1 are used as l input to IDFTRAN to determine the minimus DNBR during the transient.
l 15.4-6 Amendment 57
(
- _%m_
f!i .
. ST HL AE doa 1 -
l I
The transient response for a slow RCCA withdrawal from full pvver is shown on Figures 15.4-7.through 15.4-9. Reactor trip on overtemperature AT occurs
[ aft.eralonLerperiod. Again, the minimum DNBR.is greater than $332. M .J4(c/fyl8 u vA L
%g, Figure 15.4 10 shows the minimum DNBR as a function of reactivity insertion rate from initial full power operation for minimum and maximum reactivity feedback. It can be seen that two reactor trip channels provide protection over the whole range of reactivity insertion rates. These are the high neu-tron flux and overtemperature AT channels.
than-8:;sa. %lAwd4 Valu.t. The minimus DNBR is always greater l 1 J M ann 4 gge Figures 15.4-11 ,ind 15.4-17 show the minimum DNBR as a function of reactivity insertion rate for RCCA withdrawal incidents starting at 60 and 10 percent power respectively. The results are similar to the 100 percent power case, cxcept as the initial' power is decreased, the range over which the overten-parature AT trip is effective qs increased. In neither case does the DNBR fall below 3:;32 fbe g lAWf Vsut. l 18 Jehfy M*
The shape of the curves of minimuf DNBR versus reactivity insertion rate in the referenced figures is due both to reactor core and coolant system tran-cient responsac and to protection system action in initiating a reactor trip.
Referring to the minimum feedback case in Figure 15.411, for example, it is noted that:
- 1. For high reactivit Ak/see and 1.0x10'y insertion rates (i.e., between approximately 3 x 10 62 Ak/sec) reactor trip is initiated by the high neutron flux trip. .The neutron flux level in the core rises rapidly for these insertion rates while core heat flux and coolant system temperature lag behind due to the thermal capacity of the fuel and coolant system fluid.
Thus, the reactor is tripped prior to significant increase in heat flux or water temperature with resultant high minimum DNBRs during the transient. As reactivity insertion rate decreases, core heat flux and coolant temperatures can remain more nearly in equilibrium with the neutron flux; minimum DNBR during the transient'thus decreases with decreasing insertion rate.
- 2. The overtemperature AT reactor trip circuit initiates a reactor trip when measured coolant loop AT exceeds a setpoint based on measured RCS a.retage temperature and pressure. This trip circuit is described in detail in Chapter 7; however, it is important in this context to note that the
- average temperature contribution to the circuit is lead lag compensated in order to decrease the effect of the thermal capacity of the RCS in
-response to power increases.
- 3. With further decrease in reactivity insertion rate, the overtemperature AT and high neutron flux trips become equally g effective in terminating ;
the transient (e.g., at approximately 3 x 10 Ak/sec reactivity '
insertion rate). k2 For reactivity insert approximately5x10'jonratesbetweenapproximately3x10Ak/secand Ak/see the effectiveness of the overtemperature AT 62 trip increases (in terms of increased minimum DNBR) due to the fact that ,
with lower insertion rates the power increase rate is slower, the rate of j
, 't <
15.4-8 Amendment 62
m . . ._
. ST HbAE. 302/
_Pyl y g L
L ,
+'
15.4.2.4 Conclusions. The high neutron flux and overtemperature AT tri p channels provide adequate. prot % tion over the entire range of possible reae p l62 tivity insertion' rates (i.e.s the minimum value of DNBR is always larger than 3540). Thus, the DNB desig. basis as described'in Section 4.4 is met. b3 N.5<4LhY $%$Q } ) YC Ydh4,L
~ 15.4.3 - Rod cluster Control Assembly Misoperation h3
' 15.4.3.1' identifiestion of Causes and Accident Description. RCCA mis.
cperation accidents include:
1.. 30 One or more dropped RCCAs within the same group;
- 2. A dropped RCCA bank;
- 3. Statically misaligned RCCA; h3 L
'4. Withdrawal ~of a single RCCA.
p0 Each RCCA has's position indicator channel which displays the position of' the assembly. The displays of assembly positions are grouped for the operator's convenience. Fully inserted assemblies are further indicated by a rod at bottom signal, which actuates a local alarm and a control room annunciator.
Group demand position is also indicated.
'j RCCAs are always moved in preselected banks, and the banks are always moved in l30 the same preselected sequence.
i i
The tors.
rods comprising a group operate in parallel through multiplexing thyris-Each 53 The two groups in a bank move sequentially such that the first group is always within one step of the second. group in the bar.k. A definite schedule of actuation or deactuation of the stationary gripper, movable gripper, and lift coils of a mechanism is required to withdraw the RCCA attached to the techanism. Since the stationary gripper, movable gripper, and lift coils cssociated with the four RCCAs of a rod group are driven it. parallel, any 43 53 single failure which would cause rod withdrawal would affect a minimum of one group. Mechanical' failures are in the direction of insertion or immobility.
The dropped RCCA, dropped RCCA bank, and statically misaligned RCCA events are l53 classified as American Nuclear Society (ANS) Condition II incidents (incidents af moderate frequency) as defined in Section 15.0.1. However, the single RCCA withdrawal incident is classified as an ANS Condition III event, as discussed b31ow.
Na single electrical or mechanical failure in the rod control system could l43 ccuse the accidental withdrawal of a single RCCA from the inserted bank at
~ full power operation. . The operator could withdraw a single RCCA in the centrol chould one bank besince this feature accidentally is necessary dropped. The eventinanalyzed order to must retrieve an from result assembly l43 53 ng failures (probability for single random failure is on the
'order multiple of 10~wirj/ year; refer to section 7.7.2.2) or multiple serious operator orrors and subsequent and repeated operator disregard of event indication. l 57 The probability of such a combination of conditions is very low. The consequences, however, may include slight fuel damage. Thus, consistent with _
53 15.4 10 Amendment 62 a.
m,,,,,,,,
. ST HL.AE 3021
. , PAGE 4,2. OF 55" STP FSAR curves during the first part.of the transient, the increase in .: ora flow with cooler water results in an increase in nuclear power and a decrease in core average temperature. The minimum NBR during the transient is considerably Areater chan 3:;48: M Mg d}.h hJd Mut.
Reactivity addition for the inactive loop startup accident case is due to the decrease in core water tangerature. During the transient, this decrease is
-dua both to the increase in reactor coolant flow and, as the inactive loop flow reverses, to the cooler water entering the core from the hot leg side (colder temperature side prior to the startup of the transient) of the steam generator in the inactive loop. Thus, the reactivity insertion rate for this transient changes with time. The resultant core nuclear power transient, computed with consideration of both moderator and Doppler reactivity feedback effects, is shown on Figure 15.4-16.
The calculated aequence of events for this accident is shown in Tabis 15.4 1.
The transient results illustrated on Figures 15.4-16 through 15.4 20 indicate that a stabilized plant condition, with the reactor tripped, is approached at 30 seconds. Plant cooldown may subsequently be achieved by following normal ahutdown procedures.
15.4.4.3 Radiological Consequences. There are only minimal radiological consequences associated with startup of an inactive reactor coolant loop at an incorrect temperature. Therefore, this event is not limiting. The reactor trip causes a turbine trip and heat is removed from the secondary system through the steam generator power relief valves or safety valves. Since no fuel damage is postulated to occur from this transient, the radiological con-sequences associated with this event are less severe than the steam line break event, as discussed in Section 15.1.5.
f g,,
15.4.4.4 Conclusions. The transient esults show that the core is not adversely affected. The DNBR remains above throughout the transient; 43 thus, the DNB design basis as described in Section 4.4 is set.
15.4.5 A Malfunction or Failure of the Flow Controller in a SWR Loop That Results in an Increased Reactor Coolant Flow Rate Not applicable to South Texas.
15.4.6 Chemical and Volume Control System (CVCS) Malfunction That Results in a Decrease in Boron Concentration in the Reactor Coolant 15.4.6.1 Identification of Causes and Accident Description. Reactivity l can be added to the core by feeding unborated water into the RCS via the 1 57 CVCS. Boron dilution is a manual operation under strict administrative controls with procedures calling for a limit on the rate and duratien of dilution. A boric acid blend system is provided to permit the operator to match the boron concentration of reactor coolant makeup water during normal i charging to that in the RCS. The CVCS is designed to limit the potential rate !
sf dilution to a value which, after indication through alarms and instrumentation,.provides the operator sufficient time to correct the cituation in a safe and orderly manner.
.. 6 15.4 18 Amendment 57
15.6.1.2 Analysis of Effects and consequences. The accidental depressurization transiewt is analyzed by employing the detailed digital com-puter code IDITRAN (Ref.15.6-1). The code simulates the neutron kinetics, RCS, pressurizer, pressurizer relief and safety valves, pressurizer spray,
-steam generator (SG), and SC safety valves. The code computes pertinent plant variables including temperatures, pressures, and power level.
Plant characteristics and initial conditions are discussed in Section 15.0.3.
In order to give conservative results in calculating the departure from nucle-cte made. boiling ratio (DNBR) during the transient, the following assumptions are 1.
Initial conditions of maximum core power and reactor coolant temperature
(+4.7'T uncertainty) uncertainty) and minimum reactor coolant pressure (-46 psi are assumed.
'this results in the minimum initial margin to U 62 departure from nucleate boiling (DNB) (see Section 15.0.3).
- 2. A least negative moderator temperature coefficient of reactivity is assumed.
The spatial effect of void due to local or subcooled boiling is not core power shape.
considered in the analysis with respect to reactivity feedback et 3.
A large (absolute value) Doppler coefficient of reactivity is assumed (Fig.15.0 2) so that the resultant amount of positive feedback is con-servatively tivity feedback, high. This retards any power decrease due to moderator reac-Normal reactor control systems are not required to function. The rod control i system is assumed to be in the automatic mode in order to hold the core at full power longer and thus delay the trip. This is a worst-ease assumption; if the reactor were in manual control, an earlier trip could occur on low pressurizer pressure. The RTS functions to trip the reactor on the appropri-ete signal. No single active failure will prevent the RTS from functioning 45 properly.
Ensults The system response to an inadvertent opening of a pressurizer safety valve is chown on Figures 15.6-1 through 15.6-3. Figure 15.6-1 illustrates the nuclear .
p ver transient following the depressurization. Nuclear power is maintained et the initial value until reactor trip occurs on overtemperature AT. The pressure d3nt are given decayon transient and average Figure 15.6 2. Pressuretemperature transient following the acci.
drops more rapidly after core heat gsneration is reduced via the trip, and then slows once saturation temperature is reached in the hot leg. The DNBR transient is shown on Figure 15.6-3; DNBR 45 rcmains above M90 throughout the trgnsient.
A da M ya#4*p U /4.i+ VAA4C The calculated sequence of events for the inadvertent opening of a pressurizer safety valve incident is shown in Table 15.61.
15.6.1.3 Radiological consequences. An inadvertent opening of a pres.
curizer safety or relief valve releases primary coolant to the pressurizer rolief tank; however, even assuming a direct release to the containment atmos-phare, the radiological consequences of this event would be substantially less 3.b . -
naendment 62
Question 440.62N Provide the values of the moderator and Doppler coefficients of reactivity used in the loss of normal feedwater/ loss of offsite power analysis and verify their conservatism.
Response
h Yb.
. ram 4 s The acceptance criteria used for these tr sients are a) i thy pressurizer was not permitted to become water solid and b the minimum departure from nucleate
. o boiling ratio (DNBR) must be graater thanY1:96: Thus, the moderator and Doppler coefficients of reactivity were selected to yield conservative re-suits. This was done by selecting the moderator and Doppler coefficients of reactivity to maximize core power which in turn maximizes the volume expan-sion. The pertinent coefficients are listed in Table 15.0 2.
- l
, s.c c = c r.:
Vol . 2 Q6Jt 15.2 1N Amendment 49
UUU Ut1M F "*N " " J l Qu*<stien 440.62N 1 i
Provide the values of the moderator and Doppler coefficients of reactivity used in the loss of normal feedwater/ loss of offsite power analysis.*and verify their conservatism. '
i
Response
M (M The acceptance criteria used for these tr sients Ns-b & pressurizer are a) tho'J h was not permitted to become water solid and b the minimum departure from nucleate boiling ratio (DNBR) must be greater than 2=Nh. Thus, the moderator and Doppler coefficients of reactivity were selected to yield conservative re-sults. This was done by selecting the moderator and Doppler coefficients of reactivity to maximize core power which in turn maximizes the volume expan-sion. The pertinent coefficients are listed in Table 15.0 2.
1 a
Jol, s Q&R .15.2 .A *- 49 1
l ATTACHMENT
' AGE 6oyf,E l-
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ATTACHMENT
. ST HL AE W '
PAGE47 OF 5"6' STP FSAR 7.2.1.2.1 Cenerating Station Conditions: Generating station conditions requiring a reactor trip are the following:
- 1. DNBR approaching the design basis limit (Chapters 4 and 15).
- 2. Power density (kW/ft) approaching rated value for ANS Condition 11 faults (see Chapter 4 for fuel design limits). .
- 3. Reactor Coolant System (RCS) overpressure creating stresses approaching the limits specified in Chapter 5.
7.2.1.2.2 Cenerating Station Variables: The following variables are required to be monitored to provide reactor trips (see Table 7.2 1):
- 1. Neutron flux
- 2. Reactor coolant temperature
- 3. RCS pressure (pressurizer pressure)
- 4. Pressurizer water level
- 5. Reactor coolant flow
- 6. RCP operational status (voltage and frequency)
- 7. Steam generator water lesel (density compensated) 43
- 8. Turbine generator operational status (trip fluid pressure and stop valve position) 7.2.1.2.3 Spatially Dependent Variables: The reactor coolant tempera-rure measurement is the only spatially dependent variable. l54 43 7.2.1.2.4 Limits, Margins, and Setpoints: The parametric values that vill require reactor trip are given in Chapter 15. Chapter 15 proves that the cetpoints to be used in the Technical Specifications are conservative.
The setpoints for the various functions in the RTS were analytically deter-eined so that the operational limits so prescribed will prevent fuel rod clad damage and loss of integrity of the RCS as a result of any ANS Condition 11 incident (anticipated malfunction). As such, during any ANS Condition II incident, the RTS limits the following parameters to:
.1. Minimum DNBR - esS6 'bes,in bir b?A (a dIwaserl m Mew q.4g l43
- 2. Maximum system pressure - 2,750 psia
- 3. Fuel rod maximum linear power for determination of protection setpoints -
18.0 kW/ft 9
7.2 14 Amendment 54
7.
S -HL C gl VAGE4b' OF 35 '
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I Al I ACHMENT
. ST HL AE 30,21 2.2 2AvETY LIMITS RGE U OF SS WaSES 2.1.1 REACTOR CORE The restrictions of this Safety Limit prevent overheating of the fuel and possible cladding perforation which would result in the release of fission products to the reactor coolant. Overheating of the fuel cladding is prevented ,
by restricting fuel operation to within the nucleate boiling regime where the heat transfer coefficient is large and the cladding surface temperature is slightly above the coolant saturation temperature.
Operation above the upper boundary of the nucleate boiling regde coul'd result in excessive claddin .
from nucleate boiling (DNB)g and temperatures because the resultant sharpofreduction the onsetin ofheat departure transfer coefficient. DNB is not a directly seasurable parameter during operation and '
m therefore THERFAL POVER and reacter coolant temperature and oressurt have been r= ~
related to Dhe through thT1 h 1 correlation.
tion has been developed to predict the DNB flux and the location of DNB forTheIWT- '
axially uniform and nonuniform heat flux distributions. The local DNS heat flux ratio (DNER) is defined as the ratio of the heat flux that would cause DNE at a particular core location to the local heat flux and is indicative of the margin to DNB.
The mi " mum val of h Dh h uring ady-st e ope ion, kb ope 'ional tr sients, nd an
- ipa d trans nts is ' mite o 1.3rmal . Th \is g, value c responds e a 95. robab' ity 95% nfiden e eve at 0 -
w n o ur and chosen an a repr sarg to D a ra Ln The curves of Figure 2.1-1 show the loci of points of THERFAL POWER, Reactor Coolant System ressure and average temperature for which the minimum DNBR is no less tha , or the average enthalpy at the vessel exit is equal to the enthalpy aturated liquid.
h Th ese curves are based on an enthalpy het channel factor, FgN , of 1.52 and a reference cosine with a peak of 1.61 for a.xial power shape. An allowance is included for an increase in Fh at reduced power based on the expression:
Fh=1.52[1+0.3(1-P))
Where P is the fraction of RATED THERMAL PDWER.
These limiting heat flux conditions'are higher than those calculated for the range of all control rods fully withdrawn to the maximum allowable control rod insertion assuming the axial power imbalance is within the limits of the f3 (AI) function of the Overtemperature trip. When the axial power imbalance is not within the tolerance, the axial power imbalance effect on the Over-temperature AT trips will reduce the Setpoints to provide protection consistent with core Safety Limits. ,
fhM7 SOUTH TEXA5 - UNIT 1 B I-1
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QWACHMENT ST HL AE 30.2/
PAGESCOF c5~
i Insert A The minimum value of DNBR during steady state, normal operational transients, and anticipated transients is limited to '
1.17, the DNBR design limit of the WRB-1 correlation. This value corresponds to a 95% probability at a 954 confidence level that DNB will not occur. Margin is maintained by meeting a DNBR value ;
of 1.27 in the safety analyses. i
]
l i
C_________________-__- - - - - - - - - - - - - - - - - - - - - - " - - - - - - - - - ' ' - - ' - - -
- @LHL.AE. Jn2I PAGE 6/ OF 5"C LIMITING SAFETY SYSTEM SETTINGS BASES REACTOR TRIP SYSTER INSTRUMENTATION SETPOINTS (Continued)
The various Reactor trip circuits automatically open the Reactor trip breakers whenever a condition monitored by the Reactor Trip System reaches a -
preset or calculated level. In addition to redundant channels and trains, the design approach provides a Reactor Trip System which monitors numerous system variables, therefore providing Trip System functional diversity. The functional capability at the specified trip setting is required for those anticipatory or I diverse Reactor trips for which no direct credit was assumed in the safety analysis to enhance the overall reliability of the Reactor Trip System. The Reactor Trip System initiates a Turbine trip signal whenever Reactor trip is initiated. This prevents the reactivity insertion that would otherwise result . 1 from excessive Reactor Coolant System cooldown and thus avoids unnecessary actuation of the Engineered Safety Features Actuation System.
The Reactor Trip System includes manual Reactor trip capability.
Power Ran;e, Neutron Flux -
In each of the Power Range Neutron Flux channels there are two independent bistables, each with its own trip setting used for a High and Low Range trip setting. The Low Setpoint trip provides protection during suberitical and low power operations to mitigate the consequences of a power excursion beginning from low power, and the High Setpoint trip provides protection during power operations to mitigate the consequences of a reactivity excursion from all power levels.
The Low Setpoint trip may be manually blocked above P-10 (a power level of approximately 10% of RATED THERMAL POWER) and is automatically reinstated below the P-10 Setpoint.
Power Range, Neutron Flux, Hioh Rates The Power Range Positive Rate trip provides protection against rapid flux increases which are characteristic of a rupture of Id control rod drive housing.
Specifically, this trip complements the Power Range Neutron Flux High and Low trips to ensure that the criteria are set for rod ejection from sid-power.
The Power Range Negative Rate trip provides protection for control rod drop accidents. At high power a single or multiple rod drop accident could cause local flux peaking which could cause an unconservative local DNBR to exist. The Power Range Negative Rate trip will prevent this from occurring by tripping the reactor. No credit is taken for operation of the Power Range Negative Rate trip for those control rod drop accidents for which DNBRs will be greater than i p des.y l;s+
SOUTH TEXA5 - UNIT 1 B 2-4 [f[)W
~ ' _ _ _ _ _ _
es-nvre. g o FhGE SLOF s f POSER D15TRfBUT20N LIMITS E!.SE S
. NEAT FLUX HDT CHANNEL FACTOR and NUCLEAR ENTHALPY RISE HOT CHANNEL FACTOR (Continuen) -
- c. The control rod insertion limits of Specifications 3.1.3.5 and .
3.1.3.6 are maintained; and
- d. The axial power distribution, expressed in tems of AXIAL FLUX DIFFERENCE, is maintained within the limits.
i Ffg will be maintained within its limits provided Conditions' a. through -
- d. above are saintained. The combination of the RCS flow requirement (395,000 ppm)andtherequirementonFfg guarantees that the DNBR used in the safety analysis will be met. -The relaxation of F"g as a function of THERMAL p0'*ER allows changes in the radial power shape for all permissible rod inser-tien limits.
N When F3g is measured, no additional allowances are necessary prior to comparison with the limit. A measurement error of 4% for F has bu n allowed for in the determination of the design DNBR value. H
\ Fue rodbokgred es the alueokDNSrat'. Cr dit iskvailable to offset thi reduct An in t gener sargik The neric argins, totaling.
s i
3.3% h5R cc letely ffset ny red ow pena ties. is ma in inc ' des the s i folloaing: N r,
Nflace
- a. esign imitDNKof1.Ovs1.8, N
. Gr) Spaci (K,) .059 s D.06 and
- c. Them Diffu *on Coef cient for us in mod fed spater fac r) f 0.05 vs 0.0 ,
s
~
The a niicabl values f rod . w pena its a expla ed in AR Sec- .
tion 4. 2.2.5.
e SOUTH TEXA5 - UNIT 1 B 3/4 2-4 p[{h
. 3/4.4 ' REACTOR COOLANT SYSTEM gg .
BASES b 3/4.4.1 REACTOR COOLANT LOOPS AND COOLANT CIRCULATION The plant is designed to ope ith all reactor coolant loops in operation and maintain DNBR above during all normal operations and
. anticipated transients. In MODES I and 2 with ont. reactor coolant loop not in operation this specification requires that the plant be in at least tiDT STANDBY within 6 hours6.944444e-5 days <br />0.00167 hours <br />9.920635e-6 weeks <br />2.283e-6 months <br />.
In MODE 3. two reactor coolant loops provide sufficient heat removal capability for~ removing core decay heat even in the event of a bank withdrawal accident; however a single reactor coolant loop provides sufficient heat removal capacity If a bank withdrawal accident can be prevented, i.e,., by opening the Reactor Trip System breakers. Single faHum considerations '
require that two loops be OPERABLE at a11 times.
In MODE 4, and in MODE 5 with reactor coolant loops filled, a single '
reactor coolant loop or RHR loop provides wffiefent heat removal capability for removing decay heat; but single faGure considerations require that at least two loeps (either RHR or RCS) be OPERABLE.
In MODE 5 with reactor coolant loops not filled, a single RHR loop provides sufficient heat removal capability for removing decay heat; but single failure considerations, and the unavailability of the steam generators as a heat removing component, require that at least two RHR loops be OPERABLE.
The boron dilution analysis assumed a common RCS volume, and maximum di-lution flow rate for MODES 3 and 4, and a different volume and flow rate for MODE 5. The MDDE 5 conditions assumed limited mixing in the RCS and cooling with the RHR system only. In MDDES 3 and 4 it was assumed that at least one reactor coolant pump was operating. If at least one reactor coolant pump is not operating in MDDE 3 or 4, then the maximum possible dilution flow rate must be limited to the value assumed for MDDE 5.
The operation of one reactor coolant pump (RCP) or one RHR pump provides adequate flew to ensure mixing, prevent stratifi;ation and produce gradual reactivity changes during boron concentration reductions in the Reactor Coolant System. The reactivity change rate associated with boron reduction will, therefore, be within the capability of operator reccpnition and control.
The restrictions on starting an RCP with one or more RCS cold legs less than or equal to 350*F are provided to prevent RCS pressure transients, caused by energy additions from the Secondary Coolant System, which could exceed the limits of Appendix G to 10 CFR rart 50. The RCS will be protected against overpressure transients. and will not exceed the Itaits of Appendix G by restricting starting of the RCPs to when the secondary water temperature of each steam generator is less than 50*F above each of the RCS cald leg temperatures.
- ~
3/4.4.2 SAFETY VALVES -
The pressurizer Code safety valves operate 19 prevent the RCS from being pressurized above its Safety Limit of 2735 psig. Each safety valve is designed to relieve 504,950 lbs per hour of saturated steam at the valve setpoint of 2500 psis. The relief capacity of a single safety valve is adequate to relieve any overpressure condition which could occur during shutdown. In the event 4 that no safe' / ~9v Tre MRABI.E. an operating RHR loop, connected to the
$0UTH TEXAS - UNIT 1 8 3/4 4-1 j [ hchO
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mrn, a w--
F PAGE OF S5" POUER D25TRIBUTION LIMRTS EASES 3/4.2.5 DNB PARAMETERS (t.ontinued) initial F5AR ersumptions,pn h ve been analytically demonstrated adequate to maint > 4 a Einizac) Dh58 c7 9 throughout each analyzed transient. The ,
ind'. :.d T,yp value of 59B*F and the indicated pressurizer pressure value of 22r. pig are pro >ided assuming that the readings from four channels will be averaged before comparing with the required lisit. The flow requirement ,
(39f4,000 gpm) includes a measurement uncertainty of 3.5L ,
The 12-hour periodic surveillance of these parameters through instrument rendeut is sufficient to ensure that the parameters are restored within their limits following load changes and other expected transient operation.
1 bb (f ( ( l b 8 SOUTH TEXAS - UNIT 1 B 3/4 2-6
{g()@{
("'
~
cHACHE1ENT
. ST HL.AE. 3021 PAGE G OFf f I Insert B Fuel rod bowing reduces the value of DNB ratio. Margin has been maintained between the DNBit value used in the safety analyses (1.27) and the design limit (1.17) to offset the rod bow penalty and other penalties which may apply.
4
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