ML20082U269
| ML20082U269 | |
| Person / Time | |
|---|---|
| Site: | McGuire, Mcguire  |
| Issue date: | 12/12/1983 |
| From: | DUKE POWER CO. |
| To: | |
| Shared Package | |
| ML20082U243 | List: |
| References | |
| NUDOCS 8312160289 | |
| Download: ML20082U269 (500) | |
Text
ATTACHMENT 1
SUMMARY
OF AFFECTED TECHNICAL SPECIFICATIONS Section Description Reason for Change 2.1.1 Reactor Core Safety Limits New Core Limits for Optimized Fuel 2.2.1 Reactor Trip System Setpoints New Core Limits (OTAT and OPAT only) 3/4.1.1.1 Shutdown Margin - Tavg > 200*F Reduced Shutdown Margin 3/4.1.1.3 Moderator Temperature Positive Moderator Temperature Coefficient Coefficient 3/4.1.3.4 Rod Drop Time Longer Rod drop time (bounding value) used in the analysis 3/4.1.3.6 Control Rod Insertion Hardware Limitations / Inappropriate Limits Conversion Correlation 3/4.2.1 Axial Flux Difference Relaxed Axial Offset Control (RAOC)
Limits 3/4.2.2 Heat Flux Hot Channel New LOCA Analysis for Optimized Factor - Fq(Z) Fuel and Inclusion of Fq Surveil-lance 4
3/4.2.3 RCS Flow Rate Minimum Measured Flowrate Assumed with the Improved Thermal Design Procedure 3/4.5.1.1 1.ccumulators - Cold Leg New LOCA analysis for OFA Injection 3/4.5.4 Boron Injection System - Boron' Injection Tank Deleted BIT 6.9.1.12 Radial Peaking Factor Limit Fq Surveillance Report a
.P. -
( .
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fity.fi.'in:lidi :tifts tippii! i ppin Hi! !i iriitu;i m iii ; !!pyqggniti .!!!ifrij gi:ph 0.2 0,4 0.6 0.8 1.0 1
(- FRACTION OF RATED THERMAL POWER Re[a.c.f. Wib FIGURE 2.1 1 REACTOR CORE SAFETY LIMIT FOUR LOOPS IN OPERATION fi nrg, . .
5 McGUIRE - UNITS 1 and 2 2-2
660< -
Flow 8er Loop = 98.400 gom 655 -
650<
645- Unacceptacle Oceration E400 .
655-650- # as t, 1
k625 -
o a 620 .
615 ay 8 t t, I
93g #
610 # /*
605 \
Acceptable 600 Operation 595
, a 500 -
585
- 0. .I .2 .5 .t .5 .6 .7 .8 9 1. 1.t :.2 POVER ffracticn or nominell Figure 2.1-1 =
4eactor Core Safety Limit Four Locos in Joeration' ( ui.4 1)
{
s
(^
's NA'.<e t.14% 1 c.4 ~4-
~2. - 2 r ., *
. 1 665, ow er Loop = 95,500 g p 660 655 'l 240g ###8 650 645 " ***
- 22S0 Operation Ds 1, 640<
655<
656< 2000 Ds1g 625-y I900 Dsig
- 623<
,I615-
- 610 625 622 595'
- Acceptable Operation 590 555<
SBa<
575<
578 ,
- 8. .I .2 .5 4 .5 .6 .7 .8 9 1. 1.1 1.2 POWER t react. ion of nominell Figure 2.1-1 b Unit 2 Reactor Core Safety Limit - Four Locos in Operation l
t McGUIRE UNITS 1 and 2 0 W /*f8-2-2a
TABLE 2.2-1 j 2
n -
S REACTOR TRIP SYSTEM IN51RilMENI ATION TRIP SETPolNTS
\ r%
. FilNCTIONAL UNIT TRIP SETPOINT ALLOWA8LE VALUES N.A.
k 1. Manual Reactor Trip H.A.
d Low Setpoint - 125% of RATED Low Setpoint - 126% of RAIED g 2. Power Range, Neutron flux TilERMAL POWER TifERMAL POWER a H!gh Setpoint - $ 109% of RATED High Setpoint - 1 110% of RATED u THERMAL POWER THERMAL POWER
- 3. Power Range, Neutron Flux, < 5% of RATED THERMAL POWER with < 5.5% of RATED THERMAL POWER High Positive Rate a time constant > 2 seconds Ulth a time constant > 2 seconefs
- 4. Power Range, Neutron Flux, 1 5% of RATED TilERMAL POWER with 5 5.5% of RATED TilERMAL POWER High Hegative Rate a time constant 1 2 seconds with a time constant 1 2 seconds ,
J, 5. Intermediate Range,. Neutron i 25% of RATED TilERMAL POWER 1 30% of RATED THERMAL POWER flux 5
- 6. Source Range, Neutron Flux 1 10 counts per second 1 1.3 x 10 counts per second
- 7. Overtemperature aT See Note i See Note 3
- 8. Overpower AT See Note 2 See Note 3
- 9. Pressurizer Pressure--Low 1 1945 psig 1 1935 ps!g
- 10. Pressurizer Pressure--High 5 2385 psig i 2395 psig II. Pressarizer Water Level--High 1 92% of instrument span S 93% of instrument span
- 12. Low Reactor Coolant Flow 1 90% of design flow per loop
- 1 89% of design flow per loop *
- Design flow is gpa per loop, foo Unit 1. .
and 95,5eojpe per l..p for Uc;t 2.
- . . . - _ . . - . . _ - . . - . - . - - - . _ - - - - - . - . - = . - . - . . . _ _ - -
N TABLE 2.2-1'(Continued) g REACTOR TRIP SYSTEM INSTRUMENTATION TRIP SETPOINTS
. i + '?: s NOTATION E + Tus y NOTE 1: OVERIEMi'ERATURE AT
't N
g p 1 1 + tjS 1
$ I l+t S)
I AT (1 + t )IA o I 1 2
b5 I S}~
3(P-P') - f 3(AI)}
" l + r5s Whare: AT =
Measured AT by RTD Manifold Instrumentation,
-[ 3, 1
=
Lag compensator on measured AT, I*wi 3 '
oT; 'V3 : 7. % erg uw,9s a g 2.)
n t-9 t =
Time constants utilized in the lag compensator for aTg - 2 ;;;.
=
to AT, Indicated AT at RATED THERMAL POWER, l 5 Kg i 1.0952g ( W + N y I4 60lC ""' 'k K =
2 0.0133, (.WW 2.), o.o n~2. ( u-!i :);
I+t jy S
=
3, 3 The function generated by the lead-Idg controller for T,yg dynamic compensation, Ag-u I
fg y
I g
=
Time constants utilized in the lead-lag controller for T,yg, 4 ig= 4 sec( w,.+3 i i 2.)j --
T = Average temperature, 'F,
=
y, Lag compensator on measured T,yg, 1 =
4 Time constant utilized in the measured T,yg lag compensator, t = 2 sec( ti 4s I $1.),
T' 5 588./'t Heterence T,yg at RATED THERMAL POWER, K
3 X( 1)/ *" # 0W ( "
D D
O Q i
TABLE 2.2-1 (Continued)
N REACTOR TRIP SYSTEM INSTRUMENTATION TRIP SETPOINTS 8
"o m NOTATION (Continued) e E NOTE 1: (Continued)
"4 v
H' P = Pressurizer pressure, psig, ,
P' = 2235 psig (Nominal RCS operating pressure),
" -1 S = Laplace transform operator, sec ,
and fg(AI) is a function of the indicated difference between top and bottom detectors of the power-range nuclear ion chambers; with gains to be selected based on measured ;
instrument response during plant startup tests such that: *4.0 (o (%M N,i 4f.o T.(ts.".+ 1.), -<H 4. a.d (i) for q g qb between -36% and % f (AI) j = 0, where qg and q are percent b
to RATED THERMAL POWER in the top andjbottom halves of the core respectively, and qt*Ab is total THERMAL POWER in percent of RATED THERMAL (u.M t.), POWER;
-a41 af.(u.My th (ii) foreachpercentthatthemagnitud{gf t "b exceeds -36%g e AT Trip Setpoint of shall be automatically reduced by 0.0L,.L u.a (its value at RATED 4 3.sst1 04.h 0THERMAL POWER; and
.(iii) foreachpercentthatthemagnitudgg,fpg-q . b exceeds +g the AT Trip Setpoint shall be automatically reduced by 4H. MHdpf its value at RATED THERMAL POWER.
~ ~
("
Qu.".t 2.), IM'D fe (""3 O > >
lt %1
._ : Le.J L.3 c. g. sd .- o m me.wed er#
LseA h I+% 5
" % e. .sh.ds dLb.e. sk h \e.d leic3 75I 7'L .
' Co wic de e- k e- 6T, T, ? 3 ec , i t -
3 Sec.j ua3 n)*
Q - o w., 7, . o w Cu-a Q;
TABLE 2.2-1 (Continued)
-g REACIOR TRIP SYSTEM INSTRUMENTATION TRIP SETPOINTS h N01ATION (Continued)
?.
, i!OIE ": CVI:l P0 LIER AT E 1 Tf 1 1 Il + r 5) 1 AT, { K4 -K 5(1+sgS)I l+t ) T -K6 U(1 + r ) - T"] - f 2(OI)I
, ':lhere: AT = As defined in Note 1, o .
t + TSh h = As defined in Note 1[ I
- 7IS _ A5 A -(. 4 I4 Mde l 1 + r4
. .ms 1
I + T.5) s T y
= As defined in Note 1,
= @S de.S .s e L 9gg A I, = As defined in Note 1, Q
\
sy, j y2. 'n K
4 5 1.0908,(u ;f Q ) 1.o'70 f ( d=;* 8),
y l'5 = 0.02/ F for increasing average temperature and 0 for decreasing average j g temperature, )
7 tgS
=
3, The function generated by the rate-lag controller for T,yg dynamic 3
compensation, (7
=
rg Time constant utilized in the rate-lag controller for T,yg, rg =,5 sec(Wds t I L),
1 7 3
As defined in Note 1, l As defined in Note 1, rk
( uni t z.), . c o l W/* F ( "*Ii ')
, K =
6 0.00126/*FA for T > T" and K6 = 0 for T 5 T",
T '= As defined in Note 1, I
T"
= 1 588.A*F Reference T,yg at RATED THERIML POWER, S = . As defined in !!ote 1, and f 2(AI) = 0 f r all.AT.
.O .( R4 cneq oge.c on PW " %.
(* D ,
TABLE 2.2-1 (Continued)
N g REACTOR TRIP SYSTEM INSTRUMENTATION TRIP SETPOINTS -
,/
E NOTATION (Continued)
Note 3: The channel's maximum Trip Setpoint shall not exceed its computed Trip Setpoint by more than 2%.
g .
a-fu g 0 l
. 2.1 SAFETY LIMITS BASES 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 regime could i
- result in excessive cladding temperatures because of the onset of departure j from nucleate boiling (DNB) and the resultant sharp reduction in heat transfer coefficient. DNB is not a directly measurable parameter during operation and therefore THERMAL POWER and reactor coolant temperature and pressure have been F related to DNB through the WRB-1 correlation. The WRB-1 DNB correlation has been developed to predict the DNB flux and the location of DNB for axially uniform and nonuniform heat flux distributions. The local DNB heat flux ratio (DNBR),
defined as the ratio of the heat flux that would cause DNB at a particular core location to the local heat flux, is-indicative of the margin to DNB.
The minimum value of the DNBR during steady-state operation, normal operational transients, and anticipated transients is limited to 1.30. This C value corresponds to a 95% probability at'a 95% confidence level that DNB will not occur and is chosen as an appropriate margin to DNB for all operating conditions.
The curves of Figures 2.1-1 and 2.1-2 show the loci of points of THERMAL POWER, Reactor Coolant System pressure and average temperature for which the minimum DNBR is no less than 1.30, or the average enthalpy at the vessel exit g is equal to the enthalpy of saturated liquid.
% of 1.55 and aThesecurvesarebasedonanenthalpyhotchannelfactor,Fh,An reference cosine with a peak of 1.55Nf r axial power shape. )
allowance is included for an increase in Fg .at reduced power based on-the i expression:
F = 1.55 [1+ 0.2 (1-P)]
Where P is the fraction of RATED THERMAL POWER.
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 fy (delta I) function of the Overtemperature trip. When the axial power imbalance is not within the tolerance,. the axial power imbalance effect on the '
Overtemperature Delta T trips will reduce the Setpoints to provide protection 1 consistent with core Safety Limits. !
)
McGUIRE - UNITS 1 and 2 B~2 ;
e eu * *
- e
~ -
,---w- -,v
~
i l
SAFETY LIMITS
~...
BASES N,, ,,
LSect# )
l 2.1.2 REACTOR COOLANT SYSTEM PRESSURE '
The restriction of this Safety Limit protects the integrity of,the Reactor Coolant System from overpressurization and thereby prevents the release of radionuclides contained in the reactor coolant from reaching the
, containment atmosphere.
The reactor vessel and pressurizer are designed to Section III of the ASME Code for Nuclear Power Plants which permits a maximum transient pressure of 110% (2735 psig) of design pressure. The Safety Limit of 2735 psig is therefore consistent with the design criteria and associated code requirements.
The entire Reactor Coolant System is hydrotested at 3107 psig,125% of design pressure, to demonstrate integrity prior to initial operation.
. .es
( 9.# c.c. 9 over o*
uw (2.g t - E-t w McGkJIRE-UNITS 1and2 B 2-2
e i
6 f.c (Al.i a 3 !
the DNB design basis is as follows: there must be' at least a 95 percent probability that the minimum DNBR of the limiting rod during Condition I and II events is greater than or equal to the DNBR limit of the DNB dorrelation being used (the WRB-1 correlation in this applica- . ..
tion). The correlation DNBR limit is established based on the entire -
applicable experimental data set such that there is a 95 percent pro-bability with 95 percent confidence that DNB will not occur when the minimum DNBR is at the DNBR limit.
In meeting this design basis, uncertainties in plant operating para-meters,' nuclear and thermal parameters, and fuel fabrication parameters bl are considered statistically such that there is at least a 95 confidence g
that the minimum DNBR for the limiting red is greater than or ecual to the DNBR limit. The uncertainties in the above plant parameters are used to determine the plant DNBR uncertainty. This DNBR ur. certainty, combined with the correlation DNBR limit, establishes a design DNBR value which must be met in plant tafety analyses using values of input parameters without uncertainties.
The curves of Figure 2.1-1 show the loci of points of- THERiiAL POWER, Reactor Coolant System pressure and average temperature below which the calculated DNBR is no less than the design DNBR value or the average enthalpy at the vessel exit is less than the enthalpy of satu-rated liquid.
I --
The curves are based on a nuclear entnalpy rise hot channel facto",
3, of snace.
power 1.49 and a reference cosine with a peak of 1.55 for axial
' An allowance is included for an increase in-F3g at
'ecutec power based on the expression:
F3g.1,49 1 + o,3 (1.p) 4 where P is the fraction of RATED THERMAL POWER.
i These -limiting heat flux conditions are higher than those calculated l for the range of all control rods fully withdrawn to the maximum allow-
\ able-control rod insertion assuming the axial power imbalance is within i the -limits of the ft (AI) function of the Overtemperature trip.
- When the axial power imbalance is not within the tolerance, the axial pcwer imoalance effect on the Overtemperature ai trips will recuce the -
set::oints to provide protection consistent with core safety limits. '
{:
7* f
l LIMITING SAFETY SYSTEM SETTINGS
)
BASES I
Power Ranoe, Neutron Flux (Continued)
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. Hich Rates na 4edi., .m h 6 ==y y g%g, The Power Range Positive Rate trip provides protection against rapid flux increases which are characteristic of _ rupturc-cf a cent.el red e .R hnneia Specifically, this trip complements the Power Range Neutron Flux High and Low trips to ensure that the criteria are met for rod ejection from 4-:J ,.nnr y%l peer, T Power R Jav3M dro ge Negative / tate t ip pryvidesprotec 'on foy contr 1 rop ci nts. At hijjh powet , a r d drop act-i nt of singye or m tip e co d ca e local flu peaki whi coul cause an unc nserva ive ocal h _ ,pRoccu ring exis . The ' ower ange gati tate rip w 11 pre ent th s fr m trip ng t react r. N credi is t en for operat on o th Po r Ra e Nega ive R e tri for t ose c trol r d dro accide ts f r w ch D R's ill be reate than .30.
Intermediate and Source Ranoe, Neutron Flux The Intermediate and Source Range, Neutron Flux trips provide core protection during reactor startup to mitigate the consequences of an-uncon-
- trolled rod cluster control assembly bank withdrawal from a subcritical -
condition. These trips provide redundant protection to the Low Setpoint trip of the Power Range, Neutron Flux channels. The Source Range channels will initiate a Reactor trip at about 10'5 counts per second unless manually
- blocked when P-6 beccmes active. The Intermediate Range channels will initiate a Reactor trip at a current level equivalent to approximately 25% of RATED THERMAL POWER unless manually blocked when P-10 becomes active.
The Power Range Negative Rate trip provides protection to ensure that the' calculated DNBR is maintained above the design DN8R value for rod drop accidents. At high power a single or multiple roc drop accident ,
could cause local flux peaking which when in conjunction with nuclear- :~
(C power being maintained equivalent to turbine power by action of the automatic rod control system could cause an unconservative local DNER to-exist. The Power Range Negative Rate trip will prevent this from
.cccurring by tHpping the reactor for all single or multiple droppea rocs.
)
2 McGUIRE -~ UNITS 1 and 2 B 2-4'
-g ,. , - -,-rr- -= er
3/4.1 REACTIVITY CONTROL SYSTEMS 3/4.1.1 BORATION CONTR_0L,
~
SHUTDOWN MARGIN - T,y >200*F LIMITING CONDITION FOR OPERATION 3.1.1.1 The SHUTDOWN MARGIN shall be greater than or equal to 1.6% delta k/k (WM 2.))
' r f = h;p ;p rati =. t.1 *f. A o s k / k (.u.ai) h 6 ,t..g.qm g ,
APPLICABILITY: MODES 1, 2*, 3, and 4.
r .
ACTION: m3 Q t.'V/. elcib k/k ( u.d Q, WiththeSHUTDOWNMARGINlessthan1.6%deltak/k[immediatelyinitiateand continue boration at greater than or equal to 30 gpm of a solution containing greater than or equal to 7000 ppm boron or equivalent until the required SHUTDOWN MARGIN is restored.
(
SURVEILLANCE REOUIREMENTS 4.1.1.1.1 The SHUTDOWN MARGIN shall be determined to be greater than or equal to 1.6% delta k/kx(u./,h) , t.1't. cidh k/k ( u.;* 0 ;
- a. Within 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br /> after detection of an inoperable control rod (s) and at least once per 12 hours1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br /> thereafter while the rod (s) is inoperable.
If the inoperable control rod is immovable or untrippable, the atave required SHUTDOWN MARGIN shall be verified acceptable with an increased allowance for the withdrawn worth of the immovable or untrippable control rod (s);
- b. When in MODE 1 or MODE 2 with Keff greater than or equal to 1.0 at least once per 12 hours1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br /> by verifying that control bank withdrawal is within the limits of Specification 3.1.3.6;
- c. When in MODE 2 with K,ff less tnan 1.0, within 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br /> prior to achieving reactor criticality by verifying that the predicted critical control rod position is within the limits of Specification 3.1.3.6;
- d. Prior to initial operation above 5% RATED THERMAL POWER after each fuel loading, by consideration of the factors of Specification 4.1.1.1.le., below, with the control banks at the maximum insertion
( limit of Specification 3.1.3.6; and "See special Test Exception 3.10.1.
. McGUIRE - UNITS 1 and 2 3/4 1-1
REACTIVITY CONTROL SYSTEMS MODERATOR TEMPERATURE COEFFICIENT LIMITING CONDITION FOR OPERATION 3.1.1.3 The moderator temperature coefficient (MTC) shall be: -
, +h 1;m;+s shown in F; 31-0, 6.-[*5jpjs,_
u4t t; {t{3e than j gijg { fg. yg ug{ ,;pre.f g 7 7 _-
(
8.
, , n , _ . j
.r- -
~p u.in,)Less p'. b positive than 0 delta k/k/*F for the all rods withdrawn, beginning of cycle life (BOL), hot zero THERMAL POWER condition- I
' and '
Ge o is s + s ,
Less negative than -4.1 x 10 -4 delta k/k/*F for the all rods
$f )f.C withdrawn, and of cycle life (EOL), RATED THERMAL POWER condition.
~ ~
- " N I APPLICABILITY: Specificatio[3.'1.'1.3a. AMI.l.'5b. - M5 1**
Specification 3.1.1.3K - MODES 1, 2, and 3 only#. .
A rsotJ: C-
- a. With the MTC more positive than the limit of Specifications 3.1.1.3a.go v-
- 11. t.~3,1p above, operation in MODES land 2 may proceed provided:
6 u.;i t
( 1. dontroi rod withdrawal limits are establi.shed and maintained sufficient to restore the MTC to less positive.than 0 dte the limity, shmn in Figure 3.\-0 4tAtFF within 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> or be in HOT STANDBY within the next 6 hours6.944444e-5 days <br />0.00167 hours <br />9.920635e-6 weeks <br />2.283e-6 months <br />. These withdrawal limits shall be in addition to the
_ insertion limits of Specification 3.1.3.6;.
- 1. -
2.& a
- Con $rol rod withdrawal limits are estabifshed and maintaine sufficient to restore the MTC to less positive than 0 delta ,3 k/k/'F within 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> or be in HOT STANDBY within the next 'f 6 hours6.944444e-5 days <br />0.00167 hours <br />9.920635e-6 weeks <br />2.283e-6 months <br />. These withdrawal limits shall be in addition to the insertion limits of Specification 3.1.3.6; 3,2' The control rods are maintained within the withdrawal l limits i
established above until a subsequent calculation verifies that the MTC has been restored to within its limit for the all rods withdrawn condition; and-
'M In lieu of any other. report required by Specification 6.9.1, a Special Report is prepared and submitted to the Commission pursuant to Specification 6.9.2 within 10 days, describing the value of the measured MTC, the interim control rod withdrawal -
limits, and the predicted average core burnup necessary for restoring the positive MTC to within its limi't for the all rods withdrawn condition.
t
- b. With the MTC more negative than the limit of Specification'3.1.1.3K '.
above, be in HOT SHUTDOWN within 12 hours1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br />.
"With K,ff greater than or equal to -1.0. ~
- See Special Test Exception 3.10.3.
QCgggq Quit 5 iA2 2N lW L _
I t
i l
REACTIVITY CONTROL SYSTEMS SURVEILIANCE REQUIREMENTS s,
4.1.1.3 The MTC shall be determined to be within its limits during each fuel cycle as follows: 3,g,g, g
- a. The MTC shall be measure and compared to the BOL limit of Specifications 3.1.1.3a.4, above, prior to initial operation above 5%
of RATED THERMAL POWER, after each fuel loading; and m b. The MTC shall be measured at any THERMAL POWER and compared to
-3.2 x 10 ~4 delta k/k/*F (all rods withdrawn, RATED THERMAL POWER condition) within 7 EFPD after reaching an equilibrium boron -
concentration of 300 ppe. In the event this comparison indicates
-4 the MTC is more negative than -3.2 x 10 delta k/k/*F, the MTC shall be remeasured, and compared to the EOL MTC limit of Specifica-tion 3.1.1.3Mf), at least once per 14 EFPD during the remainder of the fuel cycle. c
('l McGUIRE - UNITS 1 and 2 3/4 1-5 l
l 1
l l
l
( c ua 1
(
R 0.5
,4 e 0.2 -
5
'O Acceptable Unacceptacle C Operation Oceration
'u v
e 0.3 -
0 3
2 y 0.2 -
=-
B
( j 4
e' 0.1 -
5 0 10 20 30 40 50 60 70 80 90 100
" of Ratec Thermal Power LAMY l
' Figure 3.1-0. Moderator Temaeraturu Coefficient vs. Power Level Mc%Rs DMD L Me 1 s/9 l-64
REACTIVITY CONTROL SYSTEMS
{~
ROD DROP TIME LIMITING CONDITION FOR OPERATION 3.1.3.4 The individual full-length shutdown and control rod drop time from the fully withdrawn position shall be less than or equal to 2.2 ;;;;;.d;
('Jnit 1) Or 3.3 seconds J ..it Cy from beginning of decay of stationary gripper coil voltage to dashpot entry with:
3
- a. T,yg greater than or equal to 551*F, and ,
- b. All reactor coolant pumps operating.
APPLICABILITY: MODES 1 and 2.
ACTION:
- a. With the drop time of any full-length rod determined to exceed the above limit, restore the rod drop time to within the above limit prior to proceeding to MODE 1 or 2.
- b. Withtheroddropt.imeswithinkimitsbutdeterminedwiththree
( ~
reactor coolant pumps operating, operation may proceed provided THERMAL POWER is restricted to less than or equal to (*) of RATED THERMAL POWER.
SURVEILLANCE REQUIREMENTS 4.1.3.4 The rod drop time of full-length rods shall be demonstrated through measurement prior to reactor criticality:
- a. For all rods following each removal of the reactor vessel head,-
- b. For specifically affected individual rods following any maintenance on or modification to the Control Rod Drive System which could affect the drop time of those specific rods, and
- c. At least once per 18 months.
"These values left blank pending NRC approval of three loop operation.
I 1
- s.
s .
, McGUIRE - UNITS 1 and 2 3/4 1-19
l r .- l 22e . - - ,
220 3
-/f (29%, 228) /_- (79%, 228)=__y 5 t
./ .
Y ,
200 ,-
j = BANK B -
180 lW '
Y~(6%,162) _
-/ (100%,161)_.
5 . -
i 5 i -
5 140 ,
z'= BANK C 5j f- / l M 120 /
'/ l e ,
/ :
5100 i
- I -
e' l f i f f .
g 80 i -'
E ,'
/ BANK D l f i
?
c:
60
[l>
pi
~
/
k g=(0%, 47) I 40 ;
- e
' 1 20 . #
(30%, 0)=-/
0
~
}
l D 20 40 60 80 100 1
Relative Power (Percent)
FIGURE 3.1-1 ROD BANK INSERTION LIMITS 1 VERSUS RELATIVE POWER MCGUIRE UNIT 1
,/
MCGUIRE UNIT 2 '
j i
26uMGM-erwriT3-- e l
$/pc$
v v k v iP 8 -
TP)S$i)0 0 IM f 3/4.2.%, AXIAL FLUX DIFFERENCE (AFD) ( L4=M O LIMITING CONDITION FOR OPERATION 3.2.14The indicated AXIAL FLUX DIFFERENCE (AFD) shall be maintained within the allowed operational space defined by Figure 3.2-14 APPLICABILITY: MODE 1 ABOYE 50 PERCENT RATED THERMAL POWER ACTION:
- a. With the indicated AXIAL FLUX DIFFERENCE outside of the Figure 3.2-1 limits, 1.) Either restore the indicated AFD to within the Figure '
3.2-lylimits within 15 minutes, or 2.) Reduce THERMAL POWER to less than 50% of RATED THERMAL POWER within 30 minutes and reduce the Power Range Neutron Flux - High Trip setpoints to less than or equal to 55 percent of RATED THERMAL POWER within.
the next 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br />.
- b. THERMAL POWER shall not be increased above 50% of RATED THERMAL POWER unless the indicated AFD is within the Figure 3.2-lap limits. -
) SURVEILLANCE REQUIREMENTS 4.2.1.14The indicated AXIAL FLUX DIFFERENCE shall be determined to be within its limits during POWER OPERATION above 50 percent of RATED THERMAL POWER by:
- a. Monitoring the indicated AFD for ea'ch' OPERABLE excore channel:
- 1. At least once per.7 days when the AFD Monitor Alarm is OPERABLE, and
- 2. At least once per hour for the first 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> after re-storing the AFD Monitor Alarm to OPERABLE status.
1 b. Monitoring and logging the indicated AXIAL FLUX. DIFFERENCE for '
l each OPERABLE excore channel at least once per hour for the first 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> and at least once per 30 minutes thereafter, -
when the AXIAL FLUX DIFFERENCE Monitor Alarm is inoperable.
The logged values ~of the indicated AXIAL FLUX DIFFERENCE shall
, be assumed to exist during the interval _ preceding each log-ging.
4.2.1.24 The indicated AFD shall be considered outside of its limits when at least 2 OPERABLE excore channels are indicating the AFD to be-outside the limits.
. Mc.A'we u.;t.1.A.A 2.
3/4qI.M
3/4.2 POWER DISTRIBUTION LIMITS l
3/4.2.1 AXIAL FLUX DIFFERENCE f D N ( d LIMITING CONDITION FOR OPERATION l
3.2.1 The indicated AXIAL FLUX DIFFERENCE (AFD) shall be maintained within the following target band (flux difference units) about the target flux difference:
- a. t 5% for core average accumulated burnup of less than or equal to 3000 MWD /MTU, and
- b. + 3% -12% for core average accumulated burnup of greater than .
3000 MWD /MTU.
i APPLICABILITY: MODE 1 above 50% of RATED THERMAL POWER *.
ACTION:
- a. With the indicated AXIAL FLUX DIFFERENCE outside of the above required target band about the target flux difference and with THERMAL POWER:
I
- 1. Above 90% of RATED THERMACPOWER, within 15 minutes either:
( a) b)
Restore the indicated AFD to within the target band limits, or Reduce THERMAL POWER to less than 90% of RATED THERMAL POWER.
- 2. Between 50% and 90% of RATED THERMAL POWER:
a) POWER OPERATION may continue provided:
- 1) The indicated AFD has not been outside of the adove required target band for more than 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br /> penalty deviation cumulative 'during the previous 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />, and
- 2) The indicated AFD is within the limits shown on Figure 3.2-1. Otherwise, reduce THERMAL POWER to less than 50% of RATED THERMAL POWER within 30 minutes and reduce the Power Range Neutron Flux-High-Trip Setpoints
' to less than or equal to 55% of RATED THERMAL POWER within the next 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br />.
b) Surveillance testing of the Power Range Neutron Flux channels may be performed pursuant to Specification 4.3.1.1 provided the indicated AFD is maintailied within the limits of Figure 3.2-1. A total of 16 hours1.851852e-4 days <br />0.00444 hours <br />2.645503e-5 weeks <br />6.088e-6 months <br /> operation may be accumulated with the'AFD outside of the target band during this testing without penalty deviation.
- b. THERMAL POWER shall not be increased above 90% of RATED THERMAL (x POWER-unless the indicated AFD is within the above required target band and ACTION a.2.a) 1), above has been satisfied.
- See Special Test Exception 3.10.2.
2.-2.-
McGUIRE - UNITK y 3/4e64sm.
og
POWER DISTRIBUTION LIMITS ACTION (Continued) 3 e
- c. THERMAL POWER shall not be increased above 50% of RATED THERMAL POWER unless the indicated AFD has not been outside of the above "equired target band for more than 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br /> pinalty deviation cumulative during the previous 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />. Power increases above 50%
of RATED THERMAL POWER do not require being within the target band provided the accumulative penalty deviation is not violated.
SURVEILLANCE REOUIREMENTS
- 4.2.1.1 The indicated AXIAL FLUX DIFFERENCE shall be determined to be within its limits during POWER OPERATION above 15% of RATED THERMAL POWER by
- a. Monitoring tne indicated AFD for each OPERABLE excore channel:
- 1) At least once per 7 days when the AFD Monitor Alarm is OPERABLE, and
- 2) At least once per hour for the first 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> after restoring tne AFD Monitor Alarm to OPERABLE status.
- b. Monitoring and logging the ind cated AXIAL FLUX DIFFERENCE for each OPERABLE excore channel at least once per hour for the first ]
24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> and at least once per 30 minutes thereafter, when the AXIAL /
) FLUX DIFFERENCE Monitor Alarm is inoperable. The logged values of the indicated AXIAL FLUX DIFFERENCE shall be assumed to exist during 1
the interval preceding each logging.
4.2.1.2 The indicated AFD snall be considered outside of its target band when t
1 two or more OPERABLE excore channels are indicating the AFD to-be outside the target band. Penalty deviation outside of the target band shall be accumulated on a time basis of:
j a. One minute penalty deviation for each 1 minute of POWER OPERATION outside of the target band at THERMAL POWER levels equal to or above 50% cf RATED THERMAL POWER, and l
- b. One-half minute penalty deviation for each 1 minute of POWER OPERATION l outside of the target band at THERMAL POWER levels between 15% and 50% of RATED THERMAL POWER.
4.2.1.3 The target flux difference of'each OPERABLE excore channel shall be determined by measurement at least once per 92 Effective Full Power Days. The provisions of Specification 4.0.4 are not applicable.
4.2.1.4- The target fiux difference shall ::c updated at least once per 31 Effective Full Pom r Ocyr by eitnce determining the target flux difference
- pursuant to Specification 4.2.1.3 aba <e or by linear interpolation between the T. .1 most recently measurco value and 0% at the end of the cycle life. The provi- j I l sions of Specification 4.0.4 are not aaM icable.
3 l.
-McGUIRE - {ITX l3 2/4 2-f ok.
l l
I I
i
. l 5.
~ i!.
Si. c.
NI a l
'i E W 5
'~ i l
= l
(-15,100) (6,100) l UNACCEPTABLE:
OPERATION UNACCEPTABLE OPERATION 80 ACCEPTABLE OPERATION
.l - . .
60 50
(-31,50 ) (17,50) 40 _.
20 0
-50 -40 -30 -20 -10 0 10 20 30 40 50 Flux Difference (AI)%
FIGURE 3.2-1k AXIAL FLUX DIFFERENCE LIMITS AS A FUNCTION OF RATED THERMAL POWER (h'.I1{
u d ,<c - u :b i e z-d s/9 M 1
l '
( .
N.
s.
Ec
=w- ',.
! EE I
i c12 4
$$_a 4
(I<
c=2 2:e onS!
100 #15 i
UNACCEPTABLEB(-11,90) (11,90lEUNACCEPTABLE
'DPERATION _"~~ :; - OPERATION , ,
80 f -
=
i -
80 ~ hCCEPTABLEEOPERATION:
( 31,50) - - (31,50) 20 4
- O j 50 30 20 '-10 0 'O- '20- 30 40' 50
. LU?' DIFFERENCE (A!) *$
11-L %
FIGURE 3d AXIAL FLUX C!FFERENC2 I !MITS AS A FUNCTION OF RATED _ -
THERMAL POWE3 ( t.A='.t t ) -
- /
s-
, .1
- CD r 3 S
.d
' McGUIRE - UNIT (+-ee 2
. - . -~ . - , .:.__-...- -
. - - , . . .a, . . . ,
~
POWER DISTRIBUTION LIf1ITS 3/4.2.2 HEAT FLUX HOT CHANNEL FACTOR - F Z) ,
1 l
LIMITING CONDITION FOR OPERATION 3.2.2 Fq (Z) shall be limited by the following relationships:
JL_f 7T < F7.1?} TV/7 g O s n c; gj ( i;, ,,, i 7Ms Uc6 : BT m h Fq (71 < r4-Qve[z)] fc7 2Q a L ? > o.r ( u 3 $
; c,5.
Where P = THERMAL POWER RATED THERMAL POWER '
, and K(Z) is the function obtained from Figure 3.2-2 for a given core height location.
AP.; _ICABILITY: MODE 1. ACTION: With Fq (Z) exceeding its limit:
- a. %
Reduce THERMAL POWER at least 1% for each 1% F ? within15minutesandsimilarlyreducethePow0r(Z)exceedsthelimit Range Neutron 2 Flux-High Trip Setpoints within the next 4 hours; POWER OPERATION may proceed for up to a total of 72 hcurs; subsequent POWER
.4 ky) OPERATION may proceed provided the Overpower Delta T Trip Setpoints(4k have and boen reouced at least 1%,for each 1% qF (Z) exceeds the limit;
( l,, g
- b. Identify and correct the cause of the.out-of-limit condition prior to increasing THERMAL POWER above the reduced limit required by ACTION a., a'ove; c THERMAL POWER may then be increased provided F (2) ;
is demonstrated through incore mapping to be within its limit. q ) l (F 9G) f D. 23 ( KGB 4, p y o,,7 (q,,;g 4 b wg Fq (z) 5 [ 2.15 ] [K(Z)] for P > 0.5 Cu-;* ) I e I
%g ~c to.s- s.n l [ K Ot] 4 - 9 i O S C. 4 t Q !
OC? Fq (z) 1 [ 2.15 ) [K(Z)] for P 10.5 (.4'.+ 1) u 0.5 , i
~
G 1 McGUIRE - UNITS 1 and 2 3/4 2-# 1 i
I l ' i POWER DISTRIEUTION LIMITS l SURVEILLANCE REQUIREMENTS ( Nd d 4.2.2.1 The provisions of Specification 4.0.4 are not applicable. 4.2.2.2 FQ (z) shail.be evaluated to determine if FQ (z) is within its limit try:
- a. Using the mcveable incore detectors to obtain a power distri-bution map at any THERMAL POWER greater than 5 percent of RATED THERMAL POWEF..
- b. Increasing tM measured Fg(z) component of the power distri-bution map by 3 percent to account for manufacturing tolerances and further increasing the value by 5 percent to account for measurement uncertainties.
- c. Satisfying the f'ollowing relationship:
M 2.15 Fg (z) S p x g(7)x g(z) for P > 0.5 Fg"(z) S y 7, l 9, .for P S 0.5 where r (z) is the measured Fg(z) increased by the allow-antes for manufacturing tolerances and measurement uncertainty, 2.15 is the Fg limit, K(z)'.is given in Figure 3.2-2, P is the relative THERMAL POWER, and W(z) is the cycle dependent function that accounts for power distribution transients encountered during normal operation. This function is given in the Peaking Factor Limit Report as per Specification 6.g.1.12. N
.d. Measuring Fg (z) according to the following schedule:
- 1. Upon achieving equilibrium conditions after exceeding by 10 percent or more of RATED THERMAL POWER, the THERMAL POWER at which Fg(z) was last determined,* or
- 2. At least once per 31 effective full power days, whichever occurs,first.
Non 0 9 I *During power esc 41ation at the beginning of each cycle, power level may (be increased achieved and auntil pow.!radistribution power level for extended operation has been map'obtained. a - o ,t e m .b
i -ATTACHMENT-A l l FCKE?. O!STF.:EUTIC;i _ MIT5 SL*?.YEILLAri;E F.EOUIF.EMEriTS iCent) '- -
- e. With measurements indicating maximus F" (z) over z qj has increased since the pretious determination of FqN (z) either of the following actions shall be taken:
- 1. F0 M(z) shall be increased by 2 percent over that specified in 4.2.2.2.c, or
- 2. Fn M(z) shall be measured at laasi; once per 7 effec-tive full power days until 2 successive maps indicate that maximum FM (z) over z is not increasing.
ggj f. With the relationships specified in 4.2.2.2.c above not being satisfied:
- 1. . Calculate the percent Fa(z) exceeds its limit by the following expression:
7 .- . 1 M ' j F0(z)xW(2) maxmum l
/ -1 $ x 100 fcr P > 0.5 I over z ~
2J5 I p I K(z) I ,: : I ~ maximum F0 "IZ)
- WII) '
I
)
over z 1 (x100 for P < D.5 2.15 I x K(z) 0.5 _ i
- 2. Either of the following actions shall be taken:
- a. Within 15 minutes, control the AFD to within new AFD limits which are determined by reducing the AFD limits of 3.2-1 by 1% AFD for each percent Fg(z) exceeds its limits as determined in 4.2.2.2.f.1. Within 8 hours, reset or the AFD alarm setpoints to these modified limits,
- b. Comply with the requirements of Specification 3.2.2 for Fg(z) exceeding its limit by the percent calculated above.
?2. a
~,
@wE.R DtSTibMnoQ us rT5 suevetuAuc.t REomEE.ueurs ( uJ t t I (Co the..L) I
- g. The limits specified in 4.2.2.2.c, 4.2.2.2.e, and 4.2.2.2.f above are not applicable in the following core plane regions:
- 1. Lower core region 0 to 15 percent inclusive.
- 2. Upper core region 85 to 100 percent inclusive.
4.2.2.3 When gF (z) is measured for reasons other than meeting the requirements of Specification 4.2.2.2 an overall measured Fg(z) shall
- be obtained from a power distribution map and increased by a percent to account.-for manufacturing tolerances and further increased by 5 percent to account for measurement uncertainty.
1 i l l h W Cd.cc - Q4 ts ( a .2. 2/q ~4r)k _m_
POWER DISTRIBUTION LIMITS SURVEILLANCE REOUIREMENTS ( (A d i d 4.2.2.1 The provisions of Specification 4.0.4 are not applicable. 4.2.2.2 F xy shall be evaluated to determine ifqF (Z) is within its limit by:
- a. Using the movable incore detectors to obtain a power distribution map at any THERMAL POWER greater than 5% of RATED THERMAL POWER,
- b. Increasing the measured F xy component of the power distribution map by 3% to account for manufacturing tolerances and further increasing the value by 5% to account for measurement uncertainties,
- c. Co:nparing the F computed (Fx) btained in Specification 4.2.2.2b.,
above, to:
- 1) The F xy limits for RATED THERMAL POWER (F x ) f r the appropriate measured core planes given in Specification 4.2.2.2e. and f.,
below, and ___
- 2) The relationship:
F =F P [l+0.2(1-P)], . Where F ' is the limit for fractional THERMAL POWER operation express as a function of FxRTP and P is the fraction of RATED THERMAL POWER at which F xy was measured.
- d. Remeasuring F according to the following schedule:
- 1) When FC ,is greater than the FxRTP 1 it r the appropriate measured core plane but less than the F relationship, additional
*Y power distribution maps shall be taken and F C compared to F RTP xy xy l and F,, either:
a) Within 24 hours after exceeding by 20% of RATED THERMAL POWER cr greater, the THERMAL POWER at which F C was xy last d:termined, or b) At least once per 31 EFPD, unichever occurs first. ok
, .74 McGUIRE - UNITK W 2 3/4 2-(
POWER DISTRTBUTION LfMITS r SURVEILLANCE REOUIREMENTS (Continued) (LA [,i ~2 )
- 2) When the F is less than or equal to the F RTP limit for the appropriate measured core plane, additional power distribution R l maps shall be taken and F compared to F and F xy at least once per 31 EFPD.
- e. The F RTP xy limits for RATED THERMAL POWER (Fxy ) shall be provided for all core planes containing Bank "0" control rods and all unrodde'd core planes in a Radial Peaking Factor Limit Report per Specifica-tion 6.9.1.12,
- f. The F xy limits of Specification 4.2.2.2e., above, are not applicable in the following core planes regions as measured in percent of core height from the bottom of the fuel: '
- 1) Lower core region from 0 to 15%, inclusive,
,... 2) Upper core region from 85 to 100%, inclusive,
- 3) Grid plane regions at 17.8 2%, 32.1 2%, 46.4 2%, 60.6 2 2%
and 74.9 2%, inclusive, and
- 4) Core plane regions within 2% of core height ( 2.88 inches) about the bank demand position of the Bank "0" control rods.
L
- g. With F exceeding F , the effects of F on Fq (Z) shall be ,
evaluated to determine ifqF (Z) is within its limits. 4.2.2.3 When qF (Z) is measured for other than F determinations, an'averall measured qF (Z) shall be obtained from a power distribution map and increased by 3% to account for manufacturing tolerances and further increased by 5% to account for measurement uncertainty. i l 7
. og McGUIRE - UNITi-1-ame 2 3/4 2-7d
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.~ ~
M McGUIRE - UNITS 1 and 2 3/42-X l
,_c ..
,. . . q-POWER OfSTRIBUTION LIMITS i
3/4.2.3 RCS FLOW RATE AND NUCLEAR ENTHALPY RISE HOT CHANNEL FACTOR i LIMITING CONDITION FOR OPERATION 3.2.3 The combination of indicated Reactor Coolant System (RCS) total flow rate and R , R. shall be maintained within the region of allowable operation 3 shown on Figure, 3.2-3 for four loop operation: Where: w (uf. (1) F" E 64
** , g( .
1A' l . 49 [1. 0 +-+d-( 1. 0. - P)] ,4 4_ I.4(Lo + o.1 O - e.3 b* c p" HFyWpje
- R @aNd 1 : El ) L
~
R M N '( t-REP (.M)
" p _
~ THERMAL POWER ,
RATED THERMAL PGWER -
- d. Fh=MeasuredvaluesofFhobtainedbyusingthemovableincore detectors to obtain a power distribution map. The measured aluesofFhshallbeusedtocalculateRsinceFigure3.2-3 includes penalties for undetected feedwater venturi fouling of
- 0. 3 and for measurement uncertainties of 1.7% for flow and 4%
for incore measurement of Fh, and e. RBP (BU) = Rod Bow Penalty as a function of region average burnup as shown in Figure 3.2-4, where a region'is defined as those assemblies with the same loading date (reloads) or enrichment (first core). ( Apphh % WW 1 - i l y). APPLICABILITY: MODE 1. ACTION: With the combination of RCS total flow rate and R), R 2 utside the region of i acceptable operation shown on Figure 3.2-3:
- a. Within 2 hours- either:
i 1. Restore R the combination of RCS total flow rate and R), ! 2 t within the above limits, or 2. Reduce THERMAL POWER to less than 50% of RATED THERMAL POWER) ! and reduce the-Power Range Neutron Flux - High Trip Setpoint to / less than or equal to 55% of RATED THERMAL POWER within the next 4 hours. j 1 McGUIRE.- UNITS I and 2 '3/4 2-8 Amendment No. 22(Unit 1) Amendment No. 3 (Unit 2) 6/2S/83
POWER DISTRIBUTION LIMITS ACTION: (Continued)
- b. Within 24 hours of initially being outside the above limits, verify through incore flux mapping and RCS total flow rate comparison that the combination of R , R and RCS total flow rate are restored to within the above lim ts,2or reduce THERMAL POWER to less than 5% of RATED THERMAL POWER within the next 2 hours. l I
- c. Identify and correct the cause of the out-of-limit condition prior to increasing THERMAL POWER above the reduced THERMAL POWER limit required by ACTION a.2. and/or b. above; subsequent POWER OPERATION may proceed provided that.the combination of R3 , R,, and indicated RCS total flow rate are demonstrated, through incoPe flux mapping and RCS total flow rate comparison, to be within the region of acceptable operation shown on Figure 3.2-3 prior to exceeding the following THERMAL POWER levels:
- 1. A nominal 50% of RATED THERMAL POWER,
- 2. A nominal 75% of RATED THERMAL POWER, and
- 3. Within 24 hours of attaining greater than or equal to 95% of RATED THERMAL POWER.
(}SURVEILLANCEREOUIREMENTS 4.2.3.1 The provisions of Specification 4.0.4 are not applicable. 4.2.3.2 The combination of indicated RCS total flow rate determined by process computer readings or digital voltmeter measurement and R and R shall be within the region of acceptable operation of Figure 3.2 3: 2
- a. Prior to operation above 75% of RATED THERMAL POWER after each fuel loading, and
- b. At least once per 31 Effective Full Power Days.
4.2.3.3 The indicated RCS total flow rate shall be verified to be within the ' region of acceptable operation of Figure 3.2-3 at least once per 12 hours when the most recently obtained values of Ry and R , btained per Specification 4.2.3.2, are assumed to exist. 2 4.2.3.4 The RCS total flow rate indicators .shall be subjected to a CHANNEL CALIBRATION at least once per 18 months. 4.2.3.5 The RCS total flow rate shall be determined by precision heat balance measurement at least once per 18 months. 3 Amendment No.22 (Unit 1)
- McGUIRE - UNITS 1 and 2 3/4 2-R Amendment No. 3- (Unit ~2) l 6/28/83 t
" 48
- i I
- f. j , . ! . ..[% _.
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t l l ! l i. r PENALTIES OF 0.1% FOR UNDETECTED
-~ - ~
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E ACCEPTABLE ~~ lI 1 FEF9 WATER VENTURI FOULING AND 5
' l 46 MEASUllEMENT UNCERTAINTIES OF "
; OPERATION /- {,
M 1.7% FOR FLOW AND 4% FOilINCOilE ~'-8
~-'~" ! "~ / ' R EGION FQR'-t----
. N --~ ~ I c MEASullEMENT OF F " AllE INCLUDED / R'2ON)<Y !
e 14 Tills FIGURE. #
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~2 FIGI)RE 3.2-3 RCS 10lAL F10WHAIE VERSUS R AND R2 - F0llR LOOPS IN OPERATION 3 .a. '
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PENALTIES OF 0.1% FOR UNDETFCTED i a: * . l 46 FEEDWATER VENTURI FOULING AND ! i ! y
- g i HEASURFHENT UNCERTAINTIES OF ,
g I '!.7% FOR FLOW AND 4% FOR INCORE g I 8 m j HEASUREMENT OF Fy ARE INCLUDED
, j ;
c ! . IN'TilIS FICURE 8 i ^ i : 1 ACCEPTABLE I id t 44 l l
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OPERATION i .
, a i l
8 l ' REGION FOR - E i- l- l. . l !. ! l ! Ry ONLY ' ! l !
.y . . . : . . .
l I i g , ACCEPTABLE , 1 i
- 4 OPERATION I ' l o l 42 ,
REGION FOR ' (1.031. 42.0)
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- , ! OPERATION -
. l l. REcION 4
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, 3g-(1.0, 38.I10)
J 6 94.--RTP (1.0, 37.721)' 5 92TRTP ~
'! s 90% RTP ! !
(1.0, 36.944) , . I i .
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36 I I ! * '
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0.90 1.00 1.02 ' 1.04 1.06 [p
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I ! l ! i j A .g. Ry i= F[g,1.49[1.0 + 0.2(1.0 - P)] .
. . i .
l ! ,. R;=R/[1-.RDP(Dil)] 7 y l 3 '
-FIGURE 3.2-3b UNIT 2 RCS Flot! RATE VERSUS R; Atl0 R2 - FOUR LOOPS IN OPERATIO!! -
. =_ - _ . . . - .
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. McGUIRE - UNITS 1'a-d 2. 3/f 2"o l "yc' --y-.- - -
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FIGURE 3.2-49 UNIT 2 R09 110W PENALTY AS A FilflCTION Of BilRilllP ,
l 3/4.5 EMERGENCY CORE COOLING SYSTEMS
, [k ' j 3/4.5.1 ACCUMULATORS i
~
COLD LEG INJECTION LIMITING CONDITION FOR OPERATION l 3.5.1.1 Each cold leg injection accumulator shall be OPERABLE with:
- a. The isolation valve open,
- b. A contained borated water volume of between:025! ;;d M00 pilun.,
c. Oc4 . A baron concentration of.between 1900 and 2100 ppa,
; d. A nitrogen cover pressure of between 400 and 454 psig, and
- e. A water level and pressure channel OPERABLE.
g potz. .a 22.5L g.il s (
- M N APPLICABILITY: MODES 1, 2, and 38
~
d g g ty (u.,4 z)
') ACTION:
- a. With one cold leg injection accumulator inoperable, except as a result of a closed isolation valve, restore the inoperable accumulator to OPERABLE status within 1 hour or be in at least HOT STAN08Y within the next 6 hours and in HOT SHUTDOWN within the following .
6 hours. [i
- b. With one cold leg injection accumulater inoperable due to the !
; isolation valve being closed, either immediately open the isolation valve or be in at least HOT STAND 8Y within 1 hour and in HOT SHUTDOWN within the following 12 hours.
SURVEILLANCE REQUIREMENTS i l
)
4.5.1.1.1 Each cold leg injection accumulator shall be demonstrated OPERABLE:
- a. At least once per 12 hours by:
- 1) Verifying the contained borated water volume and nitrogen cover-pressure in the tanks, and f- 2) Verifying that each cold leg injection accumulator isolation
( valve is open.
" Pressurizer pressure above 1000 psig.
l McGUIRE - UNITS 1 and 2 3/4 5-1 l
EMEMENCYCORECOOLINGSYSTEMS j 3/4.5 x BORON INJECTION SYSTEM
'{
BORON INJ TION TANK 'N s., LIMITING CONDIDON FOR OPERATION ' s - l 3.5.4 The boron inje tion tank shall be OPERABLE with:
- a. A minimum cont ned borated water volume 900 gallons, and
- b. Between 2000 and 00 ppm of boron.
APPLICABILITY: MODES 1, 2 and (Unit 1 y) ACTION: With the boron injection tank inoper e, restore the tank to OPERABLE status within 1 hour or be in HOT STANDBY nd rated to a SHUTDOWN MARGIN equivalent to 1% delta k/k at 200*F within e next hours; restore the tank to OPERABLE status within the next 7 days be in HOT UTDOWN within the next 12 hours. J C SURVEI' LANCE REQ REMENTS
- 1 x 4.5.4 Th oron injection tank shall be demonstrated OPERAB by: .
. Verifying the contained borated water volume at least ce per ; 7 days, and
- b. Verifying the boron concentration of the water in the' tank least once per 7 days.
( Dcle.;tocA. l
. McGUIRE - UNITS.1 and 2 3/4 5-11 l
1 l - ,, , . , . ., -
C 3/4.1 REACTIVITY CONTROL SYSTEMS BASES 3/4.1.1 BORATION CONTROL 3/4.1.1.1 and 3/4.1.1.2 SHUTDOWN MARGIN A sufficient SHUTDOWN MARGIN ensures that: (1) the reactor can be made subcritical from all operating conditions, (2) the reactivity transients associated with postulated accident conditions are controllable within
.1cceptable limits, and (3) the reactor will be maintained sufficiently subcritical to preclude inadvertent criticality in the shutdown condition.
SHUTDOWN MARGIN requirements vary throughout core life as a function of . fuel depletion, RCS boron concentration, and RCS T The most restrictive condition occurs at EOL, with T atnoloadoper$lin.g temperature, and is associatedwithapostulatedstllElinebreakaccidentandresultinguncon-trolled RCS cooldown. In the analysis of this ace' dent, a minimum SHUTDOWN MARGIN of i.C J.M.. k/$s required to control tra reactivity transient. Accordingly, the SHUTDOWN MARGIN requirement is based upon this limiting condition and is consistent with FSAR safety. analysis assumptions. With T lessthan200*F,the.reactivitytransientsresultingfromapostulated.stelM9 ( line break cooldown are minimal and a 1% delta k/k SHUTDOWN MARGIN provides adequate protection. i.(. % M+. k/h 6 3 % l.1 Y' . 3/4.1.1.3 MODERATOR TEMPERATURE COEFFICIENT Mk k / k CuJ+ ) The limitations on moderator temperature coefficient (MTC) are provided to ensure that the value of this coefficient remains within the limiting condition assumed in the FSAR accident and transient analyses. The MTC values of this specification are applicable to a specific set of plant conditions; accordingly, verification of MTC values at conditions other than those explicitly stated will require extrapolation to those conditions in order to permit an accurate comparison. _ The most negative MTC value equivalent to the most positive moderator density coefficient (MDC), was obtained by incrementally correcting the MDC used in the FSAR analyses to nominal operating conditions. These corrections involved subtracting the incremental change in the EC associated with a core condition of all reds inserted (most positive MDC) to an all rods withdrawn condition and, a conversion for the rate of change of moderator density with temperature at RATED THERMAL POWER conditions. Th transformedintothglimitingMTCvalue-4.1x10{svalueoftheMDCwasthen delta k/k/*F. The MTC value of -3.2 x 10 delta k/k/*F represents a conservative value (with corrections for-burnup and soluble boron) at a core condition of 300 ppm equilibrium boron concentration and,{s obtained by making these corrections to the limiting MTC value of -4.1 x ,10 k/k/*F. N . McGUIRE - UNITS 1 and 2 B 3/4 1-1
l l l C' 3/4.2 POWER DISTRIBUTION LIMITS l BASES c lc.ulsted " l The specifications of this sectio provide assurance of fuel integrity during Condition I (Normal Operation) nd II (Incidents of Moderate Frequency) o -{ , ,. events by: (1) maintaining the 'H- DNBR in the core ;;--m- thx x pal g '^ ' 30 during normal operation and in short-term transients, and (2) limiting the fission gas release, fuel pellet temperature, and cladding mechanical M.y U properties to within assumed design criteria. In addition, limiting the peak linear power density during Condition I events provides assurance that the initial conditions assumed for the LOCA analyses are met and the ECCS acceptance criteria limit of 2200'F is not exceeded. The definitions of certain hot channel and peaking factors as used in these specifications are as follows: F0 (Z) Heat Flux Hot Channel Factor, is defined as the maximum local heat flux cn the surface of a fuel rod at core elevation Z divided by the average fuel rod heat flux, allowing for manufacturing tolerances on fuel pellets and rods; N F g Nuclear Enthalpy Rise Hot Channel Factor, is defined as the ratio of the integral of linear power along the rod with the highest integrated C . power to the average rod power; and u,lA F (Z) Radial Peaking Factor, is defined as the ratio of peak power density
- 2. to average power density in the horizontal plane at core elevation Z.
3/4.2.1 AXIAL FLUX DIFFERENCE g g g 9 ,; 4 q 3 The limits on AXIA LUX DIFFERENCE (AFD) assure that the F n (Z) upper cound envelope of 2.32 ime he normalized axial peaking factor is not exceeded during either normal operation or in the event of xenon redistribution following power changes.
~
Target flux difference is determined at equilibrium xenon conditions. ' The full-length rods may be positioned within the core in accordance with l their respective insertion limits and should be inserted near their normal
- u. 't, position for steady-state operation at high power levels. The value of the
, target flux difference obtained under these conditions divided by the fraction of RATED THERMAL POWER is the target flux difference at RATED THERMAL POWER for the associated core burnup conditions. Target flux differences for other THERMAL POWER levels are obtained by multiplying the RATED THERMAL POWER value by the appropriate fractional THERMAL POWER level. The periodic updating of the target flux difference value is necessary to reflect core burnup
_ considerations. ( ( McGUIRE - UNITS 1 and 2 B 3/4 2-1 a.-y & = = --
l POWER DISTRIBUTI0t' LIMIT 2 Bt.SES AXIAL FLUY DI:FERECE (Continu20 . Althougn it is intencec tnct the plant will be operated with the AFD within the target band reouired by Specification 3.2.1 about the target flux difference, during rapid plant THERMAL POWER reductions, control rod motion will cause the AFD to deviate outside of the target band at reduced THERMAL POWER levels. This oeviation 5:111 not affect the xenon redistribution sufficiently to change the envelope of peaking factors which may be reached on a subsequent return to RATED THERf1AL POWER (with the AFD within the target band) provided the time curatior, of the deviation is limited. Accordingly, a 1 hour penalty deviation limit cumulative during the previous 24 hours is provided for operation outside o' the target band but within the limits of g Figure 3.2-1 unile at THERMAL PGi!ER ievels between 50% and 90% of RATED
- 2. THERMAL POWER. For THERP.AL POVEF. levels between 15% and 50% of RATED THERMAL POWEP,, devictions of tne AFD cutsico of the target band are less significant.
The penalty of 2 neur tctual time reflect: this reduced significance. Provisions for monitoring the AFD on an automatic basis are derived from the plant process computer tnrough the AFD honitor Alarm. The computer determines tr E 1 minute average of each c4 the OPERABLE excore detector outputs and provices en clarm message immediately if the AFD for two or more OPERABLE excore cnar.ncis cre outside the tarcet band and the THERMAL POWER is greater than 90% of F.;TED THEPN.t. POWER. Duiing operation at THERMAL POWER levels between 50% cnc 90% an:: between 15% ano 50% RATED THERMAL POWER, the ] j computer outputs cn clarn nossage wnen the penalty deviation accumulates beyond the limits of i hour and 2 hour:, respectively. Figure E VI 2-1 rnos: a typical monthly target band. Inwt 3/t.2.E and 2R..? MIAT FLUX HOT CHANKEL FACTOP., and RCS FLOWRATE AND Nt' CLEAR EN1HI.'_F Rin KC CHAT 4%. Fi>C10F; The limits en ne:t flu: hot enannel factor, RCS flowrate, and nuclear entncipy rise no: encnnri factar ensure that: (1) the design limits on peak iccal power cenrin cad reinimut DN3R are not exceeded, and (2) in the event of e LOCA the pa: fus. cicd temcerature will not exceed the 2200*F ECCS EcCEptcnct Critteli Iimit. Each of thest ; mc:surcole but will norm &lly only be determined pericciet?iy c: It:ccified ir Soccifications 4.2.2 and 4.2.3. This periodic surveillancc is surficier.; to insure tnat the limits are maintained provided:
- n. Cer.trM rce ; e single grcup move togsther with no. individual rod inser'c.ior. cifi r' m: by ::rc then _--it stepsg M::t:d from the 3
preup centr > nc '- en: G
- c. Cct.tr: "c. .m : rt i,u ut.r. cec W tn overiapaing groups as described e
n S'.:c:i.N ctt.t: . . i . 5. - - , W.G.'; ._ ' % i cad 2 t 4/2 2 .:
1 I D F 0Y4 !4 ca: fer .ueaitering the ATO vu ou cuivmati;$ ac c LLw't i.,ived
.. .. '"s f,un, c
~ the pl:nt pr;;;;; ::mputer throu;h th; ATO nvoivor Aiers. Jbkhe computer determines the one minute average of each of the OPERABLE excore detec-CA*i,t ter outputs and provides an alarm message immediately if the AFD for at i least 2 of 4 or 2 of 3 OPERASLE excore channels are outside the allowed al-Power operating space and the THERMAL POWER is greater than 50 percent of RATED THERMAL POWER.
ss l
1 ! POWER DISTRIBUTION LIMITS l f BASES . l HEAT FLUX HOT CHANNEL FACTOR. and RCS FLOW RATE AND NUCLEAR ENTHALPY RISE HOT CHANNEL FACTOR (Continueo) c. The control rod insertion limits of Soecifications 3.1.3.5 and 3.1.3.6 are maintaineo; and
- d. The axial power distribution, expressed in terms of AXIAL FLUX DIFFERENCE, is maintained within the limits.
N Fg will be maintained within its limits proviced Conditions a. throuan ~ d..above are maintained. As noted on Figures 3.2-3 and 3.2-4, RCS flow rate and [g may be " traded off" against one another (i.e. , a low measured RCS flow rate is acceptable if the measured F N is also low) to ensure that the calcu-lated DNBR will not be below the design DNBR value. The relaxation of Fh as a function of THERMAL POWER allows changes in the racial power shace for all permissible rod insertion limits. R. as calculated in Specification 3.2.3 and used in Figure 3.2-3, accounts for FN~ g less than or equal to 1.49. This value is used in the various ac:ident analyses wnere F N influences parameters other than DNBR, e.g., peak clad tam-perature, and thus is the maximum "as measured" value allowed. R , as defined, 2 allows for the inclusion of a penalty for Rod Bow on DNBR only. Thus, knowing the "as measur?d" values of Fh and RCS flow allows for " tradeoffs" in excess of R ecual to 1.0 for the purpose of offsetting the Rod Bow DNBR penalty.
~
~ ' ' ' ~
INSERT i di When an F q measurement is taken, an allowance for both experimental error and manufacturing tolerance must be made. An allowance of 5% is amoropriate ' for a full-core mao taken witn tne Incore Detector Flux Maccing System, and a 3% allowance is approcriate for manufacturing tolerance. l l l l Amenc=ent No . 22 Jni t l',, McGUIRE - UNITS 1 and 2 8 3M 7-4
. M_ _* " . ~ . " "No, , _ 1. ~ " . " " _ ' ' . _ _ . . _
~- .
_EC 17. l I nse.4 X Fuel rod bowing reduces the value of DNB ratio. Credit is available to partially offset this reduction. This credit comes from generic or plant
^
specific design margin. For McGuire Unit ( l nr.d 2, the margin used to partially offset rod bow penalties is 9.1 percent. This margin breaks down as follows:
- 1) Design limit DNBR 1.6%
- 2) Grid spacing K s 2.9%
- 3) . Thermal Diffusion Coefficient 1.2%
3 4) DNBR Multiplier 1.7%~
- 5) Pitch Reduction 1.7%
\%w _ =-
rnr Mennira Unit--2, the margin used to partially offset rod bow penalties is 5.9 percent with the remaining 3.2 percentLused to trade off against measured flow being as much as 2 percent lower than thermal design flow plus uncertainties. The penalties applied to FN to account for rod bow AH (Figure %3.2-4Mnit 1 and-Unit-2-)~ as a function of burnup are consistent with those described in Mr. John F. Stolz's (NRC) letter to T. M. Anderson (Westinghouse) dated April 5,1979 with the difference being due to the amount of margin each unit uses to partially offset rod bow penalties. m 1
C POWER DISTRIBUTION LIMITS i BASES l HEAT FLUX HOT CHANNEL FACTOR and RCS FLOW RATE AND NUCLEAR ENTHALPY RISE 4, HOT CHANNEL FACTOR (Continued) When RCS flow rate and F are measured, no additional allowances are
- necessary prior to comparison with the limits of Figures 3.2-3 and 3.2-4.
Measurement errors of 1.7% for RCS total ficw rate and 4% for FN have been allowed for in determination of the design ONBR value. i The measurement error for RCS total flow rate is based upon performing
..a precision heat balance and using the result to calibrate the RCS flow rate
- indicators. Potential fouling of the feedwater venturi which might not be j
detected could bias the result from the precision heat balance in a non-conservative manner. Therefore, a penalty of 0.1% for undetected fouling of the feedwater venturi is included in Figure 3.2-3. Any fouling which t might bias the RCS flow rate measurement greater than 0.1% can be detected ' by monitoring and trending various plant performance parameters. If detected, action shall be taken before performing subsequent precision heat balance {
- measurements, i.e., either the effect of the fouling shall be quantified
- and compensated for in the RCS flow rat ~e measurement or the venturi shall be cleaned to eliminate the fouling.
l ( The 12-hour periodic surveillance of indicated RCS flow is sufficient to detect only flow degradation which could lead to operation outside the accept- '
% er9 ,able region of operation shown on Figure 3.2-3. @ 3/4.2.4 OUADRANT POWER TILT RATIO The QUADRANT POWER TILT RATIO limit assures that the radial power distribution i
satisfies the design values used in the power capability analysis. Radial power i distribution power operation. measurements are made during STARTUP testing and periodically during I e
'mit of 1.02, at which corrective action is required #b unB 4; 'Q [gand linear hea limiting tilt of 1.025 can ion rate protection with v-/ pieHr power tilts. A e ore the margin for uncertainty in Fq is depleted. A .02 was se ec '
an allowance for the unce associated with the indicated power tilt. The 2-hour time allowanca for operation with a tilt condition greater i than 1.02 but less than 1.09 is provided to allow identification and correc- , l tion of a dropped or misaligned cent =1 rod. In the event such action does not i correct the tilt, the margin for uncertainty on Fqis reinstated by reducing the maxi r SivJ power by 3% for each percent of tilt in excess of 1.0. t L -fe. m EAW D r n caAt. ?oesR 1 1 Amendment No. 22 (Unit 1) j McGUIRE - UNITS 1 and 2 B 3/4 2-5 Amendment No. 3 (Unit 2) 6/28/83
d T*5 , s ~ The hot channel factor F0 M(z) is measured periodically and in-creased by. a cycle and height dependent power factor, W(z), to provide g ,.}. assurance that the limit on the hot channel factor, F (z), is met. Q W(z) accounts for the effects of normal operation transients and was 1 determined from expected power control maneuvers over the full range of burnup conditions in the core. The W(z) function for normal operation is provided in the Pe'a king Factor Limit Report per Specification
. 6.9.1 12.
z-- POWER DISTRIBUTION LIMITS BASES S
}
OUADRANT POWER TILT RATIO (Continued) For purposes of monitoring QUADRANT POWER TILT RATIO when one excore detector is inoperable, the moveable incore detectors are used to confirm that the normalized symmetric power distribution is consistent with the QUADRANT POWER TILT RATIO. The incore detector monitoring is done with a full incore flux map or two sets of four symmetric thimbles. The two sets of four symmetric thimbles is a unique set of eight detector locations. These locations are C-8, E-5, E-n, H-3, H-13, L-5, L-H, N-8. 3/4.2.5 DNB PARAMETERS The limits on the DNB-related parameters assure that each of the parameters are maintained within the normal steady-state envelope of operation assumed in the transient and accident analyses. The limits are consistent with the initial FSAR assumptions 'and have been analytically demonstrated
' adeque .e to maintain a ririca;;; 5"," vi 1."^ i.hroughout each analyzed transient.
cA o y 17.-t+ 09tg, - The 12-hour periodic surveillance of these parameters through instrument readout is sufficient to ensure that the parameters are restored within their limits following load changes and other expected transient operation. I l .
.)
McGUIRE - UNITS 1 and 2 Amendment No.22 (Unit 1) B 3/4 2-5 Amendment No. 3 .(Unit 2) / ; l 6/28/83 1
/ l l
3/4.4 REACTOR COOLANT SYSTEM BASES 3/4.4.1 REACTOR COOLANT LOOPS AND COOLANT CIRCULATION .
. es g The plant is designed to operate wi all raacha h d loops in r coolant operation and maintain DNBR above +.-99 uring all nor:nal operations and anticipated transients. In MODES 1.and 2 with one reactor coolant loop not in operation this specification requires that the plant be in at least HOT STANDBY within 1 hour. .
In MODE 3, a single reactor coolant loop provides sufficient heat removal capability for removing decay heat; however, single failure considerations require that two loops be OPERABLE. In MODE 4, and in MODE 5 with reactor coolant loops filled, a single reactor coolant loop or RHR Icop provides sufficient heat removal capability for removing decay heat; but single failure considerations require that at least two loops (either RHR or RCS) be 0PERABLE. In MODE 5 with reactor coolant loopInot 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 operation of one reactor coolant pump (RCP) or one RHR pump provides adequate flow to ensure mixing, prevent stratification and produce gradual reactivity changes during' boron concentratior, reductions in the Reactor , Coolant System. The reactivity change rate associated with boron reduction will, therefore, be within the capability of operator recognition and control. The restrictions on starting a reactor coolant pump with one'or more RCS cold legs less than or equal to 300*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 Part 50. The RCS will be protected against overpressure transients and will not exceed the limits of Appendix G by either: (1) restricting the water volume in the pressurizer and thereby providing a volume for the reactor coolant to expand into, or (2) 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 cold leg temperatures. McGUIRE - UNITS 1 and 2- B 3/4 4-1l
l EMERGENCY CORE COOLING SYSTEMS I BASES D # ECCS SUBSYSTEMS (Continued) . The limitation for a maximum of one centrifugal charging pump ard one Safety Injection pump to be GPERABLE and the Surveillance Requirement to verify all charging pumps and Safety Injection pumps except the required GPERABLE charging pump to be inoperable below 300*F provides assurance ts.at a mass PORV. addition pressure transier,- can be relieved by the operation of a single The Surveillance Requirements provicied to ensure OPERABILITY of each component ensures that at a minimum, the assumotions used in the safety , analyses are met and that subsystem OPERABILITY is maintained. Surveillance Requirements for throttle valve position stops and flow bal ace testing provide LOCA. assurance that proper ECCS flows will be maintaine,i in the event of a Maintenance of proper flow resistance and pressure drop in the piping system to each injection point is necessary to: (1) prevent total pump flow frgm exceeding runout conditions when the system is in its minimum resistance ccnfiguration, (2) provide the proper flow split between injection points in accordance with the assumptions used in the ECCS-LOCA analyses, and (3) provide an acceptable level of total ECCS flow to-all injection points equal to or above that assumed in the ECCS-LOCA analyses. 3/4.5.4 kDilQNMKTIASAT/M e OPERABILITY of the Boron Injection System for Unit 1 as part ECCS ensu the t positive reactivitv canaa+ ha iajected in ore. RCS cooldown can be cau accident, or a steam loss-of-coolant gggg The limits on . tration er at volume and boron concen-met,
...._,.....mu in we sr.eam eak analysis are e contained water volume limit includes an allowance fo not usable because of tank discharge line location or other physical charac "stics.
3/4.5.5 REFUELING WATER STOP1,GE TANK The OPERABILITY of the refueling water storage tank (RWST) as part of the ECCS ensures that a sufficient supply of borated water is available for injection by the ECCS in the event of a LOCA. The limits on RWST minimum volume and baron concentration ensure that: (1) sufficient water is available within containment to penait recirculation cooling flow to the core, and (2) the reactor will remain subcritical in the cold condition following mixing of tne RWST and the RCS water volumes with all control rods inserted except for the most reactive control assembly. These assumptions are consistent with the LOCA analyses. i J McGUIRE - UNITS 1 and 2 B 3/4 5-2 l [
l
)
ADMINISTRATIVE CONTROLS THIRTY DAY WRITTEN REPORTS (Continued)
- c. Observed inadequacies in the implementation of Administrative or Procedural Controls which threaten to cause reduction of degree of redundancy provided in Reactor Trip Systems or Engineered Safety Features Systems;
- d. Abnormal degradation of systems other than those specified in Specification 6.9.1.10c. above designed to contain radioactive material resulting from the fission process;
- e. An unplanned offsite release of: (1) more than 1 Curie of radio, active material in liquid effluents, (2) more than 150 Curies of noble gas in gaseous effluents, or (3) more than 0.05 Curie of radiciodine in
- gaseous effluents. The report of an unplanned offsite release of radioactive material shall include the following information:
- 1) A description of the event and equipment involved,
- 2) Cause(s) for the unplanned release,
- 3) Actions taken to. prevent recurrence, and
( 4) Consequences of the unplanned release.
\ .
- f. Measured levels of radioactivity in an environmental sampling medium determined to exceed the reporting level values of Table 3.12-2 when averaged over any calendar quarter sampling period.
RADIAL PEAKING FACTOR LIMIT REPORT 6.9.1.lf The W(2) function for normal operation shall be provided to *d " the Director, Nuclear Reactor Regulations, Attention Chief of the Core the Performance Branch, U. S. Nuclear Regulatory Commission, Washington, nce D.C. 20555 at least 60 days prior to cycle initial criticality. In the ofall event that these values would be submitted at some other time during ges at core life, it will be submitted 60 days prior to the date the values !he liinit would become effective unless otherwise exempted by the Commission. .bmitted Any information needed to support W(z) will be by request from the NRC and need not be included in this report.
. an n ot'beNncludad-i SPECIAL REPORTS 6.9.2 Special reports shall be submitted to the Regional Administrator of the
( NRC Regional Office within the time period specified for each report. N. , l l
. i McGUIRE - UNITS 1 and 2 6-23
i i i i + ATTACHMENT 2 JUSTIFICATION AND SAFETY ANALYSIS ~' i F 5 1 , Mr. H. B. ' Tucker's (DPC) November 14, 1983 letter to Mr. H. R. Denton (NRC/0NRR) l described planned . changes in the fuel design for McGuire Nuclear Station, Units 1 l j and 2. McGuire Units 1 and 2 have been operating with Westinghouse 17x17 low-parasitic (STD) fueled cores. It is planned to refuel both Units 1 and 2 with Westinghouse 17x17 Reconstitutable Optimized Fuel Assembly '(OFA) regions. As a re- + sult , future core loadings would range from an approximately J /3 0FA - 2/3 STD transition core to eventually an all 0FA fueled core. The OFA fuel has similar j design features compared to the STD fuel which has had subst antial operating ' ex-perience in a number of nuclear plants. The major different es' are the use-of .six
^
I intermediate (mixing vane) Zircaloy grids for the OFA fue1 versus six intermediate (mixing vane) Inconel grids for STD fuel and a reduction in fuel rod diameter. j Major advantages for utilizing the OFA are: (1) increased efficiency of the core by reducing the amount of parasitic material and (2) reduced fuel cycle costs due to an optimization of the water to uranium ratio. l The above letter provided a Reference Safety Evaluation Report summarizing the evaluation / analysis performed on the region-by-region reload transition from the 1 present McGuire Units 1 and 2 STD fueled cores to cores with all optimized fuel. ] The report examined the differences tatween the Westinghouse OFA and STD designs and evaluated the effects of these differences for the transition to an all 0FA core. The' evaluation considered the standard reload design methods described in WCAP-9272 and 9273, " Westinghouse Reload Safety Evaluation Methodology", and the transition effects described for mixed cores in Chapter 18 of WCAP-9500-A, L "Re- l ference Core Report - 17x17 Optimized Fuel Assembly."' Consistent with the(Westing - house STD reload methodology for analyzing cycle specific reloads, parameters were chosen to maximize the applicability of the transition evaluations- for each reload - cycle and to facilitate' subsequent determination of.the applicability of 10 CFR . 50.59. Subsequent cycle specific reload safety evaluations will verify that'appli-- F cable safety limits are satisfied based on' the reference. evaluation / analyses esta-blished in the reference report. . A summary of the mechanical, nuclear, thermal and hydraulic, and: accident evaluations for the.McGuire Units.1 and 2 transitions ~to.
- an all 0FA core are given in the reference report.
WCAP-8183, " Operational Experience with Westinghouse Cores". presents the operating - experience through ' December 1981 of six 17x17 0FA demonstration assemblies which '
- - have the McGuire 1 and 2 design features.. .By early 1983, two
- 17x17 0FAs will-have-
' completed three cycles of irradiation to about.28,500 MWD /MEU,~~two have completed two cycles to about 19,400 MWD /MTU, and tw have. completed: oneicycle of irradiation -
with burnups in excess of 9,000 MWD /NTU. .All demonstration 17x17.OFAs examined' were in good' condition.- This'provides evidence of favorable operation of:ZircaloyL grids and reduced fuel rod diameters;which are the' major new desi;;n features .of the 17x17 0FA. f' Thel transition Design and. Safety Evaluations considered the following nominal operating conditions: 3411 MWt core power, 2250 psia system E pressure, 559.20F core J f f I
- . _ _ , _ . .c_ ..,,;... '
...J,, _. . m --
l inlet temperature (HFP) at 2250 psia, and 386,000 gpm RCS thermal design flow. The results of evaluation / analysis and tests described in the Reference Safety Evaluation Report lead to the following conclusions:
- a. The Westinghouse OFA reload fuel assemblies for McGuire 1 and 2 are mechanical-ly compatible with the current STD design, control rods, and reactor internals interfaces. Both fuel assemblies satisfy the current design bases for the Mc-Guire units.
- b. Changes in the nuclear characteristics due to the transition from STD to 0FA fuel will be within the range normally seen from cycle to cycle due to fuel management effects.
- c. The reload 0FAs are hydraulically compatible with the current STD design.
- d. The accident analyses for the OFA transition core were shown to provide accep-table results by meeting the applicable criteria, such as, minimum DNBR, peak pressure, and peak clad temperature, as required. The previously reviewed and licensed safety limits are met. Analyses in support of this safety evaluation establish a reference design on which subsequent reload safety evaluations in-volving 0FA reloads can be based. (Attachment 2A presents those detailed non-LOCA and LOCA accident analyses of the McGuire Units 1 and 2 FSAR impacted by the proposed changes as determined in Section 6.0 of the Reference Safety Eval-uation Report. The information contained within was prepared using the NRC Standard Format and Content Guide, Regulatory Guide 1.70, Revision 3 as it ap-plies to McGuire Nuclear Station Units 1 and 2).
- e. Plant operating limitations given in the Technical Specifications affected by use of the OFA design and positive MIC will be satisfied with the proposed changes noted in Section 7.0 of the report.
Attachment 2B is the cycle-specific Reload Safety Evaluation (RSE) for McGuire Unit 1/ Cycle 2 including Fq surveillance and BAOC Technical Specifications. The RSE presents an evaluation for McGuire Unit 1, Cycle 2, which demonstrates that the core reload will not adversely affect the saf ety of the plant. This evaluation was performed utilizing the methodology described in WCAP-9273, " Westinghouse Reload Safety Evaluation Methodology". McGuire Unit 1 is operating in Cycle lA (Safety Evaluation for the W. B. McGuire Unit 1 Core Redesign, April 1983) with all Westinghouse 17x17 low parasitic (STD) _ fuel assemblies. For Cycle 2 and subsequent cycles, it is planned to refuel the McGuire Unit 1 core with Westinghouse -17x17 optimized fuel assembly (OFA) regions. In the OFA transition licensing submittal to the NRC (Reference Safety Evaluation, November 14, 1983 letter) an analyses of the safety aspects of the transition from STD fuel design to OFA design was provided. This licensing submittal justified the compatibility of the OFA design with the STD design in a_ transition core as well. as a full 0FA core. The OFA transition licensing submittal contained mechancial, nuclear, thermal-hydraulic, and accident evaluations which are applicable to the Cycle 2 safety evaluation. All of the accidents comprising the licensing bases which could potentially be i affected by the fuel reload have _been reviewed for the Cycle 2 design described
)
i i
i I i herein. The results of new analyses are included in the above mentioned licensing i submittal and in the cycle-specific Reload Safety Evaluation, and the justifica-tion for the applicability of previous results for the remaining analyses is pre-sented. The McGuire Unit 1, Cycle 2 reactor core will be comprised of 193 fuel assemblies arranged in the core loading pattern configuration shown in Figure 1 of the Cycle 2 Reload Safety Evaluation. During the Cycle 1A/2 refueling, 60 STD fuel assemblies will be replaced with 60 Region 4 optimized fuel assemblies. A summary of the Cycle 2 fuel inventory is given in Table 1 of the Cycle 2 Reload Saf ety Evaluation. From the evaluation presented in the Cycle 2 Reload Safety Evaluation, it is con-cluded that the Cycle 2 design does not cause the previously acceptable safety limits to be exceeded. This conclusion is based on the following:
- 1. Cycle 1A burnup is between 14767 and 15533 MWD /MTU.
- 2. Cycle 2 burnup is limited to 10200 MWD /MTU.
- 3. There is adherence to plant operating limitations given in the Technical Speci-fications with the exception of the proposed changes subnitted in support of the OFA transition licensing submittal and the Fq surveillance and RAOC Techni-cal Specification changes. With these revisions, there is adherence to all plant operating limitations in the Technical Specificaticns.
To ensure plant operation consistent with the design and safety evaluation conclu-sion statements made in the Cycle 2 RSE and to ensure that these conclusions re-main valid, several Technical Specifications changes will be needed for Cycle 2. These changes are those outlined in Section 7.0 of the OFA transition ' licensing submittal and the Fq surveillance and RAOC changes given in Appendix A of the cycle-specific RSE. Attachment 1 provides copies of these specifications as they presently appear in the McGuire Units 1 and 2 Technical Specifications with the appropriate changes noted. Attachment 2C is the peaking factor limit report for McGuire Unit 1/ Cycle 2 which is provided in accordance with the proposed McGuire Unit 1 Technical Specification paragraph 6.9.1.12 as givea in Attachment 1. This report provides the W(z) functions that are to be used for RAOC operation during Cycle 2. The appropriate W(z) func-tion is used to confirm that the heat flux hot channel factor, Fq(z), will be limit-ed to the values specified in the Technical Specifications. In addition to the Technical Specification changes outlined above, two non-reload related changes are included in Attachment 1. The first involves deletion of tech spec 3/4.5.4 on the boron injection system for Unit 1. A previously approved Techni-cal Specification change (ref. Mr. W. O. Parker,' Jr. 's March 2,1982 letter to Mr. H. R. Denton) reduced boron concentration in the Boron Injection Tank (BIT) from a nominal 20,000 ppm to 2000 ppm, and deleted heat tracing for the tank. This change essentially reduced the BIT to a large section of piping for all practical purposes. Westinghouse analysis in support of that change concluded that reduction of the BIT's boron concentration does not have any adverse effect on safety of
- plant operation or the health and safety of the public. Furthermore, re-analysis.
- of the steam line break accident for end-of-lif e conditions without considering l the functioning of the BIT showed that the accident is safety mitigated. This conclusion allows for complete elimination of the BIT (as evidenced by the subsequent
licensing of McGuire Unit 2 without a BlT and corresponding Technical Specifica-tion). At the time, other plant considerations (non-safety related) led to a de-cision to reduce boron concentration rather than eliminate the BlT completely. .However, current plans are te install bypass piping around the tank during the upcoming refueling outage, removing the tank from the system. This modification will eliminate any need for the BIT Technical Specification. The second non-reless Technical Specification change included in Attachment 1 in-volves revised control rod insertion limits for both McGuire Units 1 and 2 (T.S. 3/4.1.3.6). During development of the Improved Load Follow Package (5 rod D-bank, reduced temperature return-to-power) the control rod overlap was set at 50% (114 steps). It was recently noted that the overlap in the plant cannot be set at exactly 114 steps due to hardware limitations. The overlap has therefore been changed to 113 steps. In addition, when Westinghouse Nuclear Engineering (NE) per-forms calculations with control rods present, the control rod insertion is modeled, in many cases, as a fraction inserted. The original NE generated rod insertion limits were in terms of fraction inserted. When these limits were converted to steps withdrawn, an inappropriate correlation between fraction inserted and steps withdrawn was used. The revised rod insertion limits use the correct correlation, resulting in appronmately 3 more steps of allowed D bank insertion. These revis-ed limits were used throughout the McGuire safety analysis and therefore can be implemented into the Technical Specifications. I l l l l l t
4 ATTACHMENT 2A l FSAR CHAPTER 15 ACCIDENT ANALYSIS SENSITIVITY TO PROPOSED CHANGES Accident OFA Positive MTC Reduced SDM Feecwater Malfunction X Excessive Load Increase X Steamline Depressurization X X , Steamline Rupture X X Loss of Load X X Station Blackout X Loss of Normal Feedwater X Feedline Rupture X Partial Loss of Flow X X Comolete Loss of Flow X X Locked Rotor - X RCCA Withdrawal from Suberitical X X RCCA Withdrawal at Power X Startup of an Inactive Loop X RCCA Ejection X X Incdvertent ECCS Operation X RCS Depressurization X X x 9
C C NON-LOCA ANALYSIS C L 4446Q:1:/6 A-1 i l l
TABLE OF CONTENTS Section Description Page 15.0 ACCIDENT ANALYSIS . 15.0-1 15.0.1 Classification of Plant Conditions 15.0-1 15.0.2- Optimization of Control Systems 15.0-6 15.0.3 Plant Characteristics and Initial Conditions Assumed 15.0-7 15.0.4 15.0.5 in the Accident Analyses Reactivity Coefficients Assumed in the Accident Analyses 15.0-10 ( Rod Cluster Control Assembly Insertion Characteristics 15.0-10 15.0.6 Trip Points and Time Delays to Trip Assumed in the 15.0-12 Accident Analyse: 15.0.7 Computer Codes Utilized 15.0-13 15.0.8 References 15.0-16 15.1 INCREASEINHEATREMOVALBYTdESECONDARYSYSTEM 15.1-1 15.1.1 Feedwater System Malfunction Causing an Increase in 15.1-1 Feedwater Flow 15.1.2 Excessive Increase in Secondary Steam Flow 15.1-6 15.1.3 Inadvertent Opening of a Steam Generator Relief or 15.1-9 Safety Valve 15.1.4 Steam System Piping Failure 15.1-14' 15.1 ; References 15.1-22 15.2 DECREASE IN HEAT REMOVAL BY THE-SECONDARY SYSTEM 15.2-1 15.2.1 Steam Pressure Regulator Malfunction or Failure that ( 15.2-1 Results in Decreasing Steam Flow 15.2.2 Loss of External Load 15.2-1 15.2.3 Turbine Trip- 15.2-5 15.2.4 Inadvertent Closure of Main Steam Isolation Valves 15.2-10 15.2.5 Loss of Condenser Vacuum and Other Events Causing a 15.2-10 Turbine Trip l 15.2.6 Loss of Non-Emergency AC Power to the Plant Auxiliaries 15.2-11 '~ 15.2.7 Loss of Normal Feedwater Flow t 15.2-15 i 4446Q:1/6 A-il , 1 l
C TABLE OF CONTENTS (cont.) Section Description Page 15.2.8 Feedwater System Pipe Break 15.2-19 15.2.9 References 15.2-27 15.3 DECREASE IN REACTOR COOLANT SYSTEM FLOW RATE 15.3-1 15.3.1 Partial Loss of Forced Reactor Coolant Flow 15.3-1 15.3.2 Complete Loss of Forced Reactor Coolant Flow 15.3-4
* ~
15.3.3 Reactor Coolant Pump Shaft Seizure (Locked Rotor) 15.3-6 15.3.4 References 15.3-10 15.4 REACTIVITY AND POWER DISTRIBUTION ANOMALIES 15.4-1 15.4.1 Uncontrolled Rod Cluster Control Assembly Bank Withdrawal 15.4-1
, from a Suberitical or Low Power Startup Condition 15.4.2 Uncontrelled Rod Cluster Co'ntrol Assembly Bank 15.4-7
( 15.4.3 15.4.4 Withdrawal at Power Rod Cluster Control Assembly Misoperation Startup of an Inactive Reactor Coolant Pump at an 15.4-13 15.4-22 Incorrect Temperature 15.4.5 Spectrum of Rod Cluster Control Assembly Ejection 15.4-25 Accidents 15.4.6 References 15.4-39 15.5 INCREASE IN REACTOR COOLANT INVENTORY - 15.5-1 15.5.1 Inadvertent Operation of Emergency Core Cooling System 15.5-1 During Power Operation 15.5.2 References 15.5-5 15.6 DECREASE IN REACTOR COOLANT INVENTORY 15.6 15.6.1 Inadvertent Opening of a Pressurizer Safety or 15.6-1 Relief Valve 15.6.2 References 15.6-3 - 4446Q:1/6 A-iii
_ _ _ _ ~ LIST OF TABLES Table Descriotion Pace 15.0.3-1 Nuclear Steam Supply System Power Ratings 15.0-17 15.0.3-2 Summary of Initial Conditions and Computer Codes Used 15.0-18 15.0.3-3 Nominal Values of Pertinent Plant P.arameters Utilized 15.0-21 in the Accident Analyses 15.0.3-4 Nominal Values of Pertinent Plant Parameters Utilized in the Accident Analyses 15.0-22 ( 15.0.6-1 Trip Points and Time Delays in Accident Analysis 15.0-23 15.1-1 Ties Sequence of Events for Incidents Which Cause an 15.1-23 Increase in Heat Removal by the Secondary System 15.2-1 Time Sequence of Events for Incidents Which Cause a 15.2-27 Decrease in Heat Removal by the Secondary System 15.3-1 Time Sequence of Events for Incidents Which Result in 15.3-12 a Decrease in Reactor Coolant System Flow 15.3-2 Summary of Results for Locked Rotor Transients 15.3-14 15.4-1 Time Sequence of Events for Incidents Which Cause 15.4-40 Reactivity and Power Distribution Anomalies 15.4-2 Parameters Used in the Analysis of the Rod Cluster 15.4-44 Assembly Ejection Accident 15.5-1 Time Sequence of Events for Incident Which Results in 15.5-6 an Increase in Reactor Coolant Inventory 15.6-1 Time Sequence of Events for Incident Which Causes a 15.6-4 Decrease in Reactor Coolant Inventory i A-iv 44460:1/6 '
A LIST OF FIGURES Ficure Description Pace 15.0.3-1 Illustration of Overtemperature and Overpower AT 15.0-25
- Protection j 15.0.4-1 Doppler Power Coefficient Used in Accident Analysis 15.0-26
~
15.0.5-1 RCCA Position Versus Time to Dashpot 15.0-27 15.0.5-2 Normalized Rod Worth Versus Percent Inserted 15.0-28 l 15.0.5-3 Normalized RCCA Bank Reactivity Worth Versus Normalized 15.0 29 l Drop Time 15.1.1-1 Feedwater Control Valve Malfunction 15.1-26 I 15.1.1-2 Feedwater Control Valve Malfunction 15.1-27
- 15.1.2-1 Ten Percent Step Load Increase, Minimum Moderator 15.1-28
- Feedback, Manual Reactor Control l 15.1.2-2 Ten Percent Step Load Incre,ase, Minimum Reactivity 15.1-29 l Feedback, Manual Reactor Control 1
15.1.2-3 Ten Percent Step Load Increase, Maximum Reactivity 15.1-30 l Feedback, Manual Reactor Control ! 15.1.2-4 Ten Percent Step Load Increase, Maximum Reactivity 15.1-31 Feedback, Manual Reactor Control I 15.1.2-5 Ten percent Step Luad Increase, Minimum Reactivity 15.1-32 j Feedback, Automatic Reactor Control j 15.1.2-6 Ten Percent Step Load Increase, Minimum Reactivity 15.1-33 i Feedback, Automatic Reactor Control l 15.1.2-7 Ten Percent Step Load Increase, Maximum Reactivity 15.1-34 Feedback, Automatic Reactor Control 15.1.2-8 Ten Percent Step Load Increase, Maximum Reactivity 15.1-35 . 4 Feedback, Automatic Reactor Control- ) 15.1.3-1 K,ff Versus Temperature 15.1-36 15.1.3-2 Failure of a Steam Generator Safety or Dump Valve 15.1-37 15.1.3-3 Failure of a Steam Generator Safety or Dump Valve 15.1-38 4446Q: 1/6 g,y
LIST OF FIGURES (cont.) k. Figure Description Pace 15.1.4-1 Doppler Power. Feedback 15.1-39 ( 2 15.1.4-2 1.4 Ft Steamline Rupture, Offsite Power Available 15.1-40 15.1.4-3 1.4 Ft 2Steamline Rupture, Offsite Power Available 15.1-41 15.1.4-4 1.4 Ft 2Steamline Rupture, Offsite Power Available 15.1-42 2 15.1.4-5 1.4 Ft Steamline Rupture, Offsite Power Not Available ' 15.1.4-6 2 1.4 Ft Steamline Rupture, Offsite Power Not Available 15.1 15.1-44 ( 2
~15.1.4-7 1.4 Ft Steamline Rupture, Offsite Power Not Available 15.1-45 15.2.3-1 Turbine Trip' Event With Pressurizer Spray and Power 15.2-37 Operated Relief Valves, Minimum Reactivity Feedback 15.2.3-2 Turbine Trip Event With Pressurizer Spray and Power 15.2-38 Operated Relief Valves, Minimum Reactivity Feedback 15.2.3-3 Turbine Tr'i p Event With Pressurizer Spray and Power 15.2-39 Operated Rel'ief Valves, with Maximum Reactivity Feedback 15.2.3-4 Turbine Trip Event With Pressurizer Spray and Power 15.2-40 Operated,Reljef Valves, with Maximum Reactivity Feedback 15.2.3-5 Turbine Trip Event Without Pressurizer Spray or Power 15.2-41 Op'erated " Relief Valves, with Minimum Reactivity Feedback 15.2.3-6 Turbine Trip Event Without Pressurizer Spray or Power 15.2-42 Operated Rel.ief Valves, with Minimum Reactivity Feedback 15.2.3-7' Turbine Trip' Event Without Pressurizer Spray and Power 15.2-43~
Operated Relief Valves, with Maximum Reactivity Feedback 15.2.3-8 Turbine Trip Event Without Pressurizer Spray and Power 15.2-44 Operated Relief. Valves, with Maxim 4 :itetivity Feedback 15.2.6-1 Pressurizer Pressure and Water 'W., F isients'for-Loss 15.2-45
- of Offsite Power- '
15.2.6-2 Core Average Temperature Transient'and Steam Generator = 15.2-46 Press,ure for Loss of Offsite Power 44460:1/6 -. A-vi.
LIST OF FIGURES (cont.) Fioure Description Pace 15.2.7-1 Pressurizer Pressure and Water Volume Transients for Loss 15.2-47 of Normal Feedwater 15.2.7-2 Loop Temperatures and Steam Generator Pressure for Loss 15.2-48 of Normal Feedwater ( 15.2.8-1 Nuclear Power Transient, Total Core Reactivity Transient and Feedline Break Flow Transient of Main Feedline Rupture with Offsite Power Available 15.2-49 15.2.8-2 Pressurizer Pressure, Water Volume, and Relief Transient 15.2-50 for Main Feedline Rupture with Offsite Power Available 15.2.8-3 Reactor Coolant Temperature Transients for the Faulted 15.2-51 and the Intact Loops for Main Feedline Rupture with Offsite Power Available
~
15.2.8-4 Steam Generator Pressure and Core Heat Flux Transients 15.2-52 for Main Feedline Rupture with Offsite Power Available 15.2.8-5 Nuclear Power Transient, Total C' ore Reactivity, and 15.2-53 Feedline Break Flow Transient for Main Feedline Rupture Without Offsite Power Available 15.2.8-6 Pressurizer Pressure, Water Volume, and Relief Rate for 15.2-54 Main Feedline Rupture'Without Offsite Power Available 15.2.8-7 Reactor Coolant Temperature Transients for the Faulted 15.2-55 and Intact Loops for Main Feedline Rupture Without Offsite Power Available ; 15.2.8-8 Steam Generator Pressure and Core Heat Flux Transients 15.2-56 l for Main Feedline Rupture Without Offsite Power Available 15.3.1-1 Flow Transients for Four Loops in Operation, One Pump 15.3-15 Coasting Down 15.3.1-2 Nuclear Power and Pressurizer Pressure Transients for 15.3-16 Four Loops in Operation, One Pump Coasting Down L 4446Q:1/6 A-vii
l l LIST OF FIGURES (cont.) Figure Descriotion Page 15.3.1-3 Average and Hot Channel Heat Flux Transients for Four 15.3-17 C. 4 Loops in Operation, One Pump Coasting Down 15.3.1-4 DNBR Versus Time for Four Loops in Operation, One Pump 15.3-18 Coasting Down 15.3.2-1 Core Flow Coastdown for Four Loops in Opera [ ion, Four Pumps Coasting Down 15.3-19 ( , 15.3.2-2 Nuclear Power and Pressurizer Pressure Transients for 15.3-20 Four Loops in Operation, Four Pumps Coasting Down 15.3.2-3 Average and Hot Channel Heat Flux Transients for Four 15.3-21 Loops in Operation, Four Pumps Coasting Down 15.3.2-4 DNBR Versus Time for Four Loops in Operation, Four Pumps -15.3-22 Ccasting Down , 15.3.3-1 Flow Transients for Four. Loops in Operation, One Locked 15.3-23 Rotor 15.3.3-2 Peak Reactor Coolant Pressure for Four Loops in Opera- 15.3-24 tion,'One Locked Rotor 15.3.3-3 Nuclear Power Average and Hot Channel Heat Flux Tran-15.3-25 sient for-Four Loops in Operation, One Locked Rotor 15.3.3-4 Maximum Clad Temperature at Hot Spot for Four Loops in '15.3-26 Operation, One Locked Rotor 15.4.1-1 Neutron Flux Transient for Uncontrolled Rod Withdrawal 15.4-45 from a Suberitical Condition 15.4.1-2 Thermal Flux Transient for Uncontrolled Rod Withdrawal- 15.4-46 from a Suberitical Condition . 15.4.1-3 Fuel and Clad Temperature Transients for_ Uncontrolled 15.4-47 4 Rod Withdrawal from a Subcritical Condition
. 15.4.2-1 Uncontrolled RCCA Bank Withdrawal from Full-Power with. -15.4-48 .
Minimum Reactivity Feedback (75 pcm/secLWithdrawal -
. Rate)' .
III 4446Q:1/6 e 1
- . - - - -. ,_c- , , - . ,
A LIST OF FIGURES (cont.) Fiqure Description Page 15.4.2-2 Uncontrolled RCCA Bank Withdrawal from Full Power with 15.4-49 Minimum Reactivity Feedback (75 pcm/sec Withdrawal Rate) 15.4.2-3 Uncontrolled RCCA Bank Withdrawal from Full Power with 15.4-50 ( 15.4.2-4 Minimum Reactivity Feedback (75 pcm/sec Withdrawal Rate) Uncontrolled RCCA Bank Withdrawal from Full Power with 15.4-51 Minimum Reactivity Feedback (1 pcm/see Withdrawal Rate) 15.4.2-5 Uncontrolled RCCA Bank Withdrawal from Full Power with 15.4-52 Minimum Reactivity Feedback (1 pcm/see Withdrawal Rate) 15.4.2-6 Uncontrolled RCCA Bank Witidrawal from Full Power with 15.4-53 Minimum Reactivity Feedback (1 pcm/sec Withdrawal Rate) 15.4.2-7 Minimum DNBR Versus Reactivity Insertion Rate Rod With- 15.4-54 drawal from 100% Power 15.4.2-8 Minimum DNBR Versus Reactivity Insertion Rate Rod With- 15.4-55 drawal from 60% Power IS.4.2-9 Minimum DNBR Versus Reactivity Insertion Rate Rod With - 15.4-56 drawal from 10% Power 15.4.3-1 Nuclear Power Transient and Core Heat Flux Transient 15.4-57 for Dropped Rod Cluster Control Assembly 15.4.3-2 Pressurized Pressure Transient and Core Average Tempera-- 15.4-58 ture Transient for Dropped Rod Cluster Control Assembly 15.4.4-1 Improper Startup of an Inactive Reactor Coolant Pump 15.4-59 15.4.4-2 Improper Startup of an Inactive Reactor Coolant Pump 15.4-60 - 15.4.4-3 Improper Startup of an Inactive Reactor Coolant Pump 15.4-61 L 44460:1/6 A-ix
l l I i ! LIST OF FIGURES (cont.) i-Fioure Description Page 15.4.4-4 Improper Startup of an Inactive Reactor Coulant Pump 15.4-62 }- 15.4.4-5 Improper Startup of an Inactive Reactor Coolant Pump 15.4-63 j- 15.4.5-1 Nuclear Power Transient BOL HFP Rod Ejection Accident 15.4-64 , j 15.4.5-2 Hot Spot Fuel and Clad Temperature Versus Time BOL HFP 15.4-65 Rod Ejection Accident 15.4.5-3 Nuclear Power Transient EOL HZP Rod Ejection Accident- 15.4-66 15.4.5-4 Hot Spot Fuel and Clad Temperature Versus Time EOL HZP 15.4-67 Rod Ejection Accident 15.5-1 Inadvertent Operation of ECCS-During Power Operation 15.5-7. 15.5-2 Inadvertent Operation of ECCS During Power Operation 15.5-8 15.5-3 Inadvertent Operation of ECCS During Power Operation 15.5-9 15.6-1 Inadvertent Opening of a Pressurizer Safety Valve 15.6-5 15.6-2 Inadvertent Opening of a Pressurizer Safety Valve 15.6-6 i
- l
( 4446Q:1/6
..A-x' t
v e- er y- --- - f
15.0 ACCIDENT ANALYSIS 15.0.1 CLASSIFICATION OF PLANT CONDITIONS Since 1970, the American Nuclear Society (ANS) classification of plant conditions has been used which divides plant conditions into four categories in accordance with anticipated frequency of occurrence and potential radiological consequences to the public. The four categories ( are as follows:
- 1. Condition I: Normal Operation and Operational Transients.
- 2. Condition II: Faults of Moderate Frequency.
- 3. Condition III: Infrequent Faults.
- 4. Condition IV: Limiting Faults.
~
The basic principle applied in relating design requiremsnts to each of the conditions is that the most probable occurrences should yield the least radiological risk to the public and those extreme situations having the potential for the greatest risk.to the public'shall be those least likely to occur. Where applicable, reactor trip system and engineered safeguards functioning is assumed to the extent allowed by considerations such as the single failure criterion in fulfilling this principle. 15.0.1.1 CONDITION I - NORMAL OPERATION AND OPERATIONAL TRANSIENTS Condition I occurrences are those which are expected frequently or regularly in the course of power operation, refueling, maintenance, or maneuvering of the plant. As such, Condition I occurrences are accommodated with margin between any plant parameter and the value of that parameter which would require either automatic or manual protective. action. Inasmuch as Condition I occurrences occur frequently or C 0955T:1/6 15.0 i regularly, they must be considered from the point of view of affecting the consequences of fault conditions (Conditions II, III and IV). In this regard, analysis of each fault condition described is generally based on a conservative set of initial conditions ccrresponding to adverse conditions which can occur during Condition I operation. A typical list of Condition I events is listed below:
- 1. Steady state and shutdown operations
- a. Power operation
- b. Startup (K,ff)
- c. Hot standby
- d. Hot shutdown
- e. Cold shutdown
- f. Refueling
- 2. Operation with permissible deviations Various deviations which may occur during continued operation as permitted by the plant Technical Specifications must be considered in conjunction with other operational modes. These include:
- a. Operation with components or systems out of service (such as power operation with a reactor coolant pump out of service)
L T 0955T:1/6 15.0-2
1 l 1
- b. Radioactivity in the reactor coolant, due to leakage from fuel with cladding defects.
( 1) Fission products
- 2) Corrosion products
- 3) Tritium
- c. Operation with steam generator leaks up to the maximum allowed by.the Technical Specifications
- d. Testing as allowed by Technical Specifications
- 3. Operational transients
- a. Plant heatup and cooldown
- b. Step load changes
- c. Ramp load changes
- d. Load rejection up to and including design full load rejection:
transient 15.0.1.2 CONDITION II - FAULTS OF MODERATE FREQUENCY These faults
- at worst, result in a reactor trip with the plant being capable of returning to operation. By definition, these faults (or events) do not propagate to cause a more serious fault, i.e., Condition III or IV events. In addition, Condition II events are not expected to
. result in fuel rod failures or Reactor Conlant System or secondary
) system overpressurization. 0955T:1/6 15.0-3
For the purposes of this report, the following faults are included in this category: I
- 1. Feedwater system malfunctions that result in an increase in feedwater temperature.
2, Excessive increase in secondary steam flow.
- 3. Inadvertent opening of a steam generator relief or safety valve.
- 4. Loss of external electrical load.
- 5. Turbine trip.
- 6. Inadvertent closure of main steam isolation valves.
- 7. Loss of condenser vacuum and other events resulting in turbine trip.
- 8. Loss of nonemergency AC power to the station auxiliaries.
- 9. Loss of normal feedwater flow.
- 10. Partial loss of forced reactor coolant flow.
- 11. Uncontrolled rod cluster control assembly bank withdrawal from a suberitical or low power startup condition.
- 12. Uncontrolled rod cluster control assembly bank withdrawal at power.
- 13. Rod cluster control assembly misoperation (dropped full length assembly, dropped full length assembly bank, or statically misaligned full length ~ assembly).
0955T:1/6 15.0-4 4
. . , , ' *^ ' '
- 14. Startup of an inactive reactor coolant pump at an incorrect temperature.
( 15. Inadvertent operation of the Emergency Core Cooling System during power operation.
- 16. Inadvertent opening of a pressurizer safety or relief valve.
15.0.1.3 CONDITION III - INFREQUENT FAULTS By definition Condition III occurrences are faults which may occur very infrequently during the life of the plant. They will be accommodated with the failure of only a small fraction of the fuel rods although sufficient fuel damage might occur to preclude resumption of the operation for a considerable outage time. The release of radioactivity will not be sufficient to interrupt or, restrict public use of those areas beyond the exclusion radius. A Condition III fault will not, by ( itself, generate a Condition IV fault or result in a consequential loss of function of the Reactor Coolant System or Containment barriers. For the purposes of this report the following faults are included in this category:
- 1. Complete loss of forced reactor coolant flow.
- 2. Rod cluster control assembly misoperation (single rod cluster control assembly withdrawal at full power).
15.0.1.4 CONDITION IV - LIMITING FAULTS Condition IV occurrences are faults which are not expected to take place, but are postulated because their consequences would include the potential for the release of significant' amounts of radioactive material. They are the most drastic which must be designed against and reoresent' limiting design cases. Condition IV faults are not to cause a L 0955T:1/6 15.0-5 W
fission product release to the environment resulting in an undue risk to public health and safety in excess of guideline values of 10CFR100. A single Condition IV fault is not to cause a consequential loss of required functions of systems needed to cope with the fault including those of the Emergency Core Cooling System and the Containment. For the purposes of this report, the following faults have been classified in this category:
- 1. Steam system pipe break.
- 2. Feedwater system pipe break.
- 3. Reactor coolant pump shaft seizure (locked rotor).
- 4. Spectrum of rod cluster control assembly ejection accidents.
'5.0.2 OPTIMIZATION OF CONTROL SYSTEMS A control system setpoint study is performed in order to simulate performance of the reactor control and protection systems. In this study, emphasis is placed on the development of a control system which will automatically maintain prescribed conditions in the plant even under a conservative set of reactivity parameters with respect to both system stability and transient performance.
For each mode of plant operation a group of optimum controller setpoints is determined. In areas where the resultant setpoints are different, compromises based on the optimum overall performance are made and verified. A consistent set of control system parameters is derived i satisfying plant operational requirements throughout the core life and 1 for various levels of power operation. L 1 0955T:1/6 15.0-6 :
l l l i 15.0.3 PLANT CHARACTERISTICS AND INITIAL CONDITIONS ASSUMED IN THE ' ACCIDENT ANALYSES ( 15.0.3.1 DESIGN PLANT CONDITIONS Table 15.0.3-1 lists the principal power rating values which are assumed in analyses performed in this report. Two ratings are given:
- 1. The guaranteed NSSS thermal power output. This power output includes the thermal power generated by the reactor coolant pumps.
- 2. The engineered safety features design rating. The NSSS supplied engineered safety features are designed for thermal power higher than the guaranteed value in order not to preclude realization of future potential power capability. This higher thermal power value is designated as the engineered safety features design rating. This power output includes the thermal power generated by the reactor
( coolant pumps. Where the initial power operating conditions are assumed in accident l analyses, the guaranteed NSSS thermal power output is assumed. Where de:nonstration of adequacy of the containment and engineered safety features are concerned, the engineered safety features design rating is assumed. Allowances for errors in the determination of the steady-state power level are made as described in Section 15.0.3.2. The thermal power values used for each transient analyzed are given in Table 15.0.3-2. Ir, all cases where the 3579 megawatt thermal (MWt) rating is used in an analysis the resulting transients and consequences are conservative compared to using the 3425 MWt rating. The values of other. pertinent plant parameters utilized in the accident analyses are given in Tables 15.0.3-3 and 15.0.3-4. L 0955T:1/6 15.0-7
15.0.3.2 INITIAL CONDITIONS For most accidents which are DNB limited nominal values of initial conditions are assumed. The allowances on power, temperature, and pressure are determined on a statistical basis and are included in the limit DNBR, as described in WCAP-8567 (Reference 4). This procedure is known as the " Improved Thermal Design Procedure," and is discussed more fully in Section 4 of the RTSR. For accidents which are not DNB limited, or for which the Improved Thermal Design Procedure is not employed the initial conditions are I obtained by adding the maximum steady state errors to rated values. The following conservative steady state errors were assumed in the analysis:
- 1. Core Power 12 percent allowance for j c.alorimetric error
- 2. Average Reactor Coolant 14*F allowance for controller System temperature dead-band and measurement error plus an additional 1.5'F for steam generator fouling
- 3. Pressurizer pressure 130 pounds per square inch (psi) allowance for steady state fluctuations and measurement error Table 15.0.3-2 summarizes initial conditions and computer codes used in the accident. analysis, and shows which accidents employed a DNB analysis
.h using the Improved Thermal Design Procedure.
4 15.0.3.3 POWER DISTRIBUTION . The transient response of the reactor system is dependent on the initial 1 -power distribution. The nuclear _ design of the reactor _ core minimizes adverse power distribution through the placement of control rods and 0955T:1/6 15.0-8
.- ,-n -- ,-
operating instructions. Power distribution may be characterized by the radial factor (FAH) and the total peaking factor (F q). The peaking factor limits are given in the Technical Specifications. ( For transients which may be DNB limited the radial peaking factor is of importance. The radial peaking factor increases with decreasing power level due to rod insertion. This increase in FAH is included in the core limits illustrated in Figure 15.0.3-1. All transients that may be DNB limited are assumed to begin with a F c nsistent with the AH initial power level defined in the Technical Specifications. For transients which may be overpower limited the total peaking factor (Fq ) is of importance. All transients that may be overpower limited are assumed to begin with plant conditions including power distributions which are consistent with reactor operation as defined in the Technical Specifications. ( For overpower transients which are slow with respect to the fuel rod thermal time constant, for example, the Chemical and Volume Control System malfunction that results in a decrease in the baron concentration in the reactor coolant incident which lasts many minutes, and the excessive increase in secondary steam flow incident which may reach equilibrium without causing a reactor trip, the fuel rod thermal evaluations are performed as discussed in the McGuire FSAR Section 4.4. For overpower transients which are fast with respect to the fuel rod thermal time constant, for example, the uncontrolled rod cluster control ( essembly bank withdrawal from suberitical or low power startup and rod cluster cont:rol assembly ejection incidents which result in a large power rise over a few-seconds, a detailed fuel heat transfer' calculation-must be performed. Although the fuel rod thermal time constant is a function of system conditions, fuel burnup and rod power, a typical value at beginning-of-life for high power rods'is approximately 5 seconds, i
\
(b 0955T:1/6 .15.0 . _ _
l 15.0.4 REACTIVITY COEFFICIENTS ASSUMED IN THE ACCIDENT ANALYSES The transient response of the reactor system is dependent on reactivity feedback effects, in particular the moderator temperature coefficient and the Doppler power coefficient. In the analysis of certain events, conservatism requires the use of large reactivity coefficient values whereas in the analysis of other events, conservatism requires the use of small reactivity coefficient _ values. The values are given_in Table 15.0.3-2. Reference is made in that table to Figure 15.0.4-1, which shows the upper and lower bound Doppler power coefficients as a function of power used in the transient analysis. The justification for use of conservatively large versus small reactivity coefficient values are treated on an event-by-event basis. In some cases conservative combinations of parameters are used to bound the effects of core life, although these combinations may not represent possible realistic situations. 15.0.5 ROD CLUSTER CONTROL ASSEMBLY INSERTION CHARACTERISTICS The negative reactivity insertion following a reactor trip is a function of the position versus time of the rod cluster control assemblies and the variation in rod worth as a function of rod position. With respect
, to accident analyses, the critical parameter is the time of insertion up to the dashpot entry or approximately 85 percent of the rod cluster travel.
The rod cluster control assembly position versus time assumed in accident analyses is shown in Figure 15.0'.5-1. The rod cluster control assembly insertion time to dashpot entry is taken as 3.3 seconds. Drop time testing requirements are dependent on the type of cluster control assemblies actually used in the plant and are specified in the plant Technical Specifications. 0955T:1/6 15.0-10
Figure 15.0.5-2 shows the fraction of total negative reactivity insertion versus normalized rod position for a core where the axial distribution is skewed to the lower region of the core. An axial ( distribution which is skewed to the lower region of the core can arise from an unbalanced xenon distribution. This curve is used to compute the negative reactivity insertion versus time following a reactor trip which is input to all point kinetics core models used in transient analyses. The bottom skewed power distribution itself is not input into the point kinetics core model. There is inherent conservatism in the use of Figure 15.0.5-2 in that it is based on a skewed flux distribution which would exist relatively infrequently. For cases other than those associated with unbalanced xenon distributions, significant negative reactivity would have been inserted due to the more favorable axial distribution existing prior to trip. ( The normalized rod cluster control assembly negative reactivity insertion versus time is shown in Figure 15.0.5-3. The curve shown in this figure was obtained from Figures 15.0.5-1 and 15.0.5-2. A total negative reactivity insertion following a trip of 4 percent AK is assumed in the transient analyses except where specifically noted otherwise. For Figures 15.0.5-1 and 15.0.5-2, the rod cluster control assembly drop time is normalized to 3.3 seconds. The normalized rod cluster control assembly negative reactivity insertion versus time curve for an axial power distribution skewed to the bottom (Figure 15.0.5-3) is used in those transient analyses for which a point kinetics core model is used. Where special analyses-require use of three dimensional or axial one dimensional core models, the negative reactivity insertion resulting from the reactor trip.is l calculated directly by the reactor kinetics code and is not separable. from the other reactivity feedback effects. In this case, the rod L 0955T:1/6- 15.0-11
4 cluster control assembly position versus time of Figure 15.0.5-1 is used as code' input. 15.0.6 TRIP POINTS AND TIME DELAYS TO TRIP ASSUMED IN ACCIDENT ANALYSES ( A' reactor trip signal acts to open two trip breakers connected in series feeding power to the control rod arive mechanisms. The loss of power to the mechanism c.ils causes the mechanisms to release the rod cluster
- . control assemblies which then fall by gravity into the core. There are various instrumentation delays associated with each trip function, including delays in signal actuation, in opening the trip breakers, and in the release of the rods by the mechanisms. The total delay to trip is 4
defined as the time delay from the time that trip conditions are reached to the time the rods are free and begin to fall. Limiting trip
- setpoints assumed in accident analyses and the time delay assumed for each trip function are given in Table 15.0.6-1.
, Reference is made in that table to the Overtemperature and Overpower ' AT trips shown in Figure 15.0.3-1.- This figure presents the allowable Reactor Coolant loop average- temperature and AT for the design flow and power distribution as a function of primary coolant pressure. The boundaries of operation ~ defined by the overpower AT trip and the I overtemperature AT trip are represented as " protection lines" on 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 represented as a-line. The DNB lines represent the .
- locus of. conditions' for which the DNBR equals the limit' value (1.47 for-
[ the thimble cell and 1.49 for.the typical cell.) All points'below and. { to the-left of a DNB line for a given pressure have a' DNBR greater than the limit value. The diagram shows that DNB-is-prevented for all. cases - Lif the area enclosed with the maximum protection lines-is not traversed ! 'by the applicable DN3R'line'at any point. [ b
~
0955T:1/6= 15.0-12
The area of permissible operation (power, pressure and temperature) is bounded by the combination of reactor trips: high neutron flux (fixed setpoint); high pressure (fixed setpoint); low pressure (fixed setpoint); overpower and overtemperature AT (variable setpoints). The limit value, which was used as the DNBR limit for all accidents analyzed with the Improved Thermal Design Procedure (see Table 15.0.3-2), is conservative compared to the actual design DNBR value ( (1.31 for the thimble cell and 1.33 for the typical cell) required to meet the DNB design basis. The difference between the limiting trip point assumed for the analysis and the normal trip point represents an allowance for instrumentation channel error and setpoint error. Nominal trip setpoints are specified in the plant Technical Specifications. During plant startup tests it is demonstrated that actual instrument ti,me delays are equal to or less than the assumed values. Additionally, protection system channels are ( calibrated and instrument response times determined periodically in accordance with the Technical Specifications. 15.0.7 COMPUTER CODES UTILIZED e i' Summaries of some of the principal computer codes used in transient analyses are given below. The codes used in the analyses of each J transient have been listed in Table 15.0.3-2. 15.0.7.1 FACTRAN FACTRAN calcule,tes the transient temperature distribution in a cross. i section of a metal clad UO 2 fuel rod and the transient heat flux at the surface of the clad using as input the nuclear power and time-dependent coolant parameters (pressure, flow,' temperature, and density). The code uses a fuel model which exhibits the' following features simultaneously: L 0955T:1/6 15.0-13 l
i l l 1 l
- 1. A sufficiently large number of radial space increments to handle fast transients such as rod ejection accidents.
- 2. Material properties which are functions of temperature and a sophisticated fuel-to-clad gap heat transfer calculation.
- 3. The necessary calculations to handle post-DNB transients: film boiling heat transfer correlations, zircaloy-water reaction and partial melting of the materials.
FACTRAN is further discussed in Reference 1. 15.0.7.2 LOFTRAN The LOFTRAN program is used for studies of transient response of a PWR system to specified perturbations in process parameters. LOFTRAN ' simulates a multiloop system by a model containing reactor vessel, hot and cold leg piping, steam generator (tube and shell sides) and the pressurizer. The pressurizer heaters, spray, relief and safety valves are also considered in the program. Point model neutron kinetics, and reactivity effects of the moderator, fuel, boron and rods are included. The secondary side of the steam generator utilizes a homogeneous, saturated mixture for the thermal transients and a water level correlation for indication and control. The Reactor Protection System is simulated to include reactor trips on high neutron flux, Overtemperature AT, Overpower AT, high and low pressure, low flow, and high pressurizer level. Control systems are also simulated l including rod control, steam dump, feedwater. control and pressurizer ! pressure control. The Emergency Core Cooling System, including the accumulators and upper head injection, is also modeled. LOFTRAN is a versatile program which is suited to both accident-evaluation and control studies as well as parameter sizing. 0955T:1/6- 15.0-14
l I l 1 LOFTRAN also has the capability of calculating the transient value of DNBR based on the input from the core limits illustrated on Figure 15.0.3-1. The core limits represents the minimum value of DNBR as calculated for typical or thimble cell. LOFTRAN is further discussed in Reference 2. 15.0.7.3 TWINKLE The TWINKLE program is a multi-dimensional spatial neutron kinetics code, which was patterned after steady state codes presently used for reactor core design. The code uses an implicit finite-difference method to solve the two group transient neutron diffusion equations in one, two and three dimensions. The code uses six delayed neutron groups and contains a detailed multi-region fuel-clad-coolant heat transfer model for calculating pointwise Doppler and ;noderator feedback effects. The code handles up to 2000 spatial points, and performs its own steady ( state initialization. Aside from basic cross section data and thermal-hydraulic parameters, 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, pointwise 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 neutror, flux distribution. TWINKLE is further discussed in Reference 3. 15.0.7.4 THINC The THINC Code is described in Section 4.4 of the McGuire FSAR. 1 L 0955T:1/6 15.0-15
I l l 15.
0.8 REFERENCES
- 1. Hargrove, H. G., "FACTRAN - A Fortran-IV Code for Thermal Transients in a 00 Fuel Rod," WCAP-7908, June 1972.
2
- 2. Burnett, T. W. T., et al., "LOFTRAN Code Description," WCAP-7907, June 1972.
- 3. Risher, D. H., Jr. and Barry, R. F., " TWINKLE - A Multi-Dimensional Neutron Kinetics Computer Code," WCAP-7979-P-A (Proprietary), and
( WCAP-8028-A (Non-Proprietary), January 1975.
- 4. H. Chelemer, et al., " Improved Thermal Design Procedure,"
WCAP-8567-P (Proprietary), July,1975, and WCAP-8568 (Non-Proprietary), July 1975. (
- i L
.i
-1 i
0955T:1/6 ! 15.0-16 I
< TABLE 15.0.3-1 NUCLEAR STEAM SUPPLY SYSTEM POWER RATINGS Guaranteed NSSS thermal power output (MWt) 3425 Engineered safety features design rating 3579 (maximum calculated turbine rating) (MWt) Thermal power generated by the reactor coolant 14 pumps (Mdt) Reactor core- thermal power output (MWt) ~3411 C L i 0955T:1/6 15.0-17 I
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TABLE 15.0.3-3 NOMINAL VALUES OF PERTINENT PLANT PARAMETERS UTILIZED IN THE ACCIDENT ANALYSES
- Thermal output of NSSS (MWt)a 3425 Core inlet temperature (*F) 559.2 Vessel average temperature (*F) 588.6 Reactor Coolant System pressure (psia) 2250 Reactor Coolant flow per loop (gpm) 98,400 6 '
Total Reactor Coolant flow (10 lb/hr) 146.1 6 Steam flow from NSSS (10 lb/hr) 15.14 Steam pressure at steam generator outlet (psia) . 1000 Maximum steam moisture content (percent) 0.25 Assumed feedwater temperature at steam 440 generator inlet (*F) Average core heat flux (Btu /hr-ft2 ) 197,200 Control rod drop time (sec) 3.3 j ( *For accident analyses using the Improved Thermal Design Procedure l
"See Table 15.0.3-2 0955T:1/6 15.0-21
TABLE 15.0.3-4 NOMINAL VALUES OF PERTINENT PLANT PARAMETERS UTILIZED IN THE ACCIDENT ANALYSES
- Thermal output of NSSS (MWt)b 3425 Core inlet temperature ("F) 559.2 Vessel average temperature ('F) 588.6 Reactor :oolant System pressure (psia) -2250 Reactor Coolant flow per loop (gpm) 96,500 6
Total Reactor Coolant flow (10 lb/hr) 143.3 6 Steam flow from NSSS (10 lb/hr) 15.14 Steam pressure at steam generator outlet (psia) 1000 Maximum steam moisture content (percent) 0.25 Assumed feedwater temperature at steam 440 , generator inlet (*F) *
' Average core heat flux (Etu/hr-ft ) 197,200 Control rod drop time (sec) 3_. 3 "For accident analyses not using the Improved Thermal Design Procedure b
See Table 15.0.3-2 0955T:1/6 '15.0-22. f
TABLE 15.0.6-1 TRIP POINTS AND TIME DELAYS TO TRIP ASSUMED IN A,IDENT ANALYSIS Limiting Trip Point Assumed Time Delay Trip Function In Analysis (Seconds) Power range high neutron 118 percent 0.5 flux, high setting Power range high neutron 35 percent 0.5 l flux, low setting High neutron flux, P-8 85 percent 0.5 Overtemperature AT Variable, see 6.0 8 Figure 15.0.3-1 Overpower AT- Variable, see 6.0" Figure 15.0.3-1 High pressurizer pressure 2410 psig- 2.0 Low pressurizer pressure -1835 psig 2.0
- Total time delay (including RTO bypass loop fluid transport delay effect, bypass loop piping thermal capacity, RTD time reponse, and trip circuit, channel electronics delay) from the time the temperature difference in the coolant loops exceeds the. trip.setpoint until the rods are free to fall.
0955T:1/6 15.0 .
TABLE 15.0.6-1 (Page 2) TRIP POINTS AND TIME DELAYS TO TRIP ASSUMED IN ACCIDENT ANALYSIS (cont') Limiting Trip j Point Assumed Time Delay Trip Function In Analysis (Seconds) Low reactor coolant flow 87 percent loop 1.0 (From loop flow detectors) flow Undervoltage trip 68 percent nominal 1.5 Low-low steam generator 28.0 percent of 2.0 level narrow range
. level span High steam generator 87.4 percent of 2.0 level trip of the feedwater narrow range pumps and closure of feedwater level span system valves, and turbine trip e
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I c l DASHPOT 2 l 0 I ' I I ' ' I I I I I I 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0 1.1 1.2 1.3 1.4 NORMALIZED ROD DROP TIME TIME AFTER DROP BEGINS TIME TO TOP OF DASHPOT FIGURE 15.0.5-3 i Normalized RCOA Bank Reactivity Werth wrsus Normalized Orco Time
- L 15.0-29 i
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15.1 INCREASE IN HEAT REMOVAL BY THE SECONDADY SYSTEM A number of events have been postulated which could result in an increase in heat removal from the Reactor Coolant System by the Secondary System. Detailed analyses are presented for several such events which have been iden-tified as limiting cases. Discussions of the following Reactor Coolant System cooldown events are pre-( sented in this section:
- 1. Feedwater System malfunction causing an increase in feedwater flow.
- 2. Excessive increase in secondary steam flow.
- 3. Inadvertent opening of a steam generator relief or safety valve.
- 4. Steam System piping failure.
The above are considered to be ANS Condition II events, with the exception of a major steam system pipe break, which is considered to be an ANS Condition IV event. Section 15.0.1 centains a discussion of ANS classifications and appli-cable acceptance criteria. 15.1.1 FEEDWATER SYSTEM MALFUNCTION CAUSING AN INCREASE IN FEEDWATER FLOW 15.1.1.1 Identification of Causes and Accident Description Addition of excessive feedwater will cause an increase in core power by - decreasing reactor coolant temperature. Such transients are attenuated by the thermal capacity of the secondary plant and of the Reactor Coolant System (RCS). The overpower-overtemperature protection (neutron overpower, over- , teniperature and overpower AT trips) prevents any power increase which could lead to a DNBR less than the limit value. An example of excessive feedwater flow would be a full opening of a feedwater control valve due to a feedwater. control system malfunction or an operator , 1 0946T:1/6 15.1-1 I
error. At power this excess flow causas 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 tem-perature and thus a reactivity insertion due to the effects of the negative mocerator coefficient of reactivity. Continuous addition of excessive feedwater is prevented by the steam generator high-high level trip, which closes the feedwater valves. An increase in nor* , mal feedwater flow is classified as an ANS Condition II event, fault of mod-erate frequency. See Section 15.0.1 for a discussion of ANS Condition II events. 15.1.1.2 Analysis of Effects and Consecuences Method of Analysis The excessive heat removal due to a feedwater system malfunction transient is analyzed by using the detailed digital computer code LOFTRAN (Reference 1). This code simulates a multi-loop system, neutron kinetics, the pressurizer, pressurizer relief and safety valves, pressurizer spray, steam generator, and steam generator safety valves. The code computes pertinent plant variables including temperatures, pressures, and power level. The system is analyzed to demonstrate plant behavior in the event that exces-sive feedwater addition occurs due to a control system malfunction or operator error which allows a feedwater control valve to open fully. Two cases are analyzed as follows:
- 1. Accidental opening of one feedwater control valve with the reactor just critical at zero Icad conditions assuming a conservatively large negative moderator temperature coefficient.
L l l C 0946T:1/6 15.1-2
- 2. Accidental opening of one feedwater control valve with the reactor in manual control at full power.
( This accident is analyzed with the Improved Thermal Design Procedure as described in WCAP-8567. Plant characteristics and initial conditions are discussed in Section 15.0.3. The reactivity insertion rate following a feedwater system malfunction is calculated with the following assumptions: ( 1. Initial reactor power, pressure, and RCS temperatures are assumed to be at their nominal values. Uncertainties in initial conditions are included in the limit DNBR as described in WCAP-8567.
- 2. For the feedwater control valve accident at full power, one feedwater control valve is assumed to malfunction resulting in a step increase to 200 percent of nominal feedwater flow to one steam generator.
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- 3. For the feedwater control valve accident at zero load condition, a feed-( water control valve malfunction occurs which results in an increase in flow to one steam generator from zero to 100 perc'ent of the nominal full load value for one steam generator.
- 4. For the zero load condition, feedwater temperature is at a conservatively low value of 70*F.
- 5. No credit is taken for the heat capacity of the RCS and steam generator thick metal in attenuating the resulting plant cooldown.
- 6. The feedwater flow resulting from a fully open control. valve is terminated by a steam generator high-hign level trip signal which closes all feed-water control and isolation valves, trips the main feedwater pumps, and trips the turbine.
Plant characteristics and initial conditions are further discussed in Section 15.0.3. 0946T:1/6 15.1-3
Normal reactor control systems and Engineered Safety Systems are not required to function. The Reactor Protection System may function to trip the reactor due to an overpower condition. No single active failure will prevent opera-tion of the Reactor Protection System. Results ( In the case of an accidental full opening of one feedwater control valve with the reactor at zero power and the above mentioned assumptions, the maximum reactivity insertion rate is less than the maximum reactivity insertion rate analyzed in Section 15.4.1, " Uncontrolled Rod Cluster Control Assembly Bank Withdrawal from a Suberitical or Low Power Startup Condition," and therefore, the results of the analysis are not presented here. It should be noted that if the incident occurs with the unit just critical at no load, the ret.ctor may be tripped by the power range high neutron flux trip (low setting) set at approximately 25 percent of nominal full power. i The full power case (maximum reactivity feedback coefficients, manual rod control) gives the largest reactivity feedback and results in the greatest power increase. Assuming the reactor to be in the automatic rod control mode results in a'slightly less severe transient. The rod control system is, there-fore, not required to function for an excessive feedwater flow event. When the steam generator water level in the faulted loop reaches the high-high level setpoint, all feedwater isolation valves and feedwater pump discharge valves are automatically closed and the main feedwater pumps are tripped. This prevents continuous addition of the feedwater. In addition, a turbine trip is initiated. - Following turbine trip, the reactor will be tripped on a low-low steam genera-tor water level signal in the intact steam generators. If the reactor were in the automatic control mode, the control rods would be inserted at the maximum rate following turbine trip, and the ensuing transient'would then be similar to a loss of load (turbine trip event) as analyzed in Section 15.2.3. 0946T:]/6 15.1-4
Transient results, see Figures 15.1.1-1 and 15.1.1-2, show the increase in nuclear power and aT associated with the increased thermal load on the reac-tor. Following the turbine trip and feedwater isolation on the steam genera-( tor high-high level signal the reactor reaches a new stabilized condition at a reduced power level consistent with the reactivity parameters assumed to maxi-mize the initial increase in core power. The reactor is tripped on low-low steam generator water level if no action is taken by the operator to terminate the reduced power operation. The DNB ratio does not drop below the limit value. Following the reactor trip, the plant approaches a stabilized condi-tion; standard plant shutdown procedures may then be followed to further cool down the plant. Since the power level rises during the excessive feedwater flow incident, the fuel temperatures will also rise until after reactor trip occurs. The core i 1 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 well below the fuel melting temperature. The transient results show that 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 significantly above its initial value during the transient. The calculated sequence of events for this accident is shown in Table 15.1-1. 15.1.1.3. Conclusions The results of the analysis show that the DNB ratios encountered for an exces-
' sive feedwater. addition at power are above the limit value, hence, no fuel or clad damage is predicted. Additionally, it has been shown that the reactivity insertion rate which-occurs at no load conditions following excessive feed-water addition is less than the maximum value considered in the rod withdrawal from a suberitical condition analysis.
t 0946T:1/6 15.1-5
15.1.2 EXCESSIVE INCREASE IN SECONDARY STEAM FLOW 15.1.2.1 Identification of Causes and Accident Description An excessive increase in secondary system steam flow (excessive load increase incident) is defined as a rapid increase in steam flow that causes a power mismatch between the reactor core power and the steam generator load demand. The Reactor Control' System is designed to accommodate a 10 percent step load increase or a 5 percent per minute ramp load increase in the range of 15 to 100 percent of full power. Any loading rate in excess of these values may ( cause a reactor trip actuated by the Reactor Protection System. Steam flow increases greater than 10 percent are analyzed in Sections 15.1.3 and 15.1.4. This accident could result from either an administrative violation such as excessive loading by the operator or an equipment malfunction in the steam dump control or turbine speed control. . During power operation, steam dump to the condenser is controlled by reactor coolant condition signals, i.e., high reactor coolant temperature indicates a need for steam dump. A single controller malfunction does not cause steam dump; an interlock is provided which blocks the opening of the valves unless a large turbine load decreas' e or a turbine trip has occurred. Protection against an excessive load increase accident is provided by the following Reactor Protection System signals:
- 1. Overpower AT
- 2. Overten.perature AT
- 3. Power range high neutron flux An excessive load increase incident is considered to be an ANS Condition II event, fault of moderate frequency. See Section 15.0.1 for a discussion of .
Condition Il events.
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0946T:1/6 - 15.1-6
15.1.2.2 Analysis of Effects and Consecuences Method of Analysis This accident is analyzed using the LOFTRAN Code (Reference 1). The : ode simulates the neutron kinetics, Reactor Coolant System, pressurizer, pressurizer relief and safety valves, pressurizer spray, steam generator, steam generator safety valves, and feedwater system. The code computes pertinent plant variables including temperatures, pressures, and power level. Four cases are analyzed to demonstrate the plant behavior following a 10 per-cent step load increase from rated load. These cases are as follows:
- 1. Reactor control in manual with minimum moderator reactivity feedback.
- 2. Reactor control in manual with max.imum 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 value. This results in the largest amount of reactivity feedback due to changes in coolant temperature. A conservative limit on the turbine valve _ opening is assumed, and all cases are studied without credit taken for. pressurizer heaters. This accident is analyzed with the Improved Thermal Design Procedure as des-cribed in WCAP-8567. Plant characteristics and initial conditions are dis-cussed in Section 15.0.3. Initial reactor power, pressure, and RCS tempera-tures are assumed to be at their nominal values. Uncertainties in initial conditions are included in the limit DNBR as described in WCAP-8567. 0946T:1/6 15.1-7
Normal reactor control systems and Engineered Safety Systems are not reauired to function. The Reactor Protection System is assumed to be operable; how-ever, reactor trip is not encountered for many cases due to the error allow-ances assumed in the setpoints. No single active failure will pr, event the Reactor Protection System from performing its intended function. The cases which assume automatic rod control are analyzed to insure that the worst case is presented. The automatic function is not required. Results Figures 15.1.2-1 through 15.1.2-4 illustrate the transient with the reactor in the manual control mode. As expected, for the minimum moderator feedback case there is a slight power increase, and the average core temperature shows a large decrease. This results in a DNBR which increases above its initial value. For the maximum moderator feedback, manually controlled case there is a much larger increase in reactor power due to the moderator feedback. A reduc-tion in DNBR is experienced but DNBR remains above the limit value. For these cases, the plar,t rapidly reaches a stabilized condition at the higher power level. Normal plant operating procedures would then be followed to reduce power. Figures 15.1.2-5 through 15.1.2-8 illustrate the transient assuming the reactor is in the automatic control mode and no reactor trip signals occur. Both the minimum and maximum moderator feedback cases show that core power increases, thereby reducing the rate of decrease in coolant average tempera-ture and pressurizer pressure. For both of these cases, the minimum DNBR remains above the limit value. The excessive load increase incident is an overpower transient for which the fuel temperatures will rise. Reactor trip may not occur for some of the cases analyzed, and the plant reaches a new equilibrium condition at a higher power level corresponding to the increase in steam flow. 0946T:1/6 15.1-8
Since DNB does not occur at any time during the excessive load increase tran- l sients, the ability of the primary coolant to remove heat from the fuel rod is not reduced. Thus, the fuel cladding temperature does not rise significantly above its initial value during the transient. The calculated sequence of events for the excessive load increase incident is shown on Tat,le 15.1-1. 15.1.2.3 Conclusions The analysis presented above shows that for a ten percent step load increase the DNBR remains above the limit value, thereby precluding fuel or clad dam-age. The plant reaches a stabilized condition rapidly following the load increase. 15.1.3 INADVERTENT OPENING OF A STEAM GEffERATOR RELIEF OR SAFETY VALVE 15.1.3.1 Identification of Causes and Accident Description C The most severe core conditions resulting from an accidental depressurization of the main steam system are associated with an inadvertent opening of a single steam dump, relief, or safety valve. The analyses performed sssuming a rupture of a main steamline are given in Section 15.1.4. 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 energy removal from the RCS'causes a reduction of coolant temperature and pressure. In the presence of a negative moderator temperature coefficient, the cooldown results in an insertion of positive reactivity. The analysis is performed to demonstrate that the following criterion is satisfied: C 0946T:1/ 6 15.1-9
Assuming a stuck rod cluster control assembly, with offsite power avail-able, and assuming a single failure in the Engineered Safety Features System there will be no consequential damage to the core or the 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 depressurization of the secondary system is classified as an ANS Condition II event. See Section 15.0.1 for a discussion of Condition II events. The following systems provide the necessary protection against an accidental depressurization of the main steam system.
- 1. Safety Injection System actuation from any of the following:
- a. Two-out-of-three low steamline pressure signals in any one loop
- b. Two-out-of-four low pressurizer pressure signals.
1 i c. Two-out-of-three high containment pressure signals.
- 2. The overpower reactor trips (neutron flux and AT) and the reactor trip occurring in conjunction with receipt of the safety injection signal.
- 3. Redundant isolation of the main feedwater lines.
Sustained high feedwater flow would cause additional cooldown. Therefore, in addition to the normal control action which will close the main feed-water valves following reactor trip, a safety injection signal will rapidly close all feedwater control valves and feedwater isolation valves, trip the main feedwater pumps, and.close the feedyater pump discharge valves. 0946T:1/6 15.1-10
- 4. Trip of the fast-acting steamline stop valves (designed to close in less than 5 seconds) on:
- a. Two-out-of-three low steamline pressure signals in any one loop.
- b. Two-out-of-four high-high containment pressure signals.
- c. Two-out-of-three high negative steamline pressure rate signals in any one loop (used only during cooldown and heatup operations).
15.1.3.2 Analysis of Effects and Consecuences Method of Analysis The following a-alyses of a secondary system steam release are performed for this section. . ( 1. A full plant digital computer simulation using the LOFTRAN Code (Reference
- 1) to determine RCS temperature and pressure during the transient, and the effect of safety injection.
- 2. Analyses to determine that the DNB basis is met.
The following conditions are assumed to exist at the time of a secondary steam system release:
- 1. Enc *of-life shutdown margin at no-load, equilibrium xenon conditions, and with the most reactive rod cluster control assembly stuck in its fully withdrawn position. Operation of rod cluster control assembly banks during core burnup is restricted in such a way that cddition of positive reactiv-ity in a secondary system steam release accident will not lead to a more adverse condition than the case analyzed.
C 0946T:1/6 15.1-11
- 2. A negative moderator coefficient corresponding to the end-of-life redded core with the most reactive rod cluster control assembly in the fully withdrawn position. The variation of the a: ficient with temperature and pressure is included. The K,ff versus temperature at 1000 psi correspon-ding to the negative moderator temperature coefficient used is shown in Figure 15.1.3-1.
- 3. Minimum capability for injection of boric acid solution corresponding to the most restrictive single failure in the Safety Injection System. This corresponds to the flow delivered by one charging pump delivering its full contents to the cold leg header. Low concentration boric acid must be swept from the safety injection lines downstream of the boron injection tank isolation valves prior to the delivery of boric acid (2000 parts per million (ppm)) to the reactor coolant loops. This effect has been accounted for in the analysis.
- 4. The case studied is a steam flow of 248 pounds per second at 1100 pounds per square inch absolute (psia) with offsite power available. This is the maximum capacity of any single steam dump, relief, or safety valve. Ini-tial hot shutdown conditions at time zero are assumed since this repre-sents the most conservative initial condition.
Should the reactor be just critical or operating at power at the time of a steam release, the reactor will be tripped by the normal overpower protec-tion when power level reaches a trip point. Following a trip at power, the RCS contains more stored energy than at no-load, the average coolant temperature is higher than at no-load and there is appreciable energy stored in the fuel. Thus, the additional stored anergy is removed via the cooldown caused by the steam release before the no-load conditions of RCS temperature and shutdown margin assumed in the analyses are reached. After the additional stored energy has been removed, the cooldown and reactivity insertions proceed in the same manner as in the analysis which assumes no-load condition at time zero. However, since the initial steam generator water inventory is greatest at no-load, the magnitude and dura-tion of the RCS cooldown are greatest for steamline release occurring at no-load. 0946T:1/6 15.1-12
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i S. l In computing the steam flow, the Moody Curve (Reference 2) for f(L/D) = 0 is used. ! I
- 6. Perfect moisture separation in the steam generator is assumed. 1
]
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 15.1.3-2 and 15.1.3-3 show the transient results for a steam flow of 248 lb/sec at 1100 psia. . ( The assumed steam release is the maximum capacity of any single steam dump,
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relief, or safety valve. Safety injection is initiated automatically by low pressurizer pressure. Operation of one centrifugal charging pump is assumed. Boron solution at 2000 ppm enters the RCS providing sufficient negative reac-tivity to prevent core damage. .The transient is quite conservative 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 period _of about 5 minutes, the neglected stored energy is likely to have a significant effect in slowing the cooldown. 15.1.3.3 Conclusions The analysis shows that the criteria stated earlier in this section are. satis-fied. For an accidental depressurization of the main steam system, the DiiB design limits are not exceeded. This case is less limiting than the steamline rupture. case described in Section 15.1.4. L 0946T:1/6- 15.1-13 e
15.1.4 STEAM SYSTEM PIPING FAILURE 15.1.4.1 Identification of Causes and Accident Description The steam release arising from a rupture of a main steamline would result in an initial increase in steam flow which decreases during the accident as the steam pressure falls. The energy removal from the RCS causes a reduction of coolant temperature and pressure. In the presence of a negative moderator temperature coefficient, the cooldown results in an insertion of positive reactivi*y. If the most reactive rod cluster control assembly (RCCA) is assumed stuck in its fully withdrawn position after reactor trip, there is an increased possibility that the core will become critical and return to power. A return to power following a steamline rupture is a potential problem mainly because of the high power peaking factors which exist assuming the most reactive RCCA to be stuck in its fully withdrawn position. The core is ultimately shut down by the boric acid injec. tion delivered by the Safety Injection System. The analysis of a main steamline rupture is performed to demonstrate that the following criteria are satisfied: Assuming a stuck RCCA with or.without offsite power, and assuming a single failure in the engineered safety features, the core remains in place and intact. Radiation doses do not exceed the guidelines of 10CFR100. Although DNB and possible clad perforation following a steam pipe rupture are not necessarily unacceptable, the following analysis, in fact, shows that no DNB occurs for any rupture assuming the most reactive assembly stuck in its fully withdrawn position. A major steamline rupture is classified as an ANS Condition IV event. See Section 15.0.1 for a discussion of Condition IV events. C 0946T:1/6 15.1-14. I l j
j The major rupture of a steamline is the most limiting cooldown transient and is analyzed at zero power with no decay heat. Decay heat would retard the cooldown thereby reducing the return to power. A detailed analysis of this ( transient with the most limiting break size, a double-ended rupture, is pre-sented here. The following functions provide the protection for a steamline rupture:
- 1. Safety Injection System actuation from any of the following:
- a. Two-out-of-three low steamline pressure signals in any one loop.
- b. Two-out-of-four low pressurizer pressure signals.
- c. Two-out-of-three high containment pressure signals.
- 2. The overpower reactor trips (neutron flux and AT) and the reactor trip occurring in conjunction with receipt of the safety injection signal.
- 3. Redundant isolation of the main feedwater lines.
Sustained high feedwater flow would cause additional cooldown. Therefore, in addition to the normal control action which will close the main feed-water valves a safety injection signal will rapidly close all feedwater control valves and feedwater isolation valves, trip the main feedwater pumps, and close the feedwater pump discharge valves.
- 4. Trip of the fast acting steamline stop valves (designed to close in less than 5 seconds) on:
- a. Two-out-of-three low steamline pressure signals in any one loop.
- b. Two-out-of-four high-high containment pressure signals.
- c. Two-out-of-three high negative steamline pressure rate signals in any one loop (used only during cooldown and heatup operations).
0946T:1/6 15.1-15
i Fast-acting isolatic, valves are provided in each steamline; these valves will ( fully close within 10 seconds of a large break in the steamline. For breaks downstream of the isolation valves, closure of all valves would completely terminate the blowdown. For any break, in any location, no more than one steam generator would experience an uncontrolled blowdown even if one of the isola-tion valves fails to close. Steam flow is measured by monitoring dynamic head in nozzles located in the throat of the steam generator. The effective throat area of the nozzles is 1.4 square feet, which is considerably less than the main steam pipe area; thus, the nozzles also serve to limit the maximum steam flow for a break at any location. 1 15.1.4.2 Analysis of Effects and Consecuences Method of Analysis . The analysis of the steam pipe rupture has been performed to determine:
- 1. The core heat flux and RCS temperature and pressure resulting from the cooldown following the steamline break. The LOFTRAN code (Reference 1)
I has been used.
- 2. The thermal and hydraulic behavior of the core following a steamline break. A detailed thermal and hydraulic digital-computer code, THINC, has been used to determine if DNB occurs for the core conditions computed in item 1 above. '
The following conditions were assumed to exist at the time of a main steam break accident:
- 1. End-of-life shutdown margin at no-load, equilibrium xenon conditions, and the most reactive RCCA stuck ir, its fully withdrawn position. Operation of the control rod banks during core burnup is restricted in such a way that addition of positive reactivity in a steamline break accident will not lead to a more adverse condition than the case analyzed. -
0946T:1/6 15.1-16
- 2. A negative moderator coefficient corresponding to the end-of-life redded core with the most reactive RCCA in the fully withdrawn position. The variation of the coefficient with temperature and pressure has been inclu-ded. The K,ff versus temperature at 1000 psi corresponding to the nega-tive moderator temperature coefficient used is shown in Figure 15.1.3-1.
The effect of power generation in the core on overall reactivity is shown in Figure 15.1.4-1. ( The core properties associated with the sector nearest the affected steam
. generator and those associated with the remaining sector were conserva-tively combined to obtain average core properties for reactivity feedback calculations. Further, it was conservatively assumed that the core power distribution was uniform. These two conditicas cause underprediction of the reactivity feedback in the high power region near the stuck rod. To verify the conservatism of this method, the reactivity as well as the power distribution was checked for the lietting statepoints for the cases analyzed.
This core analysis considered the Doppler reactivity from the high fuel temperature near the stuck RCCA, moderator feedback from the high water enthalpy near the stuck RCCA, power redistribution and non-uniform core inlet temper.ature effects. . For cases in which steam generation occurs in the high flux regions of the core, the effect of void formation was also included. It was determined that the reactivity employed in the kinetics analysis was always larger than the reactivity calculated including the above local effects for the statepoints. These results verify conserva-tism; i.e., underprediction of negative reactivity feedback from power generation.
- 3. Minimum capability for injection of boric. acid (2000 ppm) solution corresponding to the most restrictive single failure in the Safety Injection System. The Emergency Core Cooling System consists of three systems: 1) the passive accumulators, 2) the Residual Heat Removal System, and 3) the Safety Injection System. Only the Safety Injection System and the accumulators are modeled for the steamline break accident analysis.
L 0946T:1/6 15.1-17
l The actual modeling of the Safety Injection System in LOFTRAN is described in Reference [1]. The flow corresponds to that delivered by one charging pump delivering its full flow to the cold leg header. No credit has been taken for the low concentration borated water, which must be swept from the lines downstream of the boron injection tank isolation valves prior to the delivery of high concentration boric acid to the Reactor Coolant loops. For the cases where offsite power is assumed, the sequence of events in the Safety Injection System is the following. After the generation of the safety injection signal (appropriate delays for instrumentation, logic, and signal transport included), the appropriate valves begin to operate and the high head safety injection pump starts. In 10 seconds, the valves are assumed to be in their final position and the pump is assumed to be at full speed. The volume containing the low concentration borated water is swept bafore the 2000 ppm reaches the core. This delay, described above, is inherently included in the modeling. , In cases where offsite power is not available, an additional 10 second delay is assumed to start the diesels and to load the necessary safety injection equipment onto them.
- 4. Design value of the steam generator heat transfer coefficient including allowance for steam generator tube fouling.
- 5. Since the steam generators are provided with integral flow restrictors with a 1.4 square foot throat area, any rupture with a break area greater than 1.4 square feet, regardless of location, would have the same effect on the NSSS as the 1.4-square foot break. The following cases have been considered in determining the core power and RCS transients:
- a. Complete severance of a pipe, with the plant initially at no-load t conditions, full reactor coolant flow with offsite power available.
- b. Case (a) with loss of offsite power simultaneous with the steamline break and initiation of the safety injection signal. Loss of offsite power results in reactor coolant pump coastdown.
T 0946T:1/6 15.1-18
- 6. Power peaking factors corresponding to one stuck RCCA and non-uniform core inlet coolant temperatures are determined at end of core life. The cold-est core inlet temperatures are assumed to occur in the sector with the stuck rod. The power peaking factors account for the effect of the local void in the region of the stuck control assembly during the return to power phase following the steamline break. This void in conjunction with the large negative moderator coefficient partially offsets the effect of the stuck assembly. The power peaking factors depend upon the core power, temperature, pressure, and flow, and, thus, are different for each case studied.
The core parameters used for each of the two cases correspond to values determined from the respective transient analysis. Both cases above assume initial hot shutdown conditions at time zero since this represents the most pessimistic. initial condition. Should the reac-tor be just critical or operating at power at the time of a steamline break, the reacte will be tripped by the normal overpower protection system when power level reaches a trip point. Following a trip at power, the RCS contains more stored energy than at no-load, the average coolant temperature is higher than at no-load and there is appreciable energy stored in the fuel. Thus, the additional stored energy is removed via the cooldown caused by the steamline break before the no-load conditions of RCS temperature and shutdown margin assumed in the analyses are reached. After the additional stored energy has been removed, the cooldown and reactivity insertions proceed in the same manner as in the analysis which. assumes no-load condition at time zero. A spectrum of steamline breaks at various power levels has been analyzed in Reference 3.
- 7. In computing the steam flow during a steamline break, the Moody Curve (Reference 2) for f(L/0) = 0 is used.
- 8. The Upper Head Injection (UHI) is simulated. The actuation pressure for the UHI is near the saturation pressure for the inactive coolant in the upper head. The insurge of cold UHI water keeps this inactive coolant-0946T:1/6 15.1-19
from flashing and thus retards pressure decrease. The net effect of UHI is to retard the pressure decrease which in turn permits less safety injection flow into the core due to the relatively higher back pressure. These effects are very small and results are not significantly affected. These assumptions are discussed more fully in Reference 3. Results 4 The calculated sequence of events for both cases analyzed is shown on Table 15.1-1.
; The results presented are a conservative indication of the events which would occur assuming a steamline rupture since it is postulated that all of the conditions described above occur simultaneously.
Core Power and Reactor Coolant System Transient Figures 15.1.4-2 through 15.1-4-4 show the RCS transient and core heat flux following a main steamline rupture (complete severance of a pipe) at initial no-load condition (case a). , Offsite power is assumed available so that full reactor coolant flow exists. The transient shown assumes an uncontrolled steam release from only one steam generator. Should the core be critical at near zero power when the rupture occurs the initiation of safety injection by low steamline pressure will trip the reactor. Steam release from more than one steam generator will be preven-ted by automatic trip of the fast acting isolation valves in the steamlines by low steamline pressure signals, high-high containment pressure signals, or high negative steamline pressure rate signals. Even with the failure of one valve, release is limited to no more than 10 seconds for the other steam gen-- erators while the one_ generator blows down. The steamline stop valves are designed to be fully closed in less than 5 seconds from receipt of a closure signal.
-l l
0946T:1/6 15.1-20 i
As shown in Figure 15.1.4-3 the cere attains criticality with the RCCAs inser-ted (with the design shutdown assuming one stuck RCCA) shortly before boron solution at 2000 enters the RCS. The continued addition of boron results in a peak core power significantly' lower than the nominal full power value. The calculation assumes the boric acid is mixed with, and diluted by, the water flowing in the RCS prior to entering the. reactor core. The concentra-tion after mixing 'epends upon the relative flow rates in the RCS, from the ( UHI, and from the .afety injection System. The variation of mass flow rate in the RCS due to watar density changes is included in the calculation as is the variation of flow rate in the Safety Injection System due to changes in the RCS pressure. The Safety Injection System flow calculation includes the line losses in the system as well as the pump head curve. Figures 15.1.4-5 through 15.1.4-7 show the salient parameters for case b, which corresponds to the case discussed above with additional loss of offsite power at the time the safety injection signal is generated. The Safety Injec-( tion System delay time includes 10 seconds to start the diesel in addition to 10 seconds to start the safety injection pump and open the valves. Critical-ity is achieved later and the core power increase is slower than in the simi-lar case with offsite power available. The ability of the emptying steam generator to extract heat from the RCS is reduced by.the decreased flow in the RCS. The peak power remains well below the nominal full power value. It should be noted that following a steamline break only one steam generator blows down completely. Thus, the remaining steam generators are still avail-able for dissipation of decay heat after the initial transient is over. In the case of loss of offsite power this heat is removed to the atmosphere via the steamline safety valves. Margin to Critical Heat Flux A DNB analysis was performed for both of these cases. It was found that both cases have a minimum DNBR greater than the limit value. L 0946T:1/6 15.1-21
15.1.4.3. Conclusions Although DNB and possible cladding perforation following a steam pipe rupture are not necessarily unacceptable and not precluded by the criteria, the above analysis, in fact, shows that the DNB design basis is met. ( 15.
1.5 REFERENCES
- 1. Burnett, T. W. T., et al., "LOFTRAN Code Description," WCAP-7907, June 1972.
- 2. Moody, F. S.,
" Transactions of the ASME, Journal of Heat Transfer," Fig-ure 3, page 134, February 1965.
- 3. Hollingsworth, S. D. and Wood,.D. C., " Reactor Core Response to Excessive Secondary Steam Releases," WCAP-9226, Revision 1, (Proprietary), January, 1978, and WCAP-9227, Revision 1, (Non-Proprietary), January 1978.
C L c 0946T:1/6 15.1-22
2 TABLE 15.1-1 (Page 1) TIME SEQUENCE OF EVENTS FOR INCIDENTS WHICH CAUSE AN INCREASE IN HEAT REMOVAL BY THE SECONDADY SYSTEM Accident Event Time (sec.) Excessive Feedwater One main feedwater control 0.0 flow'at full power valve fails fully open High-high steam generator 27 water level signal generated Turbine trip occurs due to 30 high-high steam generator. level Minimum DNBR occurs 30
- Feedwater isolation valves 36 close
~
Low-low steam generator level 102-1
. reactor trip setpoint reached in intact steam generators l Excessive Increase in Secondary Steam Flow
- 1. Manual Reactor 10 percent. step load increase 1 '
0 . 0 .- l' Control (Minimum ! moderator feedback)
.l
~
~j a
0946T:1/6 '15.1-23 j
TABLE 15.1-1 (Page 2) TIME SEQUENCE OF EVENTS FOR INCIDENTS WHICH CAUSE AN INCREASE IN HEAT REMOVAL BY THE SECONDARY SYSTEM Accident. Event Time (sec.) Equilibrium conditions 100 reached (approximate time only)
- 2. Manual Reactor 10 percent step load increase 0.0 Control (Maximum moderator feedback)
Equilibrium conditions 75 reached (approximate time only)
- 3. Automatic Reactor 10 percent step load increase 0.0 Control (Minimum Control (Minimum moderator feedback) moderator feedback)
Equilibrium conditions 50 reached (approximate time only)
~
- 4. Automatic Reactor 10 percent step load increase 0.0 '
Control (Maximum moderator feedback) ' Equilibrium conditions. 75~ reached (approximate. time only) 0946T:1/6 ~15.1 . . .
. .- - - - - - _ _ = _ - _ _ - _ - - _ .-- -_-_ _-
TABLE 15,1-1 (Page 3) TIME SE00ENCE OF EVENTS FOR INCIDENTS WHICH ( CAUSE AN INCREASE IN HEAT REMOVAL BY THE SECONDARY SYSTEM Accident Event Time (sec.) Inadvertent opening of Inadvertent opening of 0.0 a steam generator relief one main steam safety or safety valve or relief valve Pressurizer empties 163 Boron reaches 200 core .. ( Steam system piping failure
- 1. Case a Steamline ruptures 0 1
Pressurizer empties 14. Criticality attained 19' Boron reaches 30 core
- 2. Case b Steamline ruptures -0 Pressurizer empties 16' i
'l Criticality. attained 22 Baron reaches- , 37 core 0946T:1/6 15.1-25~ i
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Feeeweer Control valve Malfunction 15.1-26
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15.2 DECREASE IN HEAT REMOVAL BY THE SECONDARY SYSTEM i A number of transients and accidents have been postulated which could result ( in a reduction of the capacity of the secondary system to remove heat generated in the Reactor Coolant System (RCS). These events are discussed in this section. Detailed analyses are presented for several such events which have been identified as more limiting than the others. ( Discussions of the following RCS coolant heatup events are presented in Section 15.2:
- 1. Steam Pressure Regulator Malfunction
- 2. Loss of External Lead
- 3. Turbine Trip
' 4. Inadvertent Closure of Main Steam Isolation Valves
- 5. Loss of Condenser Vacuum and Other Events Resulting in Turbine Trip
~
- 6. Loss of Non-Emergency AC Power to the Station Auxiliaries
- 7. Loss of Normal Feedwater Flow
- 8. Feedwater System Pipe Break The above items are considered to be ANS Condition II events, with the exception of a Feedwater System Pipe Break, which is considered to be an ANS Condition IV event. Section 15.0.1 contains a discussion of ANS classification and applicable acceptance criteria.
15.2.1 STEAM PRESSURE REGULATOR MALFUNCTION OR FAILURE THAT RESULTS IN DECREASING STEAM FLOW There are no pressure regulators whose failure or malfunction could cause a steam flow transient. 15.2.2 LOSS OF EXTERNAL LOAD L 0965T:1/6 15.2-1
15.2.2.1 Identif"tcation of Causes and Accident Descriotion A major load loss on the plant can result from loss of external electrical load due to some electrical system disturbance. Offsite AC power remains available to operate plant components such as the reactor coolant pumps; as a result, the onsite emergency diesel generators are not required to function for this event. Following the loss of generator load, an immediate fast closure of the turbine control valves will occur. This will cause a sudden reduction in steam flow, resulting in an increase in pressure and temperature in the steam generator shell. As a result, the heat transfer rate in the steam generator is reduced, causing the reactor coolant temperature to rise, which in turn causes coolant expansion, pressurizer insurge, and RCS pressure rise. For a loss of external electrical load without subsequent turbine trip, the plant would be expected to trip from the , Reactor Protection System if a safety limit were approached. With full load rejection capability plant operation would be expected to continue without a reactor trip. A continued steam load of approximately 5 percent would exist after total loss of external electrical load because of the steam demand of plant auxiliaries. In the event that a safety limit is approached, protection would be provided by the high pressurizer pressure and overtemperature AT trips. Power and frequency relays associated with the reactor coolant pump provide no additional safety function for this event. Following a complete loss of load, the maximum turbine overspeed would be approximately 8 to 9 percent, resulting in an overfrequency of less than 6 Hz. This resulting overfrequency is not expected to damage the sensors (non-NSSS) in any way. However, it is noted that frequency testing of this equipment is required by'the Technical Specifications. Any degradation in their performance could be ascertained at that time. Any increased frequency to.the reactor coolant pump motors will result in slightly increased flowrate and subsequent additional margin to safety limits. For postulated loss of load and subsequent turbine generator overspeed, any overfrequency condition is not seen by safety related pump motors, Reactor Protection System equipment, or other safeguards loads. 0965T:1/6 15.2-2
Safeguards loads are supplied from offsite power or, alternatively, from emergency diesels. Reactor Protection System equipment is supplied from the inverters; the inverters are supplied from a DC bus energized from batteries ( or by a rectified AC voltage from safeguards buses. In the event the steam dump valves fail to open following a large loss of load, the steam generator safety valves may lift and the reactor may be tripped by the high pressurizer pressure signal, the high pressurizer water level signal, or the overtemperature AT signal. The steam generator shell side pressure and reactor coolant temperatures will increase rapidly. The pressurizer safety valves and steam generator safety valves are, however, sized to protect the Reactor Coolant System (RCS) and steam generator against overpressurization for all load losses without assuming the operation of the steam dump system, pressurizer spray, pressurizer power-operated relief valves, or automatic rod cluster control assembly control.
~
The steam generator safety valve capac1ty is sized to remove the steam flow at ( the Engineered Safety Features rating (105 percent of steam flow at rated power) from the steam generator without exceeding 110 percent of the steam system design pressure. The pressurizer safety valve capacity is sized based on a complete loss of heat sink with the plant initially operating at the maximum calculated turbine load along with operation of the steam generator safety valves. The pressurizer safety valves are then able to relieve sufficient steam to maintain the RCS pressure within 110 percent of the RCS design pressure. A more complete discussion of overpressure protection can be found in Reference 1. A loss of external load is classified as an ANS Condition II event, fault of 4
~moderate frequency. See Section 15.n.1 for a discussion of Condition II events.
A loss of external load event results in an NSSS transient that is less severe than a ',urbine trip event (see Section 15.2.3). Therefore, a, detailed transient analysis is not-presented for the loss of external' load. 0965T:1/6 15.2-3
The primary-side transient is caused by a decrease in heat transfer capability from primary to secondary due to a rapid termination of steam flow to the turbine, accompanied by an automatic reduction of feedwater flow. (Should feed flow not ce reduced, a larger heat sink would be available and the transient would be less severe). Termination of steam flow to the turbine following a loss of external load occurs due to automatic fast closure of the turbine control valves in approximately 0.3 seconds. Following a turbine trip event, termination of steam flow occurs via turbine stop valve closure, which occurs in approximately 0.1 seconds. Therefore, the transient in primary pressure, temperature, and water volume will be less severe for the loss of external load than for the turbine trip due to a slightly slower loss of heat transfer capability. 15.2.2.2 Analysis of Effects and Consecuences Method of Analysis . Refer to Section 15.2.3.2 for the method used to analyze the limiting transient (turbine trip) in this grouping of events. The results of the turbine trip event analysis are more severe than those expected for the loss of external load, as discussed in 15.2.2.1. Normal reactor control systems and Engineered Safety Systems are not required to function. The Auxiliary Feedwater System may, however, be automatically actuated following a loss of main feedwater; this will further mitigate the effects of the transient. The Reactor Protection System may be required to function following a complete - loss of external load to terminate core heat input and prevent DNB. Depending on the magnitude of the load loss, pressurizer safety valves and/or steam generator safety valves may be required to open to maintain system pressure below alloweble limits. No single active failure will prevent operation of any system required to function. 0965T:1/6 15.2-4
15.2.2.3 Conclusions Based on results obtained for the turbine trip event (Section 15.2.3) and considerations described in Section 15.2.2.1, the applicable acceptance criteria for a loss of external load event are met. 15.2.3 TURBINE TRIP 15.2.3.1 Identification of Causes and Accident Description For 'a turbine trip event, the turbine stop va'lves close rapidly (typically 0.1 sec.) on loss of trip-fluid pressure actuated by one of a number of-possible turbine trip signals. Turbine-trip initiation signals include:
- 1. Generator Trip
- 2. Low Condenser Vacuum ,
- 3. Loss of Lubricating 011
( 4. 5. 6. Turbine Thrust Bearing Failure Turbine Overspeed Main Steam Reheat High Level
- 7. Manual Trip Upon initiation of stop valve closure, steam flow to the turbine stops abruptly. Sensors on the stop valves detect the turbine trip and initiate steam dump. The loss of steam flow results in an almost immediate rise in secondary system temperature and pressure with a resultant primary system transient as described in Section 15.2.2.1 for the loss of external load event. A more severe transient occurs for the turbine. trip event due'to the more rapid loss of steam flow caused by the more rapid valve closure.
The automatic steam dump system would normally accomodate the excess steam generation. Reactor coolant temperatures and pressure do not significantly increase if the steam dump system and pressurizer pressure control system are functioni,g properly. If the turbine condenser was not available, the excess steam generation would be dumped to the atmosphere and main feedwater flow l
~
096 5T:1/6 15.2-5
I f i j would be lost. For this situation, feedwater flow would be maintained by the .) 3 Auxiliary Feedwater System to insure adequate residual and decay heat removal capability. Should the steam dump system fail to operate, the steam generator i safety valves may lift to provide pressure control. See 15.2.2.1 for a 4 further discussion of the transient. A turbine trip is classified as an ANS Condition II event, fault of moderate frequency. See Section 15.0.1 for a discussion of Condition II events. j A turbine-trip event is more limiting than loss of external load, loss of i condenser vacuum, and other turbine-trip events. As such, this event has been j analyzed in detail. Results and discussion of the analysis are presented in ) St: tion 15.2.3.2. ! 15.2.3.2 Analysis of Effects and Consecuences t 1 1 Method of Analysis l In this analysis, the behavior of the unit is evaluated for a complete loss of steam load from full power primarily to show the adequacy of the pressure relieving devices and also to demonstrate core protection margins. _The reactor is not tripped until conditions in the RCS result in a trip. No credit is taken for steam dump. Ma'in feedwater flow is terminated at the time of turbine trip, with no credit taken for auxiliary feedwater to mitigate the consequences of the transient. ~ The turbine trip transients' are analyzed by employing the detailed digital computer program LOFTRAN (Reference 2). The program simulates the neutron kinetics, RCS, pressurizer, pressurizer relief and safety valves, pressurizer
- spray, steam generator, and steam generator safety valves. The program computes pertinent plant varibles including temperatures, pressures, and power .
4-level.
. This accident is analyzed with the Improved Thermal Design Procedure.as.
described in WCAP-8567, Plant characteristics and. initial : condition's' are
~
' '{ discussed in Section 15.0.3. 0965T:1/6. q
.15.2-6
=
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i f Maaor assumptions are summarized below: 1, Initial Operating Conditions - initial reactor power, pressure, and RCS ( temperatures are assumed to be at their nominal values. Uncertainties in initial conditions are included in the limit DNBR as described in WCAP-8567. t
- 2. Moderator and Doppler Coefficients of Reactivity - the turbine trip is ,
analyzed with both maximum and minimum reactivity feedback. The maximum feedback cases assume a large negative moderator temperature coefficient and the most negative Doppler power coefficient. The minimum feedback cases assume a positive moderator temperature coefficient and the least negative Doppler coefficients. (see Figure 15.0.4-1).
- 3. Reactor Control - from the standpoint of the maximum pressures attained it is conservative to assume that the reactor is in manual control. If the reactor were in automatic control, the control rod banks would move prior
( 4. to trip and reduce the severity of the transient. Steam Release - no credit is taken for the operation of the steam dump system or steam generator -power-operated relief valves. The steam generator pressure rises to the safety valve setpoint where steam release through the safety valves limits secondary steam pressure at the setpoint value.
- 5. Pressurizer Spray and Power-Operated Relief Valves - two cases for both the minimum and maximum reactivity feedback cases are analyzed:
- a. Full credit is taken for the effect of pressurizer spray and power-operated relief valves in reducing or limiting the coolant pressure. Safety valves are also available.
- b. No credit is taken for the effect of pressurizer spray and power operated relief valves in reducing or limiting the coolant pressure.
Safety valves are operable. L 0965T:1/6 15.2-7
- 6. Feedwater Flow - main feedwater flow to the steam generators is assumed to be lost at the time of turbine trip. No cr2dit is taken for auxiliary feedwater flow since a stabilized plant condition will be reached before auxiliary feedwater initiation is normally assumed to occur; however, the auxiliary feedwater pumps would be expected to start on a trip of the main feedwater pumps. The auxiliary feedwater flow would remove core decay heat following plant stabilization.
- 7. Reactor trip is actuated by the first Reactor Protection System trip setpoint reached. Trip signals are expected due to high pressurizer (
pressure, overtemperature AT, high pressurizer water level, and low-low steam generator water level. Except as discussed above, normal Reactor Coolant System and Engineered Safety Systems are not required to function. Cases are presented in which pressurizer spray and powe'r operated relief valves ar,e assumed, but the more limiting cases where these functions are not assumed are also presented. The Reactor Protection System 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. Results The transient responses for a turbine trip from full power operation are shown j for four cases: two cases for minimum reactivity feedback and two cases for maximum reactivity feedback (Figures 15.2.3-1 through 15.2.3-8). The calculated sequence of events for the accident is shown in Table 15.2-1. l l Figures 15.2.3-1 and 15.2.3-2 show the transient responses for the total loss of steam load with a positive moderator temperature coefficient assuming full - credit for the pressurizer spray and pressurizer power-operated relief valves. No credit is taken for the steam dump. The reactor is tripped by the' 0965T:1/6 15.2-8
high pressurizer pressure trip channel. The minimum DNBR remains well above l the limit value. The pressurizer safety valves are actuated for this case and maintain system pressure below 110 percent of the design value. The Steam ( Generator Safety Valves limit the secondary steam conditions to saturation at the safety valve setpoint. Figures 15.2.3-3 and 15.2.3-4 show the response for the total loss of steam load with a large negative moderator temperature coefficient. All other plant parameters are the same as the above. The DNBR increases throughout the transient and never drops below its initial value. pressurizer relief valves and steam generator safety valves prevent overpressurization in primary and secondary systems, respectively. The pressurizer safety valves are not actuated for this case. In the event that feedwater flow is not terminated at the time of turbine trip for this case, flow would continue under automatic control with the reactor at a reduced power. The operator would take action to terminate the transient ( and bring the plant to a stabilized condition. If no action were taken by the operator the reduced power operation would continue until the condenser hotwell was emptied. A low-low steam generator water level reactor trip would be generated along with auxiliary feedwater initiation signals. Auxiliary feedwater would then be used to remove decay heat with the results less severe than those presented in Section 15.2.7, Loss of Normal Feedwater Flow. The turbine trip accident was also studied assuming the plant to be initially operating at full power with no credit taken for the pressurizer spray, pressurizer power-operated relief valves, or steam dump. Th. reactor is tripped on the high pressurizer pressure signal. Figures 15.2.3-5 and 15.2.3-6 show the transients with a positive moderator coefficient. The neutron flux remains essentially constant at full power until the reactor is tripped. The DNBR never goes below the initial value throughout the transient. In this case the pressurizer safety valves are actuated,. and maintain system pressure below 110 percent of the design value. 0965T:1/6 15.2-9
Figures 15.2.3-7 and 15.2.3-8 are the transients with maxiraum reactivity 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 are actuated to limit primary pressure.
- Reference 1 presents additional results of analysis for a complete loss of heat sink including loss of main feedwater. This analysis shows the overpressure protection that is afforded by the pressurizer and steam generator safety valves.
15.2.3.3 Conclusions Results of the analyses, including those in Reference 1, show that the plant design is such that a turbine 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 max,imum pressures to within the design limits. i The integrity of the core is maintained by operation of the Reactor Protection System, i.e., the DNBR will be maintained above the limiting value. The applicable acceptance criteria as listed in Section 15.0.1 have been met. 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 Inadvertent closure of the main steam isolation valves would result in a turbine trip. Turbine trips are discussed in Section 15.2.3. 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 15.2.3. A loss of condenser vacuum would preclude the use of steam dump to the condenser; however, since steam dump is assumed not to be available in the turbine trip 0965T:1/6 15.2-10 j
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 listec in Section 15.2.3.1 are covered by Section 15.2.3. Possible overfrequency 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. i ( 15.2.6 LOSS OF NON-EMERGENCY AC POWER TO THE PLANT AUXILIARIES 15.2.6.1 Identification of Causes and Accident Description A complete loss of non-emergency AC power may result in the loss of all power to the plant auxiliaries, i.e., the reactor coolant pumps, condensate pumps, etc. The loss of power may be caused by a complete loss of the offsite grid accompanied by a turbine generator trip at the station, or by a loss of the onsite AC distribution system. This transient is more severe than the turbine trip event analyzed in Section 15.2.3 because for this case the decrease in heat removal by the secondary system is accompanied by a flow coastdown which further reduces the capacity of the primary coolant to remove heat from the core. The reactor will trip: (1) upon reaching one of the trip setpoints in the primary and secondary systems as result of tne flow coastdown and decrease in secondary heat removal; or (2) due to loss of power to the control rod drive mechanisms as a result of the loss of power to the plant. Following a loss of AC power with turbine and reactor trips, the sequence described below will occur:
- 1. Plant vital instruments are supplied from emergency DC power sources.
t
- 2. As the steam system pressure rises following the trip, the steam generator power-operated relief valves may be automatically opened to the 0965T:1/6 15.2-11
i atmosphere. The condenser is assumed not to be available for steam dump. If the steam flow rate through the power relief valves is not available, the steam generator safety valves may lift to dissipate the sensible heat of the fuel and coolant plus the residual decay heat produced in the reactor.
- 3. As the no-load temperature is approached, the steam generator power-operated relief valves (or safety valves, if the power operated relief valves are not available) are used to dissipate the residual decay heat and to maintain the plant at the hot shutdown condition. (
4 The standby diesel generators, started on loss of voltage on the plant emergency busses, begin to supply plant vital loads. The Auxiliary Feedwater System is started automatically as follows: Two motor-driven auxiliary feedwater pumps are started on any of the following:
- a. Low-low level in any generator
- b. Any safety injection signal
- c. Loss of offsite power
- d. Trip of all main feedwater pumps
- e. Manual actuatior.
One turbine-driven auxiliary feedwater pump is started on any of the following:
- a. Low-low level in any two steam generators
- b. Loss of offsite power '
- c. Manual actuation The motor-driven auxiliary feedwater pumps are supplied power by the diesels and the turbine-driven pamp utilizes steam from the secondary system. .Both type pumps are designed to supply rated flow within' one minute of the initiating signal even if a loss of all non-emergency AC power o:: curs 0965T:1/6 '15.2-12
simultaneously with loss of normal feedwater. The turbine exhausts the secondary steam to the atmosphere. The pumps take suction from the auxiliary feedwater storage tank for delivery to the steam generators. Upon the loss of power to the reactor ecolant pumps, coolant flow necessary for core cooling and the removal of residual heat is maintained by natural circulation in the reactor coolant loops. A loss of non-emergency AC power to the station auxiliaries is classified as an ANS Condition II event, fault of moderate frecuency. See Section 15.0.1 for a discussion of Condition II events A loss of AC power event, as described above, is a more limitin5 event than the turbine-trip initiated decrease irt secondary heat removal without loss of AC power, which analyzed in Section 15.2.3. However, a loss of AC power to ( the plant auxiliaries as postulated above could result in a loss of normal feedwater if the condensate pumps lose their power supply. Following the reactor coolant pump coastdown caused by the loss of AC power, the natural circulation capability of the RCS will remove residual and decay heat from t.he core, aided by auxiliary feedwater in the secondary system. An analysis is presented here to show that the natural circulation flow in the RCS following a loss of AC power event is sufficient to remove residual heat from the core. 15.2.6.2 Analysis of Effects and Consecuences Method of Analysis A detailed analysis using the LOFTRAN Code (Reference 2) is performed to obtain the plant transient following a station blackout. The simulation describes the plant thermal kinetics, Reactor Coolant System _ (RCS) including the natural circulation, pressurizer, steam generators and feedwater system. 0965T:1/6 15.2-13
)
l l l _______a
The digital program computes pertinent variables including the steam generator level, pressurizer water level, and reactor coolant average temperature. Assumptions made in the analysis are:
- 1. The plant is initially operating at 102 percent of the Engineered Safety.
Features design rating. ,
- 2. A conservative core residual heat generation based upon long term operation at the intial power level preceding the trip.
- 3. A heat transfer coefficient in the steam generator associated with RCS natural circulation, following the reactor coolant pump coastdown.
- 4. Reactor trip occurs on steam generator low-low level. No credit is taken for immediate release of the control rod drive mechanisms caused by a loss of offsite power.
- 5. The worst single failure occurs in the auxiliary feedwater system.
- 6. Auxiliary feedwater is delivered to four steam generators.
- 7. Secondary system steam relief is achieved through the steam generator safety valves.
The assumptions used in the analysis are essentially identical to the loss of. normal feedwater flow incident (Section 15.2.7) except that power is assumed to be lost to the reactor coolant pumps at the time of reactor trip. Plant characteristics and initial conditions are further discussed in Section 15.0.3. t 0965T:1/6 15.2-14
l Results The transient response of the RCS following a loss of AC power is shown in Figures 15.2.6-1 and 15.2.6-2. The calculated sequence of events for this event is listed in Table 15.2-1. The first few seconds after the loss of power to the reactor coolant pumps will closely resemble a simulation of the complete loss of flow incident (see Section 15.3.2) 1.e., core damage due to rapidly increasing core temperatures is prevented by promptly tripping the reactor. After the reactor trip, stored and residual decay heat must be removed to prevent damage to either the RCS or the core. The LOFTRAN code results show that the natural circulation flow available is sufficient to provide adequate core decay heat removal following reactor trip and RCP coastdown. . 15.2.6.3 Conclusions Analysis of the natural circulation capability of the Reactor Coolant System has demonstrated that sufficient heat removal capability exists following RCP coastdown to prevent fuel or clad damage. 15.2.7 LOSS OF NORMAL FEEDWATER FLOW 15.2.7.1 Identification of Causes and Accident Description , A loss of normal feedwater (from pump failures, valve malfunctions, or loss of offsite AC power) results in a reduction in capability of the secondary system to remove the heat generated in the reactor core. If an alternate supply of feedwater were not supplied to the plant, core residual heat following reactor trip would heat the primary system water to the point where water relief from the pressurizer would occur, resulting in a substantial loss of water from the Reactor Coolant System (RCS), Since the plant is tripped well before the steam generator heat transfer capability is reduced, the primary system variables nev%r approach a DNB condition. 0965T:1/6 15.2-15
i i l The following occur upon loss of normal feedwater (assuming main feedwater pump failures or valve malfunctions):
- 1. As the steam system pressure rises following the trip, the steam generator power-operated relief valves are automatically opened to the atmosphere.
Steam dump to the condenser is assumed not to be available. If steam flow through the power relief valves is not available, the steam generator safety valves may lift to dissipate the sensible heat of the fuel and coolant plus the residual decay heat produced in the reactor.
- 2. As the nc load temperature is approached, the steam generator power-operated relief valves (or safety valves, if the power operated relief valves are not available) are used to dissipate the residual decay heat and to maintain the plant at the hot shutdown condition.
A loss of normal feedwater is classified as an ANS Condition II event, fault of moderate frequency. See Section 15.0.1 for a discussion of Condition II events. The Auxiliary Feedwater System is started automatically as discussed in Section 15.2.6.1. The steam oriven auxiliary feedwater pump utilizes steam from the seconcary system and exhausts to the atmosphere. The motor driven auxiliary feedwater pumps are supplied by power from the diesel generators. The pumps take suction directly from the auxiliary feedwater storage tank for delivery to the steam generators. An analysis of the system transient is presented below to show that following a loss of normal feedwater, the Auxiliary Feedwater System is capable of removing the stored and residual heat, thus preventing either overpressuriza-tion of the RCS or loss of water from the reactor core, and returning the plant to a safe condition. L C 0965T:1/6 15.2-16 b T
l 15.2.7.2 Analysis of Effects and Consecuences
- Method of Analysis A detailed analysis using the LOFTRAN Code (Reference 2) is performed in order to obtain the plant transient following a loss of normal feedwater. The simulation describes the plant thermal kinetics, RCS, pressurizer, steam generators and feedwater system. The digital program computes pertinent variables including the steam generator level, pressurizer water level, and reactor coolant average temperature.
Assumptions made in the analysis are:
- 1. The plant is initially operating at 102 percent of the Engineered Safety Features design rating.
- 2. A conservative core residual heat generation based upon long term operation at the initial power level preceding the trip. '
- 3. Reactor trip occurs on steam generator low-low level.
- 4. The worst single failure occurs in the auxiliary feedwater system.
- 5. Auxiliary feedwater is delivered to four steam generators.
- 6. Secondary systern steam relief is achieved through the steam generator
, safety valves, i The loss of nomal feedwater analysis is performed to demonstrate the adequacy of the reactor protection and engineered safeguards systems (e.g., the Auxiliary Feedwater System) in removing long term decay heat and preventing
- excessive heatup of the RCS with possible resultant RCS overpressurization or loss of RCS water.
l 0965T:1/6 15.2-17
As such, the assumptions used in this analysis are designed to minimize the energy removal capability of the system and to maximize the possibility of water relief from the coolant system by maximizing the coolant system expansion, as noted in the assumptions listed above. For the loss of normal feedwater transient, the reactor coolant volumetric flow remains at its normal value and the reactor trips via the low-low steam generator level trip. Thc reactor coolant pumps may be manually tripped at some later time to reduce heat addition to the RCS. An additional assumptiot made for the loss of normal feedwater evaluation is that only the pressurizer safety valves are assumed to function normally. Operation of the valves maintains peak RCS pressure at or below the actuation setpoint (2500 psia) throughout the transient. Plant characteristics and initial conditions are further discussed in Section 15.0.3. Normal reactor control systems are not required to function. The keactor Protection System is required to function following a loss of normal feedwater as analyzed here. The Auxiliary Feedwater System is required to deliver a minimum auxiliary feedwater flowrate. No single active failure will prevent operation of any system required to function. Results Figures 15.2.7-1 and 15.2.7-2 show the significant plant parameters following a loss of normal feedwater. Following the reactor and turbine trip from full load, the water level in the steam generators will fall due to the reduction of steam generator void fraction and because steam flow through the safety valves continues to dissipate the stored and generated heat. One minute following the initiation of the low-low level trip, at least two auxiliary feedwater pumps are l automatically started, reducing the rate of water level decrease. 0965T:1/6 15.2-18
The capacity of the auxiliary feedwater pumps are such that the water level in the steam generators being fed does not recede below the lowest level at which sufficient heat transfer area is available to dissipate core residual heat ( without water relief from the RCS safety valves. Figure 15.2.7-1 shows that at no time is there water relief from the pressurizer. The calculated sequence of events for this accident is listed in Table 15.2-1. As shown in Figures 15.2.7-1 and 15.2.7-2, the plant approaches a stabilized condition following reactor trip and auxiliary feedwater initiation. Plant procedures may be followed to further cool dcwn the plant. 15.2.7.3 Conclusions ~ Results of the analysis show that a loss of normal feedwater does not adversely affect the core, the RCS, or the steam system since the auxiliary feedwater capacity is such that reactor coolant water is not relieved from the pressurizer relief or safety valves. 15.2.8 FEEDWATER SYSTEM PIPE BREAK l 15.2.8.1 Identification of Causes and Accident Descriotion A major feedwater line rupture is defined as.a break in a feedwater line large enough to prevent the addition of sufficient feedwater to the steam generators to maintain shell side fluid inventory in the steam generators. If the break is postulated in a feedline between the check valve and the steam generator, fluid from the steam generator may also be discharged through the break. Further, a break in this location could preclude the subsequent addition of auxiliary feedwater to the affected steam generator. (A break upstream of the feedline check valve would affect the Nuclear Steam Supply System only as a loss of feedwater. This case is covered by the evaluation in Section 15.2.7). Depending upon the size of the break and the plant operating conditions at the time of the break, the break could cause either a RCS cooldown (by excessive energy discharge through the break) or a RCS heatup. Potential RCS cooldown 0965T:1/6 15.2-19
l l l l resulting from a secondary pipe rupture is evaluated in Section 15.1.4. Therefore, only the RCS heatup effects are evaluated for a feedweter line rupture. A feedwater line rupture reduces the the ability to remove heat generated by l the core from the RCS for the following reasons: 1
- 1. Feedwater flow to the steam generators is reduced. Since feedwater is sub' cooled, its loss may cause reactor coolant temperatures to increase prior to reactor trip.
- 2. Fluid in the steam generator may be discharged through the break, and would then not be available for decay heat removal after trip.
- 3. The break may be large enough to prevent the addition of any main feedwater after trip.
An Auxiliary Feedwater System is provided to assure that adequate feedwater . will be available such that:
- 1. No substantial overpressurization of the RCS shall occur.
- 2. Sufficient liquid in the RCS shall be maintained in order to provide adequate decay heat removal.
A major feedwater line rupture is classified as an ANS Condition IV event. See Section 15.0-1 for a discussion of Condition IV events. The severity of the feedwater line rupture transient depends on a number of system parameters including break size, initial reactor power, and credit taken for the functioning of various control and safety systems. A number of cases of feedwater line break have been analyzed. Based on these analy:es, it has been shown that the most limiting feedwater line ruptures are the double ended rupture of the largest feedwater line, occurring at full power with and without loss of offsite power, with no credit taken for pressurizer control. These cases are analyzed below. 0965T:1/6 15.2-20
The following provides the necessary protection 4 r a main feedwater rupture:
- 1. A reactor trip on any of the following conditions:
- a. High pressurizer pressure.
- b. Overtemperature t.T.
- c. Low-low steam generator water level in any steam generator,
- d. Safety injection signals from any of the following:
- 1) 2/3 low steam line pressure in any loop.
- 2) 2/3 high containment pressure
~
- 2. An Auxiliary Feedwater System to provide an assured source of feedwater to the steam generators for decay heat removal.
15.2.8.2 Analysis of Effects and Consecuences 4 Method of Analysis A detailed analysis using the LOFTRAN Code (Reference 2) is performed in order to determine the plant transient following a feedwater line rupture. The code describes the plant thermal kinetics,. RCS including natural circulation, pressurizer, steam generators and feedwater system, and computes pertinent variables including the pressurizer pressure, pressurizer water level, and reactor coolant average temperature. The cases analyzed assume a double-ended rupture of the largest feedwater pipe at full power. Major assumptions made in the analyses are as follows:
- 1. The plant is initially operating at 102 percent of engineered safeguards power.
L 0965T:1/6 ~15.2-21
- 2. Initial reactor coolant average temperature is 4.0*F above the nominal value, and the initial pressurizer pressure is 30 psi above its nominal value.
- 3. No credit is taken for the pressurizer power-operated relief valves or I pressurizer spray. -
- 4. Initial pressurizer level is at the nominal programmed value plus 5 percent (error); initial steam generator water level is at the nominal value plus 5 percent in the faulted steam generator and at the nominal value minus 5 percent in the intact steam generators.
- 5. No credit is taken for the high pressurizer pressure reactor trip.
- 6. Main feedwater flow to all steam generators is assumed to be lost at the time the break occurs (all main feedwater spills out through the break).
- 7. The worst possible break area is assumed. This maximizes the blowdown discharge rate following the time of trip, which maximizes the resultant heatup of the reactor coolant.
- 8. A conservative feedline break discharge quality is assumed prior to the time the reactor trip cecurs, thereby maximizing the time the trip setpoint is reached. After the trip occurs, a saturated liquid discharge is assumed until all the water inventory is discharged from the affected steam generator. This minimizes the heat removal capability of the affected steam generator.
- 9. Reactor trip is assumed to be initiated when the low-low level trip setpoint minus 10 percent of the narrow range span in the faulted steam generator is reached.
- 10. The Auxiliary Feedwater System is actuated by the low-low steam generator 4 water level signal. The Auxiliary Feedwater System is assumed to supply a total of 450 gallons per minute (gpm) to the three unaffected steam 0965T:1/6 15.2-22
generators, including allowance for possible spillage through the main feedwater line break. A 60 second delay was assumed following the low-low level signal to allow time for startup of the emergency diesel generators and the auxiliary feedwater pumps. An additional 132 seconds was assumed before the feedwater lines were purged and the relatively cold (110 F) auxiliary feedwater entered the unaffected steam generators.
- 11. No credit is taken for heat energy deposited in RCS metal during the RCS heatup.
- 12. No credit is taken for charging or letdown.
- 13. Steam generator heat transfer area is assumed to decrease as the shell side liquid inventory decreases.
- 14. Conservative core residual heat generation is assumed based upon long term operation !t the initial power lev'el preceding the trip.
- 15. No credit is taken for the following potential protection logic signals to mitigate the consequences of the accident:
- a. High pressurizer pressure.
- b. Overtemperature AT.
- c. High pressurizer level.
- d. High Containment pressure.
Receipt of a low-low steam generator water level signal in at least one steam generator starts the motor driven auxiliary feedwater pumps, which then C deliver auxiliary feedwater flow to the steam generators. The turbine driven auxiliary feedwater pump is initiated if the low-low steam generator level signal is reached in at least two steam generators. Similarly, receipt of a low steam line pressure signal in at least one steam line initiates a steam line isolation signal which closes the main steam line isolation valves in all steam lines. This signal also gives a safety injection signal which initiates flow of borated water into the RCS. The amount of safety injection flow is a function of RCS pressure. 0965T:1/6 15.2-23
l t l l i l Plant characteristics and initial conditions are further discussed in Section .. 15.0.3. No reactor control systems are assumed to function. The Reactor Protection System is required to function following a feedwater line rupture as analyzed here. No single active failure will prevent operation of this system. The engineered safsty systems assumed to function are the Auxiliary Feedwater
~
System and the Safety Injection System. For the Auxiliary Feedwater System, the worst case configuration has been used, i.e., three intact steam generators receive auxiliary feedwater following the break. The turbine driven auxiliary feedwater pump has been assumed to fail. The two motor driven pumps together-deliver 450 gpm to the three intact steam generators allowing for spillage out of the break. Only one train of safety injection has been assumed to be available. Following the trip of the reactor coolant ~ pumps for the feedline rupture without offsite power, there will be a flow-coastdown until flow in the loops reaches the natural circulation value. The natural circulation capability of the RCS has been shown in Section 15.2.6, for the loss of AC power transient, to be sufficient to remove core decay heat following reactor trip. Pump coastdown characteristics are demonstrated in Sections 15.3.1 and 15.3.2 for single and multiple reactor coolant pump trips, respectively. Results Calculated plant parameters following a major feedwater line rupture are shown in Figures 15.2.8-1 through 15.2.8-8. Results for the case with offsite power available are presented in Figures 15.2.8-1 through 15.2.8-4. Results for the case where offsite power is lost are presented in Figures 15.2.8-5 through-15.2.8-8. The calculated sequence of events for both cases analyzed are . listed in Table 15.2-1. The system response following the feedwater line rupture is similar for both L cases analyzed. Results presented in Figures 15.2.8-2 and 15.2.8-4 (with offsite power available) and Figures 15.2.8-6 and 15.2.8-8 (without offsite 0965T:1/6 15.2 power) show that pressures in the RCS and main steam system remair, below 110 j percent of the respective design pressures. Pressurizer pressure increases until reactor trip occurs on low-low steam generator water level. Pressure then decreases, due to the loss of heat input, until the Safety Infection System is actuated on low steam line pressure in the ruptured loop. Coolant expansion occurs due to reduced heat transfer capability in the steam generators; the pressurizer safety valves open to maintain primary pressure at an acceptable value. Addition of the safety injection flow aids in cooling ( down the primary and helps to ensure that sufficient fluid exists to keep the core covered with water. Figures 15.2.8-1 and 15.2.8-5 show that following reactor trip, the plant remains subcritical. RCS pressure will be maintained at the safety valve setpoint until safety injection flow is terminated by the operator. 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 Auxiliary Feedwater System and makeup is provided by the Safety Injection System. The major difference between the two cases analyzed can be seen in the plots of hot and cold leg temperatures, Figure 15.2.8-3 (with offsite power available) and Figure 15.2.8-7 (without offsite power). It is apparent from the initial portion of the transient (~200 seconds), that the case without offsite power results in higher temperatures in the hot leg. For longer times, however, the case with offsite power results in a more severe rise in temperature until the coolant pumps are turned off and the Auxiliary Feedwater System is realigned. The pressurizer fills for the case with power due to the increased coolant expansion resulting from the pump heat addition; hence, water is relieved for the case with power. As previously stated, however, the core remains covered with water for both cases. L 0965T:1/6 15.2-25
15.2.a.3 Conclusions - a Results of the analyses show that for the postulated feedwater line rupture, the assumed Auxiliary Feedwater System capacity is adequate to res.ove decay heat, to prevent overpressurizing the RCS, and to prevent uncovering the A reactor core. 15.
2.9 REFERENCES
} 1. Mangan, M. A., " Overpressure Protection for Westinghouse Pressurized Water Reactors," WCAP-7769, Rev. 1 June, 1972.
- 2. Burnett, T. W. T., et al. , "LOFTRAN Code Description," WCAP-7907, June i 1972.
l s 4 I l 0965T:1/6' - 15.2-26 l I
. m
.. a - - re , -
TABLE 15.2-1 (Page 1) < l TIME SE00ENCE OF EVENTS FOR INCIDENTS WHICH CAUSE A DECREASE IN HEAT REMOVAL BY THE SECONDAitY SYSTEM Accident Event Time (see) Turbine Trip
- 1. With pressurizer Turbine trip, 0.0 control (minimum loss of main feedwater reactivity feedback) flow t
High pressurizer pressure 7.3 reactor trip point reached , Initiation of steam 7.5 release from steam generator safety valves Rods begin to drop 9.3 Peak pressurizer pres- '11.0 sure occurs Minimum DNBR occurs 11.5 I l l k 0965T:1/6- 15.2-27
TABLE 15.2-1 (Page 2) TIME SEQUENCE OF EVENTS FOR INCIDENTS WHICH CAUSE A DECREASE IN HEAT REMOVAL BY THE SECONDARY SYSTEM Accident Event Time (sec)
- 2. With pressurizer Turbine trip, 0.0 control (maximum reac- loss of main feedwater tivity feedback) flow Peak pressurizer 7.0 pressure occurs Initiation of steam 7.5 releasefrom steam generttor safety valves Low-low steam generator 62.9 water level reactor trip point reached Rods begin to drop 64.9 Minimum DNBR occurs (1)
(1) DNBR does not decrease below its initial value. i 0965T:1/6 15.2-28 { 1 l
TABLE 15.2-1 (Page 3)
' TIME-SEQUENCE OF EVENTS FOR INCIDENTS WHICH CAUSE A DECREASE-IN HEAT REMOVAL BY THE SECONDARY SYSTEM Accident , Event Time (sec)
( 3. Without pressurizer control (minimum reactivity feedback) Turbine trip, loss of main feedwater flow 0.0 i High pressurizer pressure 3.7 reactor trip point reached Rodi begin to drop 5.7 Initiation of steam 7.0 ! release from steam
- generator safety valves Peak pressurizer pressure 7.5 t
occurs
- Mini:num DNBR occurs (1)
C (1) DNBR does not decrease below its initial value.
.l 0965T:1/6 i
15.2-29
-l
. . - , #~~
i f
- l l TABLE 15.2-1 (Page 4) i f TIME SE00ENCE OF EVENTS FOR INCIDENTS WHICH CAUSE A l
DECREASE IN HEAT REMOVAL BY THE SECONDARY SYSTEM i i Accident Event Time (sec)
- 4. Wit'hout pressurizer Turbine trip, 0.0 control (maximum loss of main feedwater
, reactivity feedback) flow High pressurizer pressure 3.7 reactor trip point reached Rodsbesintodrop 5.7 Initiation of steam 7.0 release from steam generator safety valves Peak pressurizer pressure 7.0 occurs-Minimum DNBR occurs- (1) (1) DNBR does not decrease below its initial'value. t 0965T:1/6 . 15.2-30 v
TABLE 15.2-1 (Page 5) TIME SEQUENCE OF EVENTS FOR INCIDENTS WHICH CAUSE A DECREASE IN HEAT REMOVAL BY THE SECONDARY SYSTEM Accident Event Time (sec) l ( Loss of Non-Emergency AC power Main feedwater flow stops 10 l Low-low steam generator 63 water level trip Rods begin to drop 65 Reactor coolant pumps 65 begin to coastdown Peak water level in 67 pressurizer occurs i Four steam generators 123 begin to receive auxiliary feedwater from - two motor driven auxiliary feedwater pumps Core decay heat decreases -500 to auxiliary feedwater heat removal capacity . L 0965T:1/6 15.2-31
. _ - -= . . TABLE 15.2-1 (Page 6) TIME SEQUENCE OF EVENTS FOR INCIDEllTS WHICH CAUSE A DECREASE IN HEAT REMOVAL BY THE SECONDARY SYSTEM Accident Event Time (see) Loss of Normal Feedwater Main feedwater flow 10 Flow stops Low-low steam generator 63 water level trip Rods begin to drop 65 Peak water level in 67 pressurizer occurs Four steam generators 123 begin to receive
~
auxiliary feedwater from two motor driven auxiliary feedwater pumps Core decay heat decreases - 1000 to auxiliary feedwater heat removal capacity t C 0965T:1/6 15.2 , ._- -
l l TABLE 15.2-1 (Page 7) 1 _ TIME SEQUENCE OF EVENTS FOR INCIDENTS WHICH CAUSE A DECREASE IN HEAT REMOVAL BY THE SECONDARY SYSTEM Accident Event Time (sec) Feedwater system pipe break
- 1. With offsite power Main feedline rupture occurs 10 available Low-low steam generator level 27 reactor trip setpoint reached ,
in rtJptured steam generator Rods begin to drop 29 Auxiliary feedwater is 87 delivered to intact steam generators Low steam line pressure. 384 setpoint reached in ruptured steam generator L L 0965T:1/6 15.2-33
h TABLE 15.E-1 (Page 8) TIME SE00ENCE CF EVENTS FOR INCIDENTS WHICH CAUSE A DECREASE IN HEAT REMOVAL BY THE SECONDARY SYSTEM Accident Everft Time (ser1 All main steam line 391 [ isolation valves close \ Steam generator safety 844 valve setpoint reached in intact steam g.er.erators Pressurizer water relief
.388 begins Core decay r.aat.plus 3000 pump heat decreases to auxiliary feecwatte hest removal capacity j
l 0 g .. l 0965T:1/6 *
.15.2 "
p %2 $ f Y s
%_ r J
. _ - s
TABLE 15.2-1 (Page 9) TIME SEOUENCE OF EVENTS FOR INCIDENTS WHICH CAUSE A 1 DECREASE IN HEAT REMOVAL BY THE SECONDARY SYSTEM Accident Event Time (sec)
- 2. Without offsite power Main feedline rupture 10 C.
occurs Low-low steam generator -27 level reactor trip setpoint reached in ruptured steam generator Rods ~ begin to drop, power 29 lost to the reactor coolant pumps Auxiliary.feedwater is 87 delivered to intact steam generators Low steam line pressure 451 setpoint reached in ruptured steam generator 4 C - 0965T:1/6 15.2-35
l l TABLE 15.2-1 (Page 10) TIME SE0VENCE OF EVENTS FOR INCIDENTS WHICH CAUSE A DECREASE IN HEAT REMOVAL BY THE SECONDARY SYSTEM Accident Event Time (see)- All main steam line 458 isolation valves close Steam generator safety 1108 valve setpoint reached in intact steam generators Core decay heat decreases - 2000 to auxiliary feedwater - heat removal capacity 4 i i 4 P 4 i i-g
- 0965T:1/6- 15.2 36
]
1.2 3 i
.E .8 l.
EE a- l
.g C 25
.. .t U ,5 4 -- -
0 2600 i l i
, 2400-- --
E g . m [; 2200- - 2g 3 .. ( 2 2000- - - 1800 1600 '
.. . l I
g 1400- - --
,=- ..
( g- 1200- 1, j E$ 1
- l 1000- - ..
5.,3 ..
- 800--- : ;
0 20 40 60 80 100 TIME (SECCN03) l FIGURE 15.2.3-1 ( Turnine Tno Event Witn Pressurizer Scray ( and Power Operatec Relief Vaives. Minimum Reactmty Feoobacx 15.2-37
600
~
=
- 590l .
gg 580 '
= l w<
85 570 - - 1
" 1E O
560 - -- 550 . . . w 640 E 5 -
!t T =-
gE 600 .- . 2 i W 8 w 560 ' 5.0 ' ' 4.0 . . .. l = 3.0 - - -- l E , 2.0 . 1.0 0 20 40 60 80 100 TIE (SECONDS) FIGURE 15.2.3-2 Turbine Trio Event With Pressurizer Scray and Power Operated Relief Valves. M.nimum Reactwity Feechack 15.2-38
l 1.2 , C -
=5
$! .8 - -
C =5 W= aw E .4 " 0 2600 l' 3 2400-l - l x - l
= ;; 2200- . -
I s *- .. I 5 - C 5 2000- - - " 1800 1500 I .. n I
=3 1300 -
T
-2 2 o
@j 1100 - - -
l h
- h 900
- : : .
0 20 40 60 80 100 TIME (SEccleS) FIGURE 15.2.3-3 Turbine Trio Event With Pressurtzer Scray and Power Operated Relief Vatves. With Maximum Rescuvity Feeoback ! 15.2-3'9 \ _ . - . - . - - - ---
600 E j - l
= !
i e 580 - T NT
- l [
2 ( 5
- 560 - ~
8 v l 540' 640 4 E g w k
=
F 600 - [ -- 8 v 560 , , 5.0 4.5 * -- 4.0- - a 3.5-- , f 3 3,0 . --
/ ..
- 2. 5 ~ --
, 2.0- - ~
)
1.5 : : . 0 20 14 0 60 80 100 TIME (SEC305) L FIGURE 15.2.3-4 Turbine Trio Event Wim Pressurizer Sorav l and Power Operated Relief Valves. Wim Maximum Reactivity Feecbr.cx 15.2-40 w -
l ( 1.2 i O .8 1
- i
.==
C gg a.
=B
@g .4- --
l W: su \ g -. 0 I 2600 ' 2500 -- -.I. E R l U E7 2400-ll 2300 - l i
- .E 2200 - .
E a 2100- - 3 C 2000- - 1 l l 1900< : : ; . ! l l g 1500 5 1400 - - --
"@ 1300 - ..
a-
"u 1200 -
$$ 1100h ..
$ 1000 - -
C = 900
- 800l :
l 0 % Ib 20 60 80-TIME (SECOMOS) FIGURE 15.2.3-5 Tureine Trio Event Without Pressurizer SurJy or Power Operated Relief Valves. Minimum Reactivity Feooback 15.2-41
. l 600 h -
( l 5 580 '
=
= l h --
1 5 560 - --
.=.
n! -- 540 640 l E 620 - e a
~
h
- c 600 - - -
w 3 580 - .- w 8 o 560 . 5.0 l 4.0 - - E 3.0 - - 2.0 , 0 20 W 60 80 100 TI E (SECOND$) FIGURE 15.2.3-6 Turcine Trio Event Wimout Pressurizer Sorcy or Power Operated Relief Valves, Wim Minirnum Reactwity Fescbacx 15.2-42
- c , .-
l 1.2 l u' -
- 1 E l
= .8
( a:I5= ma t8 he - SE .".I .4*
~
~
~
J { o 2600 2500 - l E
@ 2400-l ..
=
67 2300 - .. 53 0 P.:. as 2200 - - l 5
= 2100 - - ..
2000 - 1900- . : .' E 1,400 s=; .. OU j] 12pn - -- 2g -.
$ 1000 -
= 7 C * --
.L 800 l . .' .' .'
0 20 4 60 80 100 TIME (SECONOS) FIGURE 15.2.3-7 [ TurDine Trio Event Wimout Pressurizer Soray ' ( and Power Operated Relief Valves. Wim Manimum Reactivity Faseback 15.2-43
4 3 580 - 7
), _
W W 560 - - g .. 540 g 640 2 _f. _ _ = 5 600 - - -
=
5
- m ..
5 560 . . : . 5.0 . 4.0 - . 3 .. E 3.0 - - -- 2.0 , TIME (SECONO3) FIGURE 15.2.3-8~ C Tureine Trio Event Witncut Pressurizer Scray - and Power Operated Relief Valves. With Maxirnum Reec:wity Feecback 15.2-44
( 2600 , 2500 - - ( E 7 E 2400 - 2300 - 5 y 2200 - - - ( 2 5
=
2100 - -
- N
$ 2000 - - ~~
1900 - -
- 1800 " " ' '" '"
1800 ' '" '" 1700 - - -- 1600 - - -- j 1500 - - -- s [p 1400 - - -- Wd-
"o 1300 - - --
25 - 28 g- 1200 -- -- N E 1100 - - -- 1000 - - -- 900 - - -- l 800 " ' ' 100 2 5 108 2 5 102 2 5 103 2 5 10* TIME (SEC:*03) Figure 15.2.6-1. Pressuruer Pressure and Water Volume Transiona for Loss of Offsite Power 15.2-45 e-- - -
-- r
t 700 650
- MOT LIG a[ 600 ,
w N --
,E E
~
550 .', cou LIs ( A 8 500 . 450 ,,, , i 1500 " i 1250 - - "" 4 i i W j R c 1000 - - '" e + 3 i 35= i 2t W 750 - - 3 i 1
=
500 - - . 100 2 5 foi 2 5 102 2 5 103 2 5 10" TIME (SEColtOS) t 1 Figure 15.2.5-2. i Core Average Tempermaire Transent and Steam Generator Pressare for Loss of Offsite Power 15.2-46
- a. --w
2500 . . i 1 2500 *
- 7 2400- --
E 2300 - -- m 0 E 2200- - --
~_
E C N
=
2100- - --
=
2000- -- 1900- -- 1800 ' ' 1800 .' * - " 1700- - .. 1600- - .- f 1500- - --
. gE 1400- - --
32 a
=2 1307 - --
%"8 g* 1200- - ..
=
2 1100- - .- 1000- - -- l 900- - - . 800 ' ' ' ' * ' * * ' ' " - ' 3 10 2 5 10' 2 5 102 2 5 10 3 2 5 10' TIME (SECOMOS) Figure 15.2.71. Pressurizer Pressere and Water Volume Transients for Loss of Normal Feoowater 15.2-47
l - 700 . .. C 650 .. _ Not uG
= .00 _
u, h cot.o us "
\
5 550 g .. a. 8 500 . 450 , , , , , , , , ,,, b T-- 1500 7 6> 1250 -- .. w R 0 . m 1000 -- - 2 W W -- - 750 3 --
=
500 -- 250 l00 2 5 10' 2 5 102 2 5 103 2 5 IOS TlHE (SECONOS) Figure 15.2.7 2. Looo Temperatures and Steam Generator Pressars for Loss of Normal Feedwater 15.2-48
I l l 1.2 _ l.0 - ( =- 1 w ;; .8 - 83 3* .6 - 25 M:
*W .4 -
e T (
~
.2 -
l 0
.01 i:: 0.0 E
g4 .01 - 35 . - y' .02 - - ( 2
.03 .
1.0 E3 .75 - d* gl
.5 -
3 B' .25 .. W5 0.0 SW # d"~ .25 - M C .5 -
.75 "
-1 0 I 2 3 4 10 10 10 10 10 TIME (SEC303)
FIGURE 15.2.8-1 1 Nuclear Power Transient. Total Core Reactrvity Transient. and Feedline Break Flow Transient for Main Feedline Ruoture Wim Offsite Power Availacle 15.2-49 :I1 l 1
2600 E I l i E 2400 * - u
$ 2200 " "
.a.
@ 2000 + -
R 0 1800 - - 2 1600 2000 1750 - g - E 1500 " 53 0= 1250 - -
u 5l 1000 -
f I~ g 750 - - E 500 - - 250
$f E 30 u_ -
b_$m dW 20 -
" as SE h$ 10 - - - -
Y 0 1
-10 0 I 2 3 4 10 10 10 10 10 714 (SECON05)
FIGURE 15.2.8-2 Pressurize Pressure. Water Volume, and Relief Transients for Main Feed!irw Ruoture Wim Offsite Power Available 15.2-50 - l
I i
~
l 700
= 650 - -
5 M0T LEG 0-EE 1! W 600 - - - WB C h ", CQJ LEG
-5 o- = 550 - --
d 500 ( E 5 650 - - i h7 M07 LIG 5* G w=
<E
- 600 - - --
E
$h cots LEG =-
y 550 - - -- 100 2 5 10' 2 5 102 2 5 10 3 2 5 10' TIME (SECONOS) L FIGURE 15.2.8-3 Reactor Coolant Temeerature Transients for the Faulted and ne intact Lcces for Main Feedline Ructure With Offsate Power Availacie 15.2-51
I
- l
, 1500 ' i 1250 - - ~ 14 TA CT ( STEAM GENERATcn3 1000 . ! U t i = e 750 - .. i 51 , w
- E 8
500 - -
= '
em. ' 250 .. 5 , F A ut.TED 2 STEAw GEMEAATCR m n 0 ' - - . i . 1.2 t i 1.0 . 1 7
, L'
- m2
=g .8. .
0% w w .5_ 4 8 v0 .6- . te. f .
- k I .4 .
i .2 - - .. 0 100 2 5 10' 2 5 102 2 5 '103 .2 5 10' TIE (SECONO3) FIGURE 15.2.8-4 i e Steam Generater Pressure and Core Heat Flux i Transients for Main Feedline Ruoture With Offsite Power Availacle 15.2-52
e 1.2 i 1.0 - ( .E m-EE
.8 * -
1 t [g .6 - SE [5= .4 T - ( t .2 -
~
0
.01 0.0 C
_E .01 -
,0{ ~
m _
=< .02 - -
a-2 C 2 .03 -
.04 1.0 EW g5 .75 -
5 .50 "
&3 .
W5 .25 -
~-
a-
@j 0.0 s i
C --
~
.25 L
.50 -
.75 100 2 5 10' 2 5 102 2 5 103 2 5 10*
TIME (SECONOS) FIGURE 15.2.8-5 l Nuclear Pcwer Transient. Total Core Reactivir/ Transient, and Feecline Break Flow Transient for Main Feedline Ruoture Without Offsite Power Availacle 15.2-53
l i 1 l 4 l 2600 l 1 ;
- E
- 2400 - b I 4-w 5
m 0 2200 - - . i i i 4 2 2000 - E 4 N
$ 1800 -
1600 1000 W S i s* 1250
- jj 1000 "
a 55 750 - 1 28 E- 500 " g u ( 6 250 1 i i
- _ 30 ..
I m3 W
$"3 20
-r M-l j$ 10 " -
U2 - EE 0 ' 3
-10 1 10 0 I 2 3 4 10 10 10 10 l TIM! (SECONCS)
L FIG'JRE 15.2.8-6 Pressurizer Pressure, Water Volume, and Relief Rate for Main Feedline Ructure Without Offsite Power Available i 15.2-54
700 " 650 - - g M07 LIG C = E E., 600 -- N P
%5 COLD LEG
( a. as 8C S~ C 7 550 -- E 500 ' 700 . . ... . . . . . . . . l W
=
3 ( 8 "h 8~ Uw
**E 650 - -
NOT L(G "G 600 - - '
*- O --
8k w% COLD LIG M . 5 550 - - -- 500 ' , 108 2 5 108 2 5 102 2 5 103- 2 5 10 4 TIME (SECOMOS) FIGURE 15.2.8-7 i Reactor Caolant Temperature Transients for the Faultec and intact Locos for Main Feecline Ruoture Witncut Offsite Power Availacle l 15.2-55
1500 ' 1250 - - Im7AC7 STEAM GENERATCR5 2 w g 750 . .. O E E 500 - - - E W 3 5 250 - Faui.:n stEAw -- y GENERA'fCR 0 - . 1.2 . ... .. . . .. 5 1.0 - - 1 - I s - g .8 - - c u E
.6- - -
E d O ' W .4- - -- w 8
.2- -- .
O 108 .' 5 108 2 5 If2 2 5 103 2 5 10 4 TlHE(SECONDS) j FIGURE 15.2.8-8 Steam Generator Pressure and Core Heat Flux
' Transients for Main Feecline Ruoture Without Offsite Power Available 15.2-56
( 15.3 DECREASE IN REACTOR COOLANT SYSTEM FLOW RATE A number of faults are postulated which could result in a decrease in reactor coolant system flow rate. These events are discussed in this section. Detailed analyses are presented for the most limiting of these events. Discussions of the following flow decrease events are presented in Section 15.3: ( 1. 2. Partial Loss of Forced Reactor Coolant Flow Complete Loss of Forced Reactor Coolant Flow
- 3. Reactor Coolant Pump Shaft Seizure (Locked Rotor)-
Item 1 above is considered to be an ANS Condition II event, item 2 an ANS Condition III event, and item 3 an ANS Condition IV event. Section 15.0.1 contains a discussion of ANS classifications. 15.3.1 PARTIAL LOSS OF FORCED REACTOR COOLANT FLOW 15.3.1.1 Identification of Causes and Accident Description A partial loss of coolant flow accident can result from a mechanical or elec-trical failure in a reactor coolant pump, or from a fault in the power supply to the pump or pumps supplied by a reactor coolant pump bus. If the reactor is at power at the time of the accident, the immediate effect of loss of cool-ant flow is a rapid increase in the coolant temperature. This increase could result in DNB with subsequent fuel damage if the reactor is not tripped promptly. , Normal power for the pumps is supplied through individual buses connected to the generator and the offsite power system. When a generator trip cccurs,_- the buses continue to be supplied from external power lines, and the pumps con-tinue to supply coolant to the core. This event is classified as an ANS Condition II incident (an incident of mod-erate frequency) as defined in Section 15.0.1. L 0947T:6 15.3-1
The necessary protection against a partial loss of coolant flow accident is provided by the low primary coolant flow reactor trip signal which is actuated in any reactor coolant loop by two out of three low flow signals. Above Per-missive 8, low flow in any loop will actuate a reactor trip. Between approxi-mately 10 percent power (Permissive 7) and the power level corresponding to Permissive 8, low flow in any two loops will actuate a reactor trip. Above 1 Permissive 7, two or more reactor coolant pump circuit breakers opening will actuate the corresponding undervoltage relays. This results in a reactor trip which serves as a backup to the low flow trip. 15.3.1.2 Analysis of Effects and Consecuences Method of Analysis The loss of one reactor coolant pump with four loops in operation has been analyzed. This transient is analyzed by three digital computer codes. First, the LOFTRAN Code (Reference 1) is used to calculate the loop and core flow during the transient, the time of reactor trip based on the calculated flows, the nuclear power transient, and the prirrary system pressure and temperature tran-sients. The FACTRAN Code (Reference 2) is then used to. calculate the heat flux transient based on the nuclear power and flow from LOFTRAN. Finally, the 1HINC Code (see FSAR Section 4.4) is used to calculate the DNBR during the transient based on the heat flux from FACTRAN and flow from LOFTRAN. The DNBR transients presented represent the minimum of the typical or thimble cell. This accident is analyzed with the Improved Thermal Design Procedure as described in WCAP-8567. Plant characteristics and initial conditions are discussed in Section 15.0.3. Initial Ccnditions Initial reactor power, pressure, and RCS temperature are assumed to be at their nominal values. Uncertainties in initial conditions are included in the limit DNBA as described in WCAP-8567. 0947T:6 15.3-2
1 1 Reactivity Coefficients ! A conservatively large absolute value of the Doppler-only power coefficient is ( used (see Figure 15.0.4-1). This is equivalent to a total integrated Doppler reactivity from 0 to 100 percent power of 0.016 6k. A positive moderator temperature coefficient of +5 pcm/*F is assumed. This results in the maximum core power during the initial part of the transient when the minimum DNBR is reached. Flow Coastdown Normally, the flow coastdown analysis is based on a momentum balance around each reactor coolant loop and across the reactor core. This momentum balance is combined with the continuity equation, a pump momentum balance and the pump characteristics and is based on high estimates of system pressure losses.
~
However, during the course of the McGuire Unit I startup testing a flow coastdown more severe than that presented in the FSAR was measured. The measured coastdown was input to the LOFTRAN code for this analysis. Results Figures 15.3.1-1 through 15.3.1-4 :;how the transient response for the loss of one reactor coolant pump with four loops in operation. Figure 15.3.1-4 shows the DNBR to be always greater than the limit value. For both cases analyzed, 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 temperatures do not increase significantly above their respective initial values. The calculated sequence of events _ tables for the case analyzed are shown' on
~
~ Table 15.3-1. The affected reactor coolant pump will continue to coast down, and the core flow will reach a new equilibrium value corresponding to the number of pumps still in operation. With the reactor tripped, a stable plant condition will eventually be attained. Normal-' plant shutdown may then proceed, 0947T:6 15.3-3
15.3.1.3 Conclusions The analysis shows that the DNBR will not decrease below the limit value at any time during the transient. Thus, no fuel or clad damage is predicted, and all applicable acceptance criteria are met. 15.3.2 COMPLETE LOSS OF FORCED REACTOR COOLANT FLOW 15.3.2.1 Identification of Causes and Accident Descrietion A complete loss of forced reactor coolant flow may -result from a simultaneous loss of electrical supplies to all reactor coolant pumps. If the reactor is at power at 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 riot tripped promptly. , Normal power for the reactor coolant pumps is supplied through buses from a ' transformer connected to the generatur and the offsite power system. Each pump is on a separate bus. . When a gv -rator trip occurs, the buses continue to be supplied from external power lines and the pumps continue to supply coolant flow to.the core. . This event is classified as an ANS Condition III incident (an infrequent-inci-dent) as defined in Section 15.0.1. The following signals provide the necessary protection against a complete loss of flow. accident:
- 1. Reactor coolant pump power supply undervoltage or underfrequency.
- 2. Low reactor coolant loop flow.
The reactor trip'on reactor coolant pump undervoluge is provided to protect against conditions which can cause a loss of voltage to all reactor coolant _ pumps, i.e., station blackout. This function is blocked below approximately-l 10~ percent power (Permissive 7). e 0947T:6- 15.3 .
The reactor trip on reactor coolant pump underfrequency is provided to trip the reactor for an underfrequency condition, resulting from frequency distur-bances on the power grid. Reference 3 provides analyses of grid frequency disturbances and the resulting Nuclear Steam Supply System protection require-ments which are generally applicable. The reactor trip on low primary coolant loop flow is provided to protect against loss of flow conditions which affect only one reactor coolant loop. This function is generated by two out of three low flow signals per reactor coolant loop. Above Permissive 8, low flow in any loop will actuate a reactor trip. Between approximately 10 percent power (Permissive 7) and the power level corresponding to Permissive 8, low flow in any two loops will actuate a reactor trip. 15.3.2.2 Analysis of Effects and Consecuences The complete loss of flow transient has been analyzed for a loss of all four reactor coolant pumps with four loops in operation. This transient is analyzed by three digital computer codes. First. the LOFTRAN Code (Reference 1) is used to calculate the loop and core flow during the transient, the time of reactor trip based on the calculated flows, the nuclear power transient, and the primary system pressure and temperature tran-sients. The FACTRAN Code (Reference 2) is then used to calculate the heat flux transient based on the nuclear power and flow from LOFTRAN. Finally, the THINC Code is used to calculate the DNBR during the transient based an the heat flux from FACTRAN and flow from LOFTRAN. The DNBR transients presented represent the minimum of the typical or thimble cell. The method of analysis and the assumptions made regarding initial operating conditions and reactivity coefficients are identical to those discussed in Section 15.3.1, except that following the loss of power supply to all pumps at power, a reactor trip is actuated by either reactor coolant pump power supply ; undervoltage or underfrequency.
. l 0947T:6 15.3-5
I Results Figures 15.3.I-1 tnrough 15.3.2-4 show the transient response for the loss of power to all reactor coolant pues with four loops in operation. The reactor is assumed te be tripped on undervoltage signal. Figure 15.3.2-4 shows the i DNBR to be always greater than the limit value. 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 temperatures do not increase significantly above their respective initial values. The calculated sequence of events for the case analyzed is shown on Table 15.3-1. The reactor coolant pumps will continue to coast down, and natural circulation flow will eventually be established, as demonstrated in Section , 15.2.6. With the reactor tripped, a stable. plant condition would be attained. Normal plant shutdown may then procsed. 15.3.2.3 Conclusior.s The analysis performed has demonstrated that for the complete loss of forced reactor coolant. flow, the DNBR does not decrease below the limit value at any time during the transient. Thus, no fuel or clad damage is predicted, and all applicable acceptance criteria are met. 15.3.3 REACTOR COOLANT PUMP SHAFT SEIZURE (LOCKED ROTOR) 15.3.3.1 Identification of Causes and Accident Description The accident postulated is an instantaneous seizure of a reactor coolant pump , rotor. Flow through the affected reactor coolant loop is rapidly reduced, leading to an initiation of a reactor trip on a low flow signal. Following initiation of the reactor trip, heat stored in the fuel rods con-tinues to be transferred to the coolant causing the coolant to' expand.' At the same time, heat transfer to the shell side of the- steam generatdrs is reduced, _- 0947T:6 15.3-6
d 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 plant trip). The rapid expansion of the coolant in the reactor core, combined C with reduced heat transfer in the steam generators causes an insurge into the pressurizer and a pressure increase throughout the Reactor Coolant System. The insurge inte the pressurizer compresses the steam volume, actuates the auto-matic spray system, opens the power-operated relief valves, and opens the pressurizer safety valves, in that sequence. The three power-operated relief valves are designed for reliable operation and would be expected to function properly during th'e accident. However, for conservatism, their pressure reduc-ing effect as well as the pressure-reducing effect of the spray is not included in the analysis. This event is classified as an ANS Condition IV incident (a limiting fault) as defined in Section 15.0.1. . 15.3.3.2 Analysis of Effects and Consecuences Method of Analysis l Two digital-computer codes are used to analyze this transient. The LOFTRAN Coce (Reference 1) is used to calculate the resulting loop and core flow tran-sients following the pump seizure, the time of reactor trip based on the loop flow transients, the nuclear power following reactor trip, and to determine the peak pressure. The thermal behavior of the fuel located at the core hot spot is investigated using the FACTRAN Code (Reference 2), which uses the core flow and the nuclear power calculated by LOFTRAN. The FACTRAN Code includes a film boiling heat transfer coefficient. At the beginning of the postulated locked rotor accident, i.e., at the time the shaft in one of the reactor coolant pumps is assumed to seize, the plant is assumed to be in operation under the most adverse steady state operating condition, i.e., maximum guaranteed steady state thermal power, maximum steady _ state pressure, and maximum steady state coolant average temperature. Plant characteristics and initial conditions are further discussed in Section 15.0.3. 0947T:6 15.3-7
For the peak pressure evaluation, the initial pressure is conservatively esti-mated as 30 psi above nominal pressure (2250 psia) to allow for errors in the pressurizer pressure measurement and control channels. This is done to obtain the highest possible rise in the coolant pressure during the transient. The pressure response shown in Figure 15.3.3-2 is at the point in the Reactor Coolant System having the maximum pressure. ( Evaluation of the Pressure Transient After pump seizure, the neutron flux is rapidly reduced by control rod inser-tien. Rod motion is assumed to begin one second after the flow in the affec-ted loop reached 87 percent of nominal flow. No credit is taken for the pres-sure reducing effect of the pressurizer relief valves, pressurizer spray, steam dump or controlled feedwater flow after plant trip. Although these operations are expected to occur and would result in a lower peak pressure, an additional degree of conservatism is provided by ignoring their effect. The pressurizer safety valves are full open at 2575 psia and their capacity for steam relief is as described in FSAR Section 5.5. Evaluation of DNB in the Core During the Accident For this accident, DNB is assumed to occur in the core, and therefore, an evaluation of the consequences with respect to fuel rod thermal transients is performed. Results obtained from analysis of this " hot spot" condition repre-sent the upper limit with respect to clad temperature and zirconium water reaction. In the evaluation, the rod power at the hot spot is assumed to be 2.5 times the average rod power (i.e., Fg = 2.5) at the initial core power level. Film Boiling Coefficient The film boiling coefficient is calculated in the FACTRAN Code using the Bishop-Sandberg-Tong film boiling correlation. The fluid properties are evalu-0947T:6 15.3-8
- - _ _ _ _ = _ _ _ _ _ . _ _ _ _ _ .
ated at film temperature (average between wall and bulk temperatures). The program calculates the film ceefficient at every time step taased upon the actual heat transfer conditions at the time. The neutren flux, system presa sure, bulk density, and mass flow rate as a function of time are used as program input. For this analysis, the initial values of the pressure and the bulk density are used throughout the transient since they are the most conservative with respect to clad temperature response. For conservatism, DNB was assumed to start at the beginning of the accident. Fuel Clad Gap Coefficient The magnitude and time dependence of the heat transfer coefficient between fuel and clad (gap coefficient) has a pronounced influence on the thermal results. The larger the value of the gap. coefficient, the more heat is trans-ferred between pellet and clad. Based on investigations on the effect of the i ( gap coefficient upon the maximum clad temperature during the transient, the gap coefficient was assumed to increase from a steady state value consistent with initial fuel temperature to 10,000 BTU /hr-ft 2 *F at the initiation of the transient. Thus, the large amount of energy stored in the fuel because of the small initial value is released to the clad at the initiation of the tran-sient. Zirconium Steam Reaction The zirconium-steam reaction can become significant above 1800*F (clad tem-perature). The Baker-Just parabolic rate. equation shown below is used to define the rate of the zirconium steam reaction. 2 6 d(w )
- 33.3 x 10 -45,500 dt **E[1.986T]
where: w = amount reacted, mg/cm 2 0947T:6 15.3-9
t = time, sec ' T = temperature, 'F The reaction heat is 1510 cal /gm. The effect of zirconium-steam reaction is included in the calculation of the
" hot spot" t. lad temperature transient. 'Results The transient results are shown in Figures 15.3.3-1 through 15.3.3-4. The results of these calculations are also summarized in Table 15.3-2. The peak Reactor Coolant System pressure reached during the transient is less than that which would cause stresses to exceed the faulted condition stress limits.
i Also, the peak clad surface temperature is considerably less'than 2700*F. It should be noted that the clad temperature was conservatively calculated assum- . ing that DNB occurs at the initiation of the transient. 4 15.3.3.3 Conclusions
- 1. Since the peak Reactor Coolant System pressure reached during the tran-sient is less than that which would cause stresses to exceed the faulted condition stress limits, the integrity of the primary coolant system is
; not endangered.
4
- 2. Since the peak clad surface temperature calculated for the hot spot -
during the worst transient remains considerably less than 2700*F the core will remain in place and intact with no loss of core cooling capability. 15.
3.4 REFERENCES
- 1. Burnett, T. W. T., et al., "LOFTRAN Code Description", WCAp-7907, June 1972.
0947T:6 15.3-10~- , " m , .s, ,4a + e . p +
1
- 2. Hargrove, H. G., "FACTRAN - A Fortran - IV Code for Thermal Transients in a U02 Fuel Red", WCAP-7908, June 1972.
l ( 3. Baldwin, M. S., Merrian, M. M. , Schenkel, H. S. and Van De Walle, D. J.,
"An Evaluation of Loss of Flow Accidents Caused by Power System Frequency Transients in Westinghouse PWRs," WCAP-8424, Revision 1, June 1975.
1 4 i [
)i 4 ; i-4 1
0947T:6 15.3-11
TABLE 15.3-1 (Page 1) T1ME SEQUENCE OF EVENTS FOR INCIDENTS W
,HICH RESULT IN A DECREASE IN REACTOR COOLANT SYSTEM FLOW i
Accident Partial Loss cf Forced Event Time (sec.)' ( Reactor Coolant Flow Coastdown begins 0. Low flow reactor trip 1.04 Rods begin,to drop 2.04 Minimum DNBR occurs 3.40 Complete Loss of Forced Reactor Coolant Flow
. . All operating . -0.
pumps lose power and begin coast-ing down Reactor coolant O. pump undervoltage j trip point reached [ Rods begin to
-drop- 1.5 .
Minimum DNBR occurs 3.4, l - l 0947T:6 15.3 ;- ?
< ~ -
. . . . . - _ . = . - - .
i t C TABLE 15.3-1 (Page 2) l TIME SEOUENCE OF EVENTS FOR INCIDENTS WHICH RESULT IN A DECREASE IN REACTOR COOLANT SYSTEM FLOW l Accident Event Time (sec.) Reactor Coolant Pump
! Shaft Seizure (Locked Rotor) 4, I
Rotor on one pump locks 0.
~
Low flow trip , ,. point reached 0.06 Rods begin to drop 1.06
, Maximum RCS pressure occurs 3.3 I
Maximum clad temperature occurs 3.5 C L 10947T:6- 15.3-13
t l l l 3 TABLE 15.3-2
SUMMARY
OF RESULTS FOR LOCKED ROTOR TRANSIENTS
~
Maximum Reactor Coolant System Pressure (psia) 2593 Maximum Clad Temperature (*F) Core Hot Spot 1943
~
) Zr-H2 O reaction at core ] hot spot (percent by weight) 0.5-4 l } s t 4 1
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0 2 4 6 8 10 TIME (SECONOS) FIGURE 15.3.1-1 I
. Flow Transients for Four Loops in Operation.
One Pump Coasting Down 15.3-15 1
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e 2 2100 -. 2000 -- - -- 1900 0 2 4 6 8 10 TIME (SECONOS) FIGURE 15.3.1-2 Nuclear Power and Pressurizer Pressure Transients for Four Loops in Operation. One Pump Coastmq Down 15.3-16 I
i . 1.2 1.0 ( -
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0 0 2 4 6 8 10 l TlHE(SECONOS) FIGURE 15.3.1-3 Average and Hot Channel Heat Flux Transients for Four Loops in Operation. One Pump Coasting Down 15.3-17
I l C 2.4 - 2.2 . . . 2.0 .
/
s 1. 8 - - E .. 1.6 .
- 1. 4 - -
1.2 . . . O I 2 3 4 5 TIME (SECOMOS) b FIGURE 15.3.1-4 DNBR vs. Time for Four Loops in Operation. One Pump Coasting Down 15.3-18
( 1.2 -' 1.0 - - 4 E
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W O .2- . 0 ' O 2 4 6 8 10 TIME (SECOMOS) b . FIGURE 15.3.2-1 Core Flow Coastdo :n for Four Loops in Operation, Four Pumps Coasting Down L 15.3-19
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( 2000 -- .. 1900 0 2 4 6 8 10 TIME (SECONOS) FIGURE 15.3.2-2' . Nue! ear Power and Pressurizer Pressure . Tranments for Four Loops in Operation. Four Pumps Coasting Down 5.3-20
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0 2 4 6 8 10 TIME (SECONDS) FIGURE 15.3.2-3 Average and Hot Channel Heat Flux Transients for Four Loops in Operation. Four Pumps Coasting Down 15.3-21
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TI E (SECOMOS) i FIGURE 15.3.2-4 DNBR vs. Time for Four Looos in Operation. Four Pumps Coasting Down 15.3-22 4
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1.2 ' O 1 2 3 4 5 TIME (SECONOS) FIGURE 15.3.2-4 DNBR vs. Time for Four Loops in Operation, Four Pumps Coasting Down 15.3-22 4 e
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0 2 4 6 8 10 TIME (SECONDS) FIGURE 15.3.3-1 L Flow T'ransients for Four Loops in Operation,- One Locned Rotor 15.3-23
C ( 2700 2600 - - -. 2500 - - -- 2 IC [ 2400- -
.E /
p 2300 - - T U 2200- . -. 2100- - -. 2000 - - 0 2 4 6 8 10 TIME (SECON03) FIGURE 15.3.3-2 Peak Reactor Coolant Pressure for Four Loops in Operation, One Locked Rotor 15.3-24 e
.. .. .. -._ _ - - _ - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - -- - o
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= _
5 .8- - AVERAGE enANaCL w8 0g - g ;- .6- - W 5 .4 - -
.2 '
0 2 4 6 8 10 TIME (SECONDS) 1 FIGURE 15.3.3-3 Nuclear Power. Average And Hotbhannel Heat Flux Transients for Four Loops - 15.3-25 in Operation. One Locked Rotor
l l 3000
= - 2500 - -
i I Y E 2000 - O as - W f 1500 - - -- e 3 v 1000 - - -- 500 0 2 4 6- 8 10 TIME (SECONOS) l FIGURE 15.3.3-4 Maximum Clad Temperature at Hot Soot for i
/
Four Looos in Operation, One Locked Rotor ( I 15.3-26 t
15.4 REACTIVITY AND POWER DISTRIBUTION ANOMALIES A number of faults have been postulated which could result in reactivity and power distribution anomalies. Reactivity changes could be caused by control C rod motion or ejection, boron concentration changes, or addition of cold water to the Reactor Coolant System. Power distributton changes could be caused by control rod motion, misalignment, or ejection, or by static means such as fuel assembly mislocation. These events are discussed in this section. Detailed analyses are presented for the most limiting of these events. Discussions of the following incidents are presented in Section 15.4:
- 1. Uncontrolled Rod Cluster Control Assembly Bank Withdrawal from a Suberitical or Low Power Startup Condition
- 2. Uncontrolled Rod Cluster Control A.ssembly Bank Withdrawal at Power
( 3. 4. Rod Cluster Control Assembly Misoperation Startup of an Inactive Reactor Coolant Pump at an Incorrect Temperature
- 5. Spectrum.of Rod Cluster Control Assembly Ejection Accidents Items 1, 2, 3, and 4 are considered to be ANS Condition II events, and Item 5 an ANS Condition IV event. Item 3 entails both Condition II and III events.
Section 15.0.1 contains a discussion of ANS classifications. 15.4.1 UNCONTROLLED ROD CLUSTER CONTROL ASSEMBLY BANK WITHDRAWAL FROM l A SUBCRITICAL OR LOW POWER STARTUP CONDITION i 15.4.1.1 Identification of Causes and Accident Description A rod cluster control assembly (RCCA) withdrawal accident is defined as an uncontrolled addition of reactivity to the reactor core caused by withdrawal of RCCA's resulting in a power excursion. Such a transient could be caused by. l 0352T:6 15.4-1 i
l i a malfunction of the reactor control or rod control systems. This could occur with the reactor either subcritical, hot zero power or at power. The "at power" case is discussed in Section 15.4.2. Although the reactor is normally brought to power from a suberitical condition by means of RCCA withdrawal, initial startup procedures with a clean core call for boron dilution. The maximum rate of reactivity increase in the case of boron dilution is less than that assumed in this analysis. The RCCA drive mechanisms are wired into preselected bank configurations which are not altered'during reactor life. These circuits prevent the RCCA's from being automatically withdrawn in other than their respective banks. Power supplied to the banks is controlled such that no more than two banks can be withdrawn at the same time and in their proper withdrawal sequence. The RCCA drive :aechanisms are of the magnetic latch type and coil actuation is sequenced to provide variable speed trave.l. The maximum reactivity insertion rate analyzed in the detailed plant analysis is that occurring with the simultaneous withdrawal of the combination of two sequential control banks ' having the maximum combined worth at maximum speed. This event is classified as an ANS Condition II incident (an incident of muderate frequency) as defined in Section 15.0.1. The neutron flux response to a continuous reactivity insertion is charac-terized by a very fast rise terminated by the reactivity feedback effect of the negative Doppler coefficient. This self limitation of the power excursion of primary importance since it limits the power to a tolerable level during the delay time for protective action. Should a continuous RCCA withdrawal accident occur, the transient will be terminated by the following automatic features of the Reactor Protection System:
- 1. Source Range High Neutron Flux Reactor Trip - actuated when either of two independent source range channels indicates a neutron flux level above a preselected manually adjustable setpoint. This trip function may be manually bypassed only after an intermediate range flux channel indicates 0952i:6 15.4-2
a flux level above a specified level. It is automatically reinstated when both intermediate range channels indicate a flux level below a specified level. ( 2. Intermediate Ranoe Hioh Neutron Flux Reactor Trip - actuated when either of two independent intermediate range channels indicates a flux level above a preselected manually adjustable setpoint. This trip function may be manually bypassed only after two of tne four power range channels are ( reading above approximately 10 percent of full p u er and is automatically reinstated when three of the four channels indicate a power level below this value. ,
- 3. power Range High Neutron Flux Reactor Trip (Low settina) - actuated when two out of the four power range channels indicate a power level above approximately 25 percent of full power. This trip function maybe manually bypassed when two of the four power range channels indicate a power level above approximately 10 percent of full power and is automatically reinstated only after three of the four channels indicate a power level below this value.
- 4. Power Range High Neutron Flux Reactor Trip (High Settino) - actuates when two out of the four power range channels indicate a power level above a preset setpoint. This trip function is always active.
- 5. High Nuclear Flux Rate Reactor Trip - actuated when the positive rate of change of neutron flux on two out of four nuclear power range channels indicate a rate above the preset setpoint. This trip function is always active. .
In addition, control rod stops on high intermediate range flux level (one of two) and high power range flux level (one out of four) serve to discontinue rod withdrawal and prevent the need to actuate the intermediate range flux level trip and the power range flux level trip, respectively. L 09527:6 15.4-3
15.4.1.2 Analysis of Effects and Consecuences Method of Analysis The analysis of the Uncontrolled RCCA bank withdrawal from subcritical accident is performed in three stages: first an average core nuclear power transient calculation, then an average core heat transfer calculation, and finally the DNBR calculation. The average core nuclear calculation is performed using spatial neutron kinetics methods (TWINKLE (Reference 1) to determine the average power generation with time including the various total core feedback effects, i.e., Doppler reactivity and moderator reactivity. The average heat flux and temperature transients are determined by performing a fuel rod transient heat transfer calculation in FACTRAN (Reference 2). The average heat flux is next used in THINC for transient DNBR calculation. This accident is analyzed using the Improved Thermal Design Procedure as described in WCAP-8567. Plant characteristics and initial conditions are discussed in Section 15.0.3. In order to give conservative results for a startup accident, the following assumptions are made:
- 1. Since the magnitude of the power peak reached during the initial part of the transient for any given rate of reactivity insertion is strongly f dependent on the Doppler coefficient, conservatively low values as a function of power are used. See Section 15.0.4 and Table 15.0.3-2.
, 2. Contribution of the moderator reactivity coefficient is negligible during the initial part of the transient because the heat transfer time between the fuel and the moderator is much longer than the neutron flux response time. However, after the initial neutron flux peak, the succeeding rate of power increase is affected by the moderator reactivity coefficient. A highly conservative value is used_ in the analysis to yield the maximum peak heat flux.
- 3. The reactor is assumed to be at hot zero power. This assumption is more conservative than tnat of_a lower initial system temperature. The higher j 0952T:6 15.4-4
l i initial system temperature yields a large fuel-water heat transfer coefficient, larger specific heats, and a less negative (smaller absolute magnitude) Doppler coefficient, all of which tend to reduce the Doppler feedback effect thereby increasing the neutron flux peak. The initial effective multiplication factor is assumed to be 1.0 since this results in the worst nuclear power transient.
- 4. Reactor trip is assumed to t,e initiated by power range high neutron flux (low setting). The most adverse combination of instrument and setpoint errors, as well as delays for trip signal actuation and rod cluster control assembly release, is taken into account. A 10 percent increase is
~
assumed for the power range flux trip setpoint raising it from the nominal value of 25 percent to 35 percent. Since.the rise in the neutron flux is so rapid, the effect of errors in the trip setpoint on the actual time at which the rods are released is negligible. In addition, the reactor trip insertion characteristic is based on the assumption that the highest worth rod cluster control assembly is stuck in its fully withdrawn position. ( See Section 15.0.5 for rod cluster control assembly insertion characteristics.
- 5. The maximum positive reactivity insertion rate assumed is greater than '
that for the simultaneous withdrawal of the combination of the two sequential control banks having the greatest combined woh.h at maximum speed (45 inches / minute).
- 6. The most limiting axial and radial power shapes, associated with having the two highest combined worth banks in thtir high worth position, is assumed in the DNB analysis.
l
- 7. The initial power level was assumed to be below the power level expected for any shutdown condition (10 -9 of nominal power). This combination of highest reactivity insertion rate and -lowest initial power produces the highest peak heat flux.
- 8. Two reactor coolant pumps are assumed to be in operation. This is conservative with respect to DNB.
( 0952T:6 15.4-5
Results Figures 15.4.1-1 through 15.4.1-3 show the transient behavior for the ! uncontrolled RCCA bank withdrawal incident, with the accident terminated by reactor trip at 35 percent of nominal power. The reactivity insertion rate 1 used is greater than that calculated for the two highest worth sequential contro,1 banks, both assumed to be in their highest incremental worth region. Figure 15.4.1-1 shows the neutron flux transient.
~
The energy release and the fuel temperature increases are relatively small. The thermal flux response, of interest for DNB considerations, is shown on Figure 15.4.1-2. The beneficial effect of the inherent thermal lag in the fuel is evidenced by a peak heat flux much less than the full power nominal value. There is a large margir. 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 15.4.4-3 shows the response of the average fuel and cladding temperature. The average fuel temperature increases to a value lower tnan the nominal full power value. The minimum DNBR at all times remains above the limit value. The calculated sequence of events for this accident is shown on Table 15.4-1. With the reactor tripped, the plant returns to a stable condition. The plant may subsequently be cooled down further by following normal plant shutdown proceaures. 15.4.1.3 Conclusions In the event of a RCCA withdrawal accident from the subtritical condition, the core and the Reactor Coolant System are not adversely affected, since the combination of thermal power and the coolant temperature result in a DNBR greater than the limit value. Thus, no fuel or clad damage is predicted as a result of DNB. C 0952T:6 15.4-6
i l 15.4.2 UNCONTROLLED R00 CLUSTER CONTROL A EMBLY BANK WITHDRAWAL AT POWER 15.4.2.1 Identification of Causes and Accident Descriotion ( Uncontrolled rod cluster control assembly (RCCA) bank withdrawal at power results in an increase in the core heat flux. Since the heat extraction from , the steam generator lags behind the core power generation until the steam generator pressure reaches the relief or safety valve setpoint, there is a net increase in the reactor coolant temperature. Unless terminated by manual or automatic action, the power mismatch and resultant coolant temperature rise could eventually result in DNB. Therefore, in order to avert damage to the fuel clad the Reactor Protection System is designed to terminate any such transient before the DNBR falls below the limit value. This event is classified as an ANS Condition II incident (an incident of mode ~ rate frequency) as defined in Section 15.0.1. i ( The automatic features of the Reactor Protection System which prevent core damage following the postulated accident include the following:
- 1. Power range neutron flux instrumentation actuates a reactor trip if two-of-four channels exceed an overpower n' 'oint.
- 2. Reactor trip is actuated if any two-out-of-four AT channels exceed an overtemperature AT setpoint. This setpoint is automatically varied with axial power imbalance, coolant temperature and pressure to protect against DNB. -
- 3. Reactor trip is actuated if any two-out-of-four AT channels exceed an overpower AT setpoint. This setpoint is automatically varied with axial power imbalance to ensure that the allowable heat generation rate (kw/ft) is not exceeded.
, 4 A high pres:urizer pressure reactor trip actuated from any two-out-of-four
- ,ressure channels which is set at a fixed point. This set pressure is less than the set pressure for the pressurizer safety valves.
03b21:6 15.4-7 I
e
- 5. A high pressurizer water level reactor trip actuated from any two-out-of-
[ three level channels when the reactor power is above approximately 10 per-cent (Permissive 7). In addition to the above listed reactor trips, there are the following RCCA withdrawal blocks:
- 1. High neutrcn flux (one-out-of-four power range)
- 2. Overpower AT (two-out-of-four)
- 3. Overtemperature AT (two-out-of-four)
The manner in which the combination of overpower and overtemperatare AT trips provide protection over the full range of Reactor Coolant System conditions is described in FSAR Chapter 7,. Figure 15.0.3-1 presents allowable reactor coolant loop average temperature and AT for the design power distribution and flow as a function of primary coolant. pressure. The boundaries of operation defined by the overpower AT trip and the overtemperature AT trip are represented as " protection lines" on 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 a given DNBR can be represented as a line. The DNB lines represent the locus of conditions for which the DNBR equals the limit value (1.47 for the thimble cell and 1.49 for the typical cell). All points below and to the left of a DNB line for a given pressure have a DNBR greater than the limit value. The diagram shows that DNB is prevented for all cases if the area enclosed with the maximum protection lines is not traversed by the applicable DNBR line at any' point. The area of permissible operation (power, pressure and temperature) is bounded by the combination of reactor trios: high neutron flux (fixed setpoint); high pressure (fixed setpoint); icw pressure (fixed setpoint); overpower and over-temperature AT (variable setpoints). 0952T:6 15.4-8
15.4.2.2 Analysis of Effects and Consecuences Method of Analysis The transient is analyzed by the LOFTRAN Code (Reference 3). This code simulates the neutron kinetics, Reactor Coolant System, pressurizer, pressurizer relief and safety valves, pressurizer spray, steam generator, and steam generator safety valves. The code computes pertinent plant variables including temperature, pressures, and power level. The core limits as illustrated in Figure 15.0.3-1 are used as input to LOFTRAN to determine the , minimum DNBR during the transient.
- This accident is analyzed with the Improved Thermal Design Procedure as described in WCAP-8567. Plant characteristics and initial conditions are discussed in Section 15.0.3.
- 1. Initial reactor power, pressur~e, and RCS temperatures are assumed to be at i
( their nominal values. Uncertainties in initial conditions are included in the limit DNBR as described in WCAP-8567.
- 2. Reactivity Coefficients - Two cases are analyzed:
- a. Minimum Reactivity Feedback. A least negative moderator coefficient of reactivity is assumed corresponding to the beginning of core life.
A variable Doppler power coefficient with core power is used in the l analysis. A conservatively small (in absolute magnitude) value is assumed.
- b. Maximum Reactivity Feedback. A conservatively large positive moderator density coefficient and a large (in absolute magnitucs?
negative Doppler power coefficient are assumed. L 0952T:6 15.4-9 _ _,___.-_m-_- . _ _ _ - - - - - -
1 l
- 3. The reactor trip on high neutron flux is assumed to be actuated at a C conservative value of 118 percent of nominal full power. The AT trips include all adverse instrumentation and setpoint errors; the delays for trip actuation are assumed to be the maximum values
- 4. The RCCA trip insertion characteristic is based on the assumption that the highest worth assembly is stuck in its fully withdrawn position.
- 5. The maximum positive reactivity insertion rate is greater than that for the simultaneous withdrawal of the two control banks having the maximum combined worth at maximum speed. 2 The effect of RCCA movement on the axial core power distribution is accounted for by causing a decrease in overtemperature AT trip'setpoint proportional to the decrease in margin to DNE.
No single active failure in any plant system will adversely affect the consequences of the accident. Results Figures 15.4.2-1 through 15.4.2-3 show the transient response for a rapid RCCA withdrawal incident starting from full power. Reactor trip on high neutron flux occurs shortly after the start of the accident. Since this is rapid with respect to the thermal time constants of the plant, small changes in T,yg and pressure result and margin to DNB is maintained. The transient response for a slow RCCA withdrawal from full power is shown in Figures 15.4.2-4 through 15.4.2-6. Reactor trip on overtemperature AT occurs after a longer period and the rise in temperature and pressure is consequently larger than for rapid RCCA withdrawal. Again, the minimum DNBR is greater than the limit value. Figure 15.4.2-7 shows the minimum DNBR as a function of reactivity insertion rate from initial' full power operation for minimum and maximum reactivity l l 0952T:6 15.4-10 i
feedcack. It can be seen that two reactor trip channels provide- protection over the whole range of reactivity insertion rates. These are the high neutron flux and overtemperature AT channels. The minimum DNBR is never ( less than the limit value. Figures 15.4.2-8 and 15.4.2-9 shows the minimum DNBR as a function of reactivity inse-tion rate for RCCA withdrawal incidents starting at 60 and 10 percent power, respectively. The results are similar to the 100 percent power case, except as the initial power is decreased, the range over which the overtemperature AT trip is effective is increased. In neither case does the DNBR fall below the limit value. The shape of the curves of minimum DNB ratio versus reactivity insertion rate in the reference figures is due both to reactor core and coolant system transient response and to protection system action in initiating a reactor trip. ( Referring to Figure 15.4.2-8, for example, it is noted that:
- 1. For high reactivity insertion rates (i.e., above - 2 x 10"# oK/sec) reactor trip is initiated by the high neutron flux trip for the minimum reactivity feedback cases. The neutron flux Jevel in the core rises rapidly for these insertion rates while core heat flux and coolant system temperature lag behind due to the thermal caoacity 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 DNB ratios during the transient. As reactivity insertion rate cecreases, core heat flux and coolant temperatures can remain more nearly in equilibrium with the neutron flux; minimum DNB ratio 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 Reactor Coolant System average temperature and pressure. It is important to note that the average temperature contribution to the circuit is C
0952T:6 15.4-11
lead-lag compensated in order to decrease the effect of the thermal capacity of the Reactor Coolant System in response to power increases.
- 3. With further decrease in reactivity insertion rate, the overtemperature AT and high neutron flux trips become equally effective in terminating the transient (e.g. , at - 2 x 10 -4 6K/sec reactivity insertion rate).
For reactivity insertion rates between - 2 x 10 -4 6K/sec and - 6 x 10 6K/see the effectiveness of the overtemperature AT trip increases (in terms of increased minimum DNB ratio) due to the fact that with lower insertion rates the power increase rate is slower, the rate of rise of average coolant temperature is slower and the system lags and delan become less significant.
- 4. For reactivity insertion rates less than - 6 x 10-5 6K/sec, the rise in the reactor coolant temperature is sufficiently high so that the steam generator safety valve setpoint is reached prior to trip. Opening of these valves, which act as an additional heat load on the Reactor Coolant System, sharply decreases the rate of increase of Reactor Coolant System average temperature. This decrease in rate of increase of the average coolant system temperature during the transient is accentuated by the lead-lag compensation causing the overtemperature AT trip setpoint to be reached later with resulting lower minimum DNB ratios.
Figures 15.4.2-7, 15.4.2-8, and 15.4.2-9 illustrate minimum DNBR calculated for minimum and maximum reactivity feedback. Since the RCCA withdrawal at power incident is an overpower transient, the fuel temperatures rise during the transient until after reactor trip occurs. For high reactivity insertion rates, the overpower transient is fast with respect to the fuel rod thermal time constant, and the core hsat flux lags behind the neutron flux response. Due to this lag, the peak core heat flux does not exceed 118 percent of its nominal value (i.e.', the high neutron flux trip setpoint assumed in the analysis). Taking -into account the effect of the RCCA withcrawal en the axial core power distribution, the peak fuel centerline temperature will still remain below the fuel mel' ting temperature. 0952T:6 15.4-12
For slow reactivity insertion rates, the core heat flux remains more nearly in equilibrium with the neutron flux. The overpower transient is terminated by the overtemperature AT reactor trip before a DNB condition is reached. The peak heat flux again is maintained below 118 percent of its nominal value. C Taking into account the effect of the RCCA withdrawal on the axial core power distribution, the peak fuel centerlir.e temperature will remain below the fuel melting temperature. Since DNB dot.s not occur at any time during the RCCA withdrawal at power transient, the ability of the primary coolant to remove heat from the fuel rod is not reduced. Thus, the fuel cladding temperature does not rise significantly above its initial value during the transient. The calculated sequence of events for this accident is shown on Table 15.4-1. With the reactor tripped, the plant eventually returns to a stable condition. The plant rr.ay subsequently be cooled down further by following normal plant shutdown procedures. 15.4.2.3 Conclusions The high neutron flux and overtemperature AT trip channels provide adequate protection over the entire range of possible reactivity insertion rates, i.e., the minimum value of DNBR is always larger than the limit value. 15.4.3 ROD CLUSTER CONTROL ASSEMBLY MISOPERATION 15.4.3.1 Identification of Causes and Accident Descriotion Rod cluster control assembly misoperation accidents include: : A. One or more dropped RCCAs within the same group ~ B. A dropped RCCA bank C. Statically misaligned RCCA 0952T:6 15.4-13
D. Withdrawal of a single RCCA. Each RCCA has a 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. Full length RCCAs are always moved in preselected banks, and the b2riks are , always moved in the same preselected sequence. Each bank of RCCAs is divided into two groups. The rods comprising a group operate in parallel through multiplexing thyristors. The two groups in a bank move sequentially such that the first group is always within one step of the second group in the bank. A definite schedule of actuation (or deactuatien of the stationary gripper, movable gripper, and lift coils of a mechanism) is required to withdraw the RCCA attached to the mechanism. Since the stationary gripper, movable gripper, and lift coils associated with the four RCCAs of a rod group are driven in parallel, any single failure which would cause rod withdrawal would affect a minimum of one group. Mechanical failures are in the direction of insertion, or immobility. i The dropped RCCA, dropped RCCA bank, and statically misaligned RCCA events are classified as ANS Condition II incidents (incidents of moderate frequency) as' defined in subsection 15.0.1. However, the single RCCA withdrawal incident is classified as an ANS Condition III event, as discussed below. No single electrical or mechanical failure in the rod control system could cause the accidental withdrawal of a single RCCA from the inserted bank at full power operation. The operator could withdraw a single RCCA in the con-trol bank since this feature is necessary in order to retrieve an assembly should one be accidentally dropped. The event analyzed must result from multiple wiring failures or multiple serious operator errors,' and subsequent and repeated operator disregard of. event indication. The probability of such a combination of conditions is so low that the limiting consequences may include slight fuel damage. 0952T:6' 15.4-14
Thus, consistent with the philosophy and format of ANSI N18.2, the event is classified as a Condition III event. By definition, " Condition III occurrences include incidents, any one of which may occur during the lifetime ( of a particular plant", and "shall not cause more than a sn.all fraction of fuel elements in the reactor to be damaged ...". This selection of criterion is in accordance with General Design Criterion (GDC) 25 which states, "The protection system shall be designed to assure that specified acceptable fuel design limits are not exceeded for any single mal-function of the reactivity control systems, such as accidental withdrawal (not ejection or dropout) of control rods." (Emphases have been added). It has been shown that single failures resulting in RCCA bank withdrawals do not violate specified fuel design limits. Moreover, no single malfunction can result in the withdrawal of a single RCCA. Thus, it is concluded that criterion established for the single rod withdrawal at power is appropriate and in accordance with GDC 25. ,, ( A dropped RCCA or RCCA bank is detected by: A. Sudden drop in the core power level as seen by the nuclear instrumen-tation system B. Asymmetric power distribution as seen on out-of-core neutron detectors or core exit thermocouples C. Rod at bottom signal ( D. Rod deviation alarm E. Rod position indication. Misaligned RCCAs are detected by: A. Asymmetric power distribution as seen on out-of-core neutron detectors or core exit thermocouples L 0952T:6 15.4-15
l B. Rod deviation alarm C. Rod position indicators The resolution of the rod position indicator channel is +5 percent of span {+7.2 inches). Deviation of any RCCA from its group by twice this distance (10 percent of span, or 14.4 inches) will not cause power distributions worse than the design limits. The deviation alarm alerts the operator to rod devia-tion with respect to the group position in' excess of 5 percent of span. If the rod deviation alarm is not operable, the operator is required to take action as required by the Technical Specifications. If one or more rod position indicator channels should be out of service, detailed operating instructions shall be followed to assure the alignment of the non-indicated RCCA. The operator is also required to take action as required by the Technical Specifications.. In the extremely unlikely event of simultaneous electrical failures which could result in single RCCA withdrawal, rod deviation and rod' control failure would both be displayed on the plant annunciator, and the rod position indica- ' tors would indicate the relative positions of the assemblies in the bank. The. urgent failure alarm also inhibits automatic rod motion in the group in which it occurs. Withdrawal of a single RCCA by operator action, whether deliberate or by a combination of errors, would result in activation of.the same alarm and the same visual indications. Withdrawal of a single RCCA results in both positive reactivity insertion tending to increase core power, and an increase - in local power density in the core area associated with the RCCA. Automatic prctection for this event is provided by the overtemperature AT reactor trip;-although, due to the increase in local power density, it is not possible in all cases to provide assurance that the core safety limits will not be violated. 15.4.3.2 Analysis of Effects and Consecuences 15.4.3.2.1 Dropped RCCAs, Dropped RCCA Bank, and Statically Misaligned RCCA -0952T:6 15.4-16
l 15.4.3.2.1.1 Method of Analysis I A. One or more dropped RCCA ' rom the same group For evaluation of the eaopped RCCA event, the transient system response is calculated using the LOFTRAN(3) code. The code simulates the neutron kinetics, Reactor Coolant System, pres-surizer, pressurizer relief and safety valves, pressurizer spray, steam generator, and steam generator safety valves. The code computes pertinent plant variables including temperatures, pressures, and power level. Statepoints are calculated and nuclear models are used te obtain a hot l channel factor consistent with the primary system conditions and reactor power. By incorporating the primary conditions from the tran- , l siert and the hot channel factor from the nuclear analysis, the DNB ! l ( design basis is shown to be met using the THINC code. The transient response, nuclear peaking factor analysis, and DNB design basis con-firmation are performed in accordance with the methodology described in Reference 4. B. Statically Misaligned RCCA Steady state power distributions are analyzed using the computer codes as described in the McGuire FSAR Section 4.4. The peaking factors are then used as input +.o the THINC code to calculate the DNBR.
-15.4.3.2.1.2 Results A. One or more diopped RCCAs -
Single or multiple RCCAs within the ~same group result in a negative-
. reactivity insertion which may be detected by the power range negative neutron flux rate trip circuitry. If detected, the reactor is tripped 0952T:6 15.4-17
within approximately 2.5 seconds following the drop of the RCCAs. The core is not adversely affected during this period, since power is decreasing rapidly. Following reactor trip, normal shutdown proce-dures are followed. The operator may manually retrieve the RCCA by following approved operating procedures. For those dropped RCCAs which do not result in a reactor. trip, power may be reestablished either by reactivity feedback or control bank wi thdrawal . Following a dropped rod event in manual rod control, the plant will establish a new equilibrium condition. The equilibrium process without control system interaction is monotonic, thus removing power overshoot as a concern, and establishing the automatic rod con-trol mode of operation as the limiting case. ; For a dropped RCCA event in the automatic rod control mode, the Rod Control System detects the drop io power and initiates control bank withdrawal. Power overshoot may occur due to this action by the auto-matic rod controller after which the control system will insert the control bank in order to restore nominal power. Figures 15.4.3-1 and I 15.4.3-2 show a typical transient response to a dropped RCCA (or RCCAs) in automatic control. In all cases the minimum DNBR remains above the limit value. B. Dropped RCCA Bank A dropped RCCA bank typically results in a reactivity insertion greater than 500 pcm which will be detected by the power range nega-tive neutron flux rate trip circuitry. The reactor is tripped within approximately 2.5 seconds following the drop of an RCCA. The core is not adversely affected during this period, since the power is decreas-ing rapidly. Following reactor trip, normal shutdown procedures may subsequently be followed to further cool down the plant. E 0952T:6 15.4-12
C. Statically Misaligned RCCA The most severe misalignment situations with respect to DNBR at signi-( ficant power levels arise from cases in which one RCCA is fully inser-ted, or where bank D is fully inserted with one RCCA fully withdrawn. Multiple independent alarms, including a bank insertion limit alarm, alert the operator well before the postulated conditions are approached. The bank can be inserted to its insertion limit with any one assembly fully withdrawn without the DNBR falling below the limit value. Any action required of the operator to maintain the plant in a stabilized condition will be in a time frame in excess of ten minutes following the incident. The insertion limits in the Technical Specifications n'ay vary from time to time depending on a number of limiting criteria. It is pre-ferable, therefore, to analyze the misaligned RCCA case at full power for position of the control bank as deeply inserted as the criteria on minimum DNBR and power peaking factor will allow. The full power C insertion limits on control bank D must then be chosen to be above that position and will usually be dictated by other criteria. Detailed results will vary from cycle to cycle depending on fuel arrangements. . For this RCCA misalignment, with bank D inserted to its full power-insertion limit and one RCCA fully withdrawn, DNBR does not fall below the limit value. This case is analyzed assuming the initial reactor power, pressure, and RCS temperature are at their nominal values as given in Table 15.0.3-3, but with the increased radial peaking factor associated with the misaligned RCCA. f DNB calculations have not been performed'specifically-for RCCAs mis-sing from other banks; however, power shape calculations have been performed as required for the RCCA ejection analysis. Inspection of the power shapes'shows that the DNB and peak kw/ft situation is less-severe than the bank D case discussed above assuming insertion limits on the other banks equivalent to a bank D full-in insertion limit. L 0952T:6 -15.4-19
For RCCA misalignments with one RCCA fully inserted, the DNBR does not fall below the limit value. This case is analyzed assuming the ini-tial reactor power, pressure, and RCS temperatures are at their nomi- ; nal values as given in Table 15.0.3-3 but with'the increased radial peaking factor associated with the misaligned RCCA. i l DNB does not occur for the RCCA misalignment incident and thus the ability of the primary coolant to remove heat from the fuel rod is not reduced. The peak fuel temperature corresponds to a linear heat gene-
~
ration rate based on the radial peaking factor penalty associated with the misaligned RCCA and the design axial power distribution. The resulting linear heat generation is well below that which would cause fuel melting. Following the identification of a RCCA misalignment condition by the opera-tor, the operator is required to take action as required by the plant Tech-nical Specifications and operating instructions. 15.4.3.2.2 Single RCCA Withdrawal 15.4.3.2.2.1 Method of Analysis. Power. distributions within the core are-calculated by the computer. codes as described in FSAR Table 4.1-2. ~ The peak-ing factors are then used by THINC to calculate the minimum DNBR for the event. The case of the worst rod withdrawn from bank D inserted at the 4 insertion limit, with the reactor initially at full power, was analyzed. This' incident is assumed to occur at beginning-of-life since this results in the minimum value of moderator temperature coefficient. -This assumption maximizes the power rise and minimizes the tendency of increased moderator. temperature - to flatten the power distribution. 4 15.4.3.2.2.2 Results of the Analysis. For the single rod withdrawal event, two cases have been considered as follows. .? c 0952T:6 <15.4-20
1 A. If the reactor is in the manual control mode, continuous withdrawal of .! a single RCCA results in both an increase in core power and coolant temperature, and an increase in the local hot channel factor in the l area of the withdrawing RCCA. In terms of the overall system (' response, this case is similar to those presented in subsection 15.4.2; however, the increased local power peaking in the area of the withdrawn RCCA results in lower minimum DNBRs than for the withdrawn bank cases. Depending on initial bank insertion and location of the withdrawn RCCA, automatic reactor trip may not occur sufficiently fast to prevent the minimum core DNBR from falling below the limit value. Evaluation of this case at the power and coolant conditions at which the overtemperature AT trip would be expected to trip the plant shows that an upper limit for the number of rods with a DNBR less than the limit value is 5 percent. B. If the reactor is in the automatic control mode, the multiple failures that result in the withdrawal of a single RCCA will result in the ( immobility of the other RCCAs in the controlling bank. The transient will then proceed in the same manner as case A described above. For such cases as above, a reactor trip will ultimately ensue, although not sufficiently fast in all cases as to prevent a minimum DNB ratio in the core of less than the limit value. Following reactor trip, normal shutdown procedures may be followed. 15.4.3.3 Conclusions For cases of dropped RCCAs or dropped banks, for which the reactor is tripped by the power range negative neutron flux rate trip, there is no reduction'in the margin to core thermal limits, and consequently the DNB design basis is met. It is shown for all cases which do not result in reactor trip that the DNBR remains greater than the limit value and, therefore, the DNB design is met. t , 0952T:6 15.4-21
l For all cases of any RCCA fully inserted, or bank D inserted to its rod inser-tion limits with any single RCCA in that bank fully withdrawn (static mis-alignment), the DNBR remains greater than the limit value. For the case of the accidental withdrawal of a single RCCA, with the reactor in the automatic or manual control mode and initially operating at full power with bank D at the insertion limit, an upper bound of the number of fuel rods experiencing DNB is 5 percent of the total fuel rods 'n the core. 15.4.4 STARTUP OF AN INACTIVE REACTOR COOLANT PUMP AT AN INCORRECT TEMPERATURE 15.4.4.1 Identificatio i of Causes and Accident Description If the plant is operating with one pump out of service, there is reverse flow througn the inactive loop due to the pressure difference across the reactor vessel. The cold leg temperature in an inactive loop is identical tt the cold leg temperature of the active loops (the reactor core inlet temperature). If the reactor is operated at power, and assuming the secondary side of the steam generator in the inactive loop is not isolated, there is a temperature drop across the steam generator in the inactive loop anc, with the reverse flow, the hot leg temperature of the inactive loop is lower than the reactor core inlet temperature. Starting of an idle reactor coolant pump without bringing the inactive loop hot leg temperature close to the core inlet temperature would result in the injection of cold water into the core, which would cause a reactivity insertion and subsequent power increase. This-event is classified as an ANS Condition II incident (an incident of moderate frequency) as defined in Section 15.0.1. . Should the startup of an inactive reac' tor coolant pump accident occur, the transient vill be terminated automatically by a reactor trip on low coolant loop flow when the power range neutron flux (two out of four channels) exceeds the P-8 setpoint, which has been previously reset for three loop operation. 0952T:6 15.4-22
15.4.4.2 Analysis of Effects and Consecuences Method of Ana' lysis This transient is analyzed by three digital computer codes. The LOFTRAN Code (Reference 3) is used to calculate the loop and core flow, nuclear power and core press.ure and temperature . transients following the startup of an idle pump. FACTRAN (Reference 2) is used to calculate the core heat flux transient based on core flow and nuclear power from LOFTRAN. The THINC Code is tnen used to calculate the DNBR during the transient based on -system conditions (pressure, temperature, and flow) calculated by LOFTRAN and heat flux as' calculated by FACTRAN. This accident is analyzed with the Improved Thermal Design Procedure as described in WCAP-8567. Plant characteristics and initial conditions are discussed in Section 15.0.3. _ ( In order to obtain conservative results for the startup of an inactive pump accident, the following assumptions are made:
- 1. Initial reactor power, pressure, and RCS temperatures are assumed to be at <-
their nominal N-1 loop operation values. Uncertainties in initial conditions are included in the limit DNBR as described in WCAP-8567.
- 2. Following initiation of startup of the idle pump, the inactive loop flow reverses and accelerates to its nominal flow value in approximately 20 seconds. This value is faster than the expected startup time, and 1s -
f conservative for this analysis, l
- 3. A conservatively large moderator density coefficient.
- 4. A conservatively small (absolute value) negative Doppler ' power coefficient.
- 5. The initial reactor coolant luop flows are at the appropriate values for one pump out of service.
l 1 ! 0952T:6 15.4-23 I .
- i. ,
1
- 6. The reactor trip is assumed to occur on low coolant flow when the power range neutron flux exceeds the P-8 setpoint. The P-8 setpoint is conser-vatively assumed to be 84 percent of rated power which corresponds to the nominal setpoint plus 9 percent for nuclear instrumentation errors.
Results , The results following the startup of an idle pump with the above listed assumptions are shown in Figures 15.4.4-1 through 15.4.4-5. As shown in these curves, during the first part of the transient, the increase in core flow with cooler water results in an increase in nuclear power and a decrease in core average temperature. The minimum DNBR during the transient is considerably greater than the limit value. Reactivity addition for the inactive loop stcrtup accident is due to the decrease in core water temperature. Duri.ng the transient, this decrease is due both to the increase in reactor coolant ficw and, as the inactive loop flow reverses, to the colder water entering the. core from the hot leg side - (colder temperature side prior to the start 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 in Figure 15.4.4-1. The calculated sequence of events for this accident is shown in Table 15.4-1. The transient results illustrated in Figures 15.4.4-1 through 15.4.4-5 indicate that a stabilized plant condition, with the reactor tripped, is approached rapidly. Plant cooldown may subsequently be achieved by following-normal shutdown procedures. 15.4.4.3 Conclusions The transient results show that the core is not adversely affected. There is consicerable margin to the limit DNBR value; thus, no-fuel or clad damage is predicted. l 0952T:6 15.4-24 l
( 15.4.5 SPECTRUM OF RCD CLUSTER CONTROL ASSEMBLY EJECTION ACCIDENTS 15.4.5.1 Identification of Causes and Accident Description ( This accident is defined as the mechanical failure of a control rod mechanism pressure housing resulting in the ejection of a rod cluster control assembly (RCCA) and drive shaft. The consequence of this mechanical failure is a rapid positive reactivity insertion together with an adverse core power distribu-tion, possibly leading to localized fuel rod damage. 15.4.5.1.1 Design Precautions and Protection Certain features are intended to preclude the possibility of the rod ejection accident, or to limit the consequences if the accident were to occur. .These include a sound, conservative mechanical design of the rod housings, togetner with a thorough quality control (testing) program during assembly, and a nuclear design which lessens the potential ejection worth of RCCAs, and ( minimizes the number of assemblies inserted at high power levels. Mechanical Design Mechanical design and quality control procedures intended to preclude the possibility of a RCCA drive mechanism housing failure are listed below: 1. Each full-length control rod drive mechanism housing is completely assembled and shop tested at 4100 psi.
- 2. The mechanism housings are individually hydrotested after they are attached to the head adapters in the reactor vessel head, and checked during the hydrotest of the completed reactor coolant system.
1 0952T:6 15.4-25.
- ___ _ _ _ - _ _ = _ _ _
- 3. Stress levels in the mechanism are not affected by anticipated system transients at power, or by the thermal movement of the coolant loops.
Moments induced by the design-basis earthquake can be accepted within the allowable primary working stress range specified by the ASME Code, Section III, for Class 1 cumponents.
- 4. The latch mechanism housing and rod travel housing are each a single ls length of forged Type 304 stainless steel. This material exhibits I excellent notch toughness at all temperatures which will be encountered.
A significant margin of strength in the elastic range together with the large energy absorption capability in the plastic range gives additional assurance that gross failure of the housing will not occur. The joints between the f latch mechanism housing and head adapter, and between the latch mechanism housing and rod travel housing, are threaded joints reinforced by canopy type rod welds. Administrative regulations re. quire periodic inspections of these l (and other) welds. Nuclear Design Even if a rupture of a RCCA drive mechanism housing is postulated, the operation of a plant utilizing chemical shim is such that the severity of an , ejected ECCA is inherently limited. In general, the reactor is operated with the RCCA's inserted only far enough to permit load follow. Reactivity changes caused by core depletion and xenon transients are compensated by boron changes. Further, the location and grouping of control RCCA banks are selected during the nuclear design to lessen the severity of a RCCA ejection accident. Therefore, should a RCCA be ejected from its normal position during - full power operation, only a minor reactivity excursion,-at worst, could be expected to occur. However, it may be occasionally desirable to operate with-larger than normal insertions. For this reason, a rod insertion limit is defined as a function of power level. Operation with the RCCAs above this limit guarantees 0952T:6 15.4-26' m _ _ __.__._ _____
adquate shutcown capability and acceptable power distribution. The position of all RCCAs is continuously indicated in the control room. An alarm will occur if a bank of RCCAs approaches its insertion limit or if one RCCA ( deviates from its bank. Operating inscructions require boration at low level alarm and emergency boration at the low-low alarm. Reactor Protection The reactor protection in the event of a rod ejection accident has been described in Reference 5. The protection for this accident is provided by high neutron flux trip (high and low setting) and high rate of neutron flux increase trip. Effects on Adjacent Housings Disregarding the remote possibility of the occurrence of,a RCCA mechanism hcusing failure, investigations have shown that failure of a housing due to either longitudinal or circumferential cracking would not cause damage to adjacent housings. However, even if damage is postulated, it would not be expected to lead to a more severe transient since RCCAs are inserted in the core in symmetric patterns, and control rods immediately adjacent to worst ejected rods are not in the core when the reactor is critical. Damage to an adjacent housing could, at worst, cause that RCCA not to fall on receiving a trip signal; however, this is already taken into account in the analysis by assuming a stuck rod adjacent to the ejected rod. 15.4.5.1.2 Limiting Criteria This event is classified as an ANS Conditien IV incident. See Section 15.0.1 for a discussion of ANS classifications. Due to the extremely low probability of a RCCA ejection accident, some fuel damage could be considered an -
)
acceptable consequence. L 0952T:6 15.4-27
l Comprehensive studies of the threshold of fuel failure and of the threshold or i significant conversion of the fuel thermal energy to mechanical energy, have been carried out as part of the SPERT project by the Idaho Nuclear Corporation (Reference 6). Extensive tests of zirconium clad UO 2 fuel r ds representa-tive of those in Pressurized Water Reactor type cores have demonstrated failure thresholds in the range of 240 to 257 cal /gm. However, other rods of a sightly different design have exhibited failures as low as 225 cal /gm. These results differ significantly from the TREAT (Reference 7) results, which indicated a failure threshold of 280 cal /gm. Limited results have indicated that this threshold decreases by about 10 percent with fuel burnup. The clad failure mechanism appears to be melting for zero burnup rods and brittle fracture for irradiated rods. Also important is the conversion ratio of thermal to mechanical energy. This ratio becomes marginally detectable above 300 cal /gm for unirradiated rods and 200 cal /gm for irradiated rods; catastrophic failure, (large fuel dispersal,-large pressure rise) event for irradiated rods, did not occur below 300, cal /gm. In view of the above experimental results, criteria are applied to ensure that there is little or no possibility of fuel dispersal in the coolant, gross lattice distortion, or severe shock waves. These criteria are:
- 1. Average fuel pellet enthalpy at the hot spot below 225 cal /gm for unirradiated fuel and 200 cal /gm for irradiated fuel.
- 2. Average clad temperature at the hot spot below the teaperature at which clad embrittlement may be expected (2700'F).
- 3. Peak reactor coolant pressure less than that which could cause stresses to-exceed the faulted condition stress -limits.
- 4. Fuel melting will be limited to less than ten percent of the fuel volume at the hot spot even if the average fuel pellet enthalpy is below the limits of criterion 1 above.
0952T:6 15.4-28
't I
15.4.5.2 Analysis of Effects and Consecuences Method of Analysis ( The calculation of the RCCA ejection transient is performed in two stages, first an average core channel calculation and then a hot region calculation. The average core calculation is performed using spatial neutron kinetics methods to determine the average power generation with time including the various total core feedback effects, i.e., Doppler reactivity and moderator reactivity. Enthalpy and temperature transients in the hot spot are then determined by multiplying the average core energy generation by the hot channel factor and performing a fuel rod transient heat transfer calculation. The power distribution calculated without feedback is pessimistically assumed to persist throughout the transient. AdetaileddiscussionofthemethodofanalysiscanbefoundinReference(5]. ( Average Core Analysis The spatial kinetics computer code, TWINKLE (Reference 1), is used for the average core transient analysis. This code solves the two group neutron diffusion theory kinetic equation in one, two or three spatial dimensions (rectangular coordinates) for six delayed neutron groups and up to 2000 spatial points. The computer code includes a detailed multiregico, transient fuel-clad-coolant heat transfer model for calculation of pointwise Doppler and moderator feedback effects. In this analysis, the code is used as a one dirtansional axial kinetics code since it allows a more realistic representa-tion of the spatial effects of axial moderator feedback and RCCA movement. l However, since the radial dimension is missing, it is still necessary to l employ very conservative methods (described in the following) of calculating the ejected rod worth and hot channel factor. Further oescription of TWINKLE appears in Section 15.0.7. l l l . \ i 0952T:6 15.4-29 l
Hot Spot Analysis In the hot spot analysis, the initial heat flux is equal to the nominal times the design hot channel factor. During the transient, the heat flux hot channel factor is linearly increased to the transient value in 0.1 second, the time for full ejection of the rod. Therefore, the assumption is made that the hot spot before and after ejection are coincident. This is very conservative since the peak after ejection will occur in or adjacent to the assembly with the ejected rod, and prior to ejection the power in this region will neces-sarily be depressed. The hot spot analysis is performed using the detailed fuel and cladding transient heat transfer computer code, FACTRAN (Reference 2). This computer code calculates the transient temperature distribution in a cross section of a metal clad UO 2 fuel r d, and the heat flux at the surface of the rod, using as input the nuclear power versus time and the 1 scal coolant conditions. The zirconium-water reaction is explicitly represented, and all material proper-ties are represented as functions of temperature. A conservative pellet radial power distribution is used within the fuel rod. FACTRAN uses the Dittus-Boelter or Jens-Lottes correlation to determine the film heat transfer before DNB, and the Bishop-Sandberg-Tong correlation (Reference 8) to determine the film boiling coefficient after DNB. The BST correlation is conservatively used assuming zero culk fluid quality. The DNB ratio is not calculated, instead the code is forced into DNB by specifying a conservative DNB heat flux. The gap heat transfer coefficient can be calcu-lated by the code; however, it is adjusted in order to force the full power-steady-state temperature distribution to agree with the fuel heat transfer - design codes. Further description of FACTRAN appears in Section 15.0.7. l c 0952T:6 15.4-30 -
l l l t ( i System Overpressure Analysis ! 1 Because safety limits for fuel damage specified earlier are not exceeded, there is little likelihood of fuel dispersal into the ecolant. The pressure (~ surge may therefore be calculated on the basis of conventional heat transfer from the fuel and prompt heat generation in the coolant. The pressure surge is calculated by first performing the fuel heat transfer calculation to determine the average and hot spot heat flux versus time. Using this heat flux data, a THINC calculation is conducted to determine the volume surge. Finally, the volume surge is simulated in a plant transient computer code. This code calculates the pressure transient taking into account fluid transport in the reactor coolant system and heat transfer to the steam generators. No credit is taken for the possible pressure reduction caused by the assumed failure of the control rod pressure housing. 15.4.5.2.1 Calculation of Basic Parameters ( Input parameters for the analysis are conservatively selected on the basis of values calculated for this type of core. The more important parameters are discussed below. Table 15.4-2 presents the parameters used in this analysis. Ejected Rod Worths and Hot Channel Factors ' The values for ejected rod worths and hot channel factors are calculated using either three dimensional static methods or by a synthesis method employing one dimensional and two dimensional calculations. Standard nuclear design codes are used in the analysis. No credit is taken for the flux flattening effects of reactivity feedback. The calculation is performed for the maximum allowed bank instrtion at a given power level, as determined by the red insertion limits. Adverse xenon distributions are considered in the calculation. Appropriate margins are added to the ejected rod worth and hot channel factors to account for any calculational uncertainties, including an allowance for nuclear power peaking due to densification. L - 0952T:6 15.4-31.
1 1 Power distributions before and after eje-: tion for a " worst case" can be fou1d ; in Reference 5. During plant startup physics testing, ejected rod worths and power distributions are measured in the zero and full power rodded ' configurations and compared to values used in ths analysis. It has been found that the ejected rod worth and power peaking factors are consistently overpredicted in the analysis. Reactivity Feedback Weichting Factors The largest temperature rises, and hence the largest reactivity feedback.e, occur in channels where the power is higher than average. Since the weight of a region is dependent on flux, thase regions have high weights. This means that the reactivity feedback is larger than that indicated by a simple channel analysis. Physics calculations have been carried out for temperature changes with a flat temperature distribution, and with a large number of axial and radial temperature distributions. Reactivity changes were compared and y effective weighting factors determined. These weighting factors take the form of multipliers which when applied to single channel feedbacks correct them to effective whole core feedbacks for the appropriate flux shape. In this analysis, since a one-dimensional (axial) spatial kinetics method is employed, axial weighting is not necessary if the initial condition is made to match the ejected rod configuration. In addition, no weighting is applied to the moderator feedback. A conservative radial weighting factor is applied to the transient fuel temperature to obtain an effective fuel temperature as a function of time acccunting for the missing spatial dimension. These weighting factors have also been shown to be conservative compared to three dimensional snalysis (Reference 5). Moderator and Doppler Coefficient The critical boron concentrations at the beginning of life and end of life are adjusted in the nuclear code in order to obtain moderator _ density coefficient curves which are conservative compared to actual design conditions for the - plant. As discussed above, no weighting factor is apolied to these results. 0952T:6 15.4-32
l l The Doppler reactivity defect is determined as a function of power level using a one dimensional steady-state computer code with a Doppler weighting factor of 1.0. The Doppler defect used is given in Section 15.0.4. The Doppler weighting factor will increase under accident conditions, as discussed above. Delayed Neutron Fraction, B ( Calculations of the effective delayed neutron fraction (8,ff) typically yield values no less than 0.70 percent at beginning of life and 0.50 percent at end of life for the first cycle. The accident is sensitive to 8 if the ejected rod worth is equal to or greater than B as in zero power transients. In order to allow for future cycles, pessimistic estimates of S of 0.50 percent at beginning of cycle and 0.44 percent at end of cycle were used in the analysis. Trio Reactivity Insertion ( The trip reactivity insertion assumed is given in Table 15.4-2 and includes the effect of one stuck RCCA. These values are reduced by the ejected rod reactivity. The shutdown reactivity was simulated by dropping a rod of the required worth into the core. The start of rod motion occurred 0.5 seconds after the high neutron flux trip point was reached. This delay is assumed to consist of 0.2 seconds for the instrument channel to produce a signal, 0.15 seconds for the trip breaker to open and 0.15 seconds for the coil to release the rods. A curve of trip rod insertion versus time was used which assumed that insertion to the dashpot does not occur until 3.3 seconds after the start of fall. The choice of such a conservative insertion rate means that there is over one second after the trip point is reached before significant shutdown reactivity is inserted into the core. This is a particularly important conservatism for hot full power accidents. The minimum design shutdown available for this plant at-HZP may be reached only at end of life in the equilibrium cycle. This value includes-an allowance for the worst stuck rod, adverse xenon distribution, conservative L 0952T:6 15.4-33 1
Doppler and moderator defects, and an allowance for calculational uncertain-ties. Physics calculations for this plant have shown that the effect of two stuck RCCAs (one of which is the worst ejected rod) is to reduce tne shutdown by about an additional one percent Ak, Therefore, following a reactor trip resulting from an RCCA ejection accident, the reactor will be suberitical when the core returns to HZP. Depressurization calculations have been performed for a typical four-loop plant assuming the maximum possible size break (2.75 inch diameter) located in the reactor pressure vessel head. The results show a rapid pressure drop and a decrease in system water mass due to the break. The safety injection system ' is actuated on the coincidence of low pressurizer pressure and level within one minute after the break. The reactor coolant pressure continues to drop and reaches saturation (1100 to 1300 psi depending on the system temperature) in about two to three minutes. Due to the large thermal inertia of primary and secondary system, there has been no significant decrease in the reactor coolant system temperature below no-load by this time, and the depressuriza-tion itself has caused an increase in shutdown margin by about 0.2 percent Ak'due to the pressure coefficient. The cooldown transient could not absorb the available shutdown margin until more than ten minutes after the break. The addition of borated (2000 ppm) safety injection flow starting one minute after the break is much more. than sufficient to ensure that the core remains subcritical during the cooldown. Reactor Protection As discussed in Section 15.4.5.1.1, reactor protection for a red ejection is provided by high neutron flux trip (high and low setting) and high rate of neutron flux increase trip. These protection functions are part of the reactor trip system. No single failure of the reactor trip system will negate the protection functions required for the rod ejection accident, or adversely affect the consequences of the accident. 0952T:6 15.4-34 l
l I l Results
)
I Cases are presented for both beginning and end of life at zero and full power.
- 1. Beginning of Cycle, Full Power Control bank D was assumed to be inserted to its insertion limit The worst ejected rod worth and hot channel factor were conservatively calculated to be 0.23 percent Ak and 5.3 respectively. The peak hot spot clad average temperature was 2270*. Tne peak hot spot fuel center temperature reached melting, conservatively assumed at 4900*F. However, melting was restricted to less than 10 percent of the pellet.
- 2. Beginning of Cycle, Zero. Power For this condition, control bank D was assumed to be fully inserted and banks B and C were at their insertion limits. The worst ejected rod is
( located in control bank D and has a worth of 0.75 percent Ak and a hot channel factor of 11.5. The peak het spot clad temperature reachec 2621 F, the fuel center temperature was 4101*F.
- 3. End of Cycle, Full Power.
Control bank D was assumed to be inserted to its insertion limit. The ejected rod worth and hot channel factors were conservatively calculated to be 0.23 percent Ak and 5.9 respectively. This resulted in a peak clad temperature of 2175*F. The peak hot spot fuel temperature reached melting conservatively assumed at 4800 F. However, melting was restricted to less than 10 percent of the pellet. L 0952T:6 15.4-35
- 4. End of Cycle, Zero Power
, The ejected rod worth and hot channel factor for this case were obtained assuming control bank D to be fully inserted and bank C at its insertion limit. The results were 0.90 percent Ak and 20.0 respectively. The peak clad and fuel center temperatures were 2685'F and 3993 F. The ( Doppler weighting factor for this case is significantly higher than for the other cases due to the very large transient hot channel factor. A summary of the cases presented above is given in Table 15.4-2. The nuclear power and hot spot fuel and clad temperature transients for the worst cases are presented in Figures 15.4.5-1 through 15.4.5-4. (beginning-of-life full power and end-of-life zero power.) The calculated sequence of events for the worst case rod ejection accidents, as shown in Figures 15.4.5-1 through 15.4.5-4, is presented in Table 15.4-1. For all cases, reactor trip occurs very early in the transient, after which the nuclear power excursion is terminated. As discussed previously in Section - 15.4.5.2.2, the reactor will remain subcritical following reactor trip. The ejection of an RCCA constitutes a break in the Reactor Coolant System, located in the reactor pressure vessel head. Following the RCCA ejection, the operator would follow the same emergency instructions as for any other loss of coolant accident to recover from the event. Fission Product Release It is assumed that fission products are released from the gaps of all rods entering DNB. In all cases considered, less than 10 percent of the rods entered DNB based on a detailed three dimensional THINC analysis (Reference 5). t 4 -0952T:6 15.4-36 _________________m__ _
I Pressure Surce A detailed calculation of the pressure surge for an ejection worth of one dollar at beginning of life, hot full power, indicates that the peak pressure C does not exceed that which would cause stress to exceed the faulted conditio.1 stress limits (Reference 5). Since the severity of the present analysis does not exceed the " worst case" analysis, the accident for this plant will not result in an excessive pressure rise or further damage to the Reactor Coolant System. Lattice deformations A large temperature gradient will exist in the region of the hot spot. Since the fuel rods are free to move in the vertical direction, differential expansion between separate rods cannot produce distortion. However, the temperature gradients across individual rods may produce a differential expansion tending to bow the midpoint of the rods toward the hotter side of ( the rod. Calculations have indicated that this bowing would result in a-negative reactivity effect at the hot spot since Westingnouse cores are under-moderated, and bowing will tend to increase the under-moderation at the hot spot. Since the 17 x 17 fuel design is also under-moderated, the same effect would.be observed. In practice, no significant bowing is anticipated, since the structural rigidity of the core is more than sufficient to withstand the forces produced. Boiling in the hot spot region would produce a net flow away from that region. Hewever, the heat from the fuel is released to the water relatively slowly, and it is considered inconceivable that crossflow will be sufficient to produce significant lattice forces. Even if massive and rapid boiling, sufficient to distort the lattice, is hypothetically postu-lated, the large void fraction in the hot spot region would produce a reduc-tion in the total core moderator to fuel ratio, and a large reduction in this ratio at the hot spot. The net effect would therefore be a negative feed-
.back. It can be concluded that no conceivable mechanism exists for a net positive feedback resulting from lattice deformation. In fact, a small negative feedback may result. The effect is conservatively ignored in the analysis.
0952T:6 15.4-37
! 15.4.5.3 Corciusions - Even on a pessimistic basis, the analyse's indicate that the described fuel and clad limits are not exceeded. It is concluded that there is no danger of sudden fuel dispersal into the coolant. Since the peak pressure does not exceed that which would cause stresses to exceed the faulted condition stress limits, it is concluded that there is no danger of further consequential damage to the Reactor Coolant System. The analyses have demonstrated that the upper limit in fission product release as a result of a number of fuel rods entering DNB amounts to ten percent, t i 0952T:6 15.4 i
1 15.
4.6 REFERENCES
- 1. Risher, D. H. , Jr. and Barry, R. F., " TWINKLE - A Multi-Dimensional
( Neutron Kinetics Computer Code," WCAP-7979-A (Proprietary) and WCAP-8028-A (Non-Proprietary), January 1975.
- 2. Hargrove, H. G., "FACTRAN - A Fortran-IV Lode for Thermal Transients in a UO2 Fuel Rod," WCAP-7908, June 1972.
- 3. Burnett, T. W. T. , et al . , "LOFTRAN Code Description," WCAP-7907, June 1972.
- 4. Morita, T. , et al. , " Dropped Rod Methodology for Negative Flux Rate Trip Plants," WCAP-10297-P-A, June 1983.
- 5. Risher, D. H., Jr., "An Evaluation of the Rod Ejection Accident in Westinghouse Pressurized Water Reactors Using Spatial Kinetics Methods,"
4 WCAP-7588, Revision I-A, January 1975.
- 6. Taxelius, T. G. (Ed), " Annual Report - SPERT Project, October,.1968, September,1969," Idaho Nuclear Corporation IN-1370, June 1970.
- 7. Liimataninen, R. C. and Testa, F. J., " Studies in TREAT of Zircaloy-2-Clad, UO2-Core Simulated Fuel Elements," ANL-7225, January -
June 1966, p.177, November 1966.
- 8. Bishop, A. A., Sandberg, R. O., and Tong, L. S., " Forced Convection Heat Transfer at High Pressure After the Critical Heat Flux," ASME 65-HT-31, .
August 1965. l l l i 0952T:6 ~15.4-39 -l l
\
y - -- ,n
L TABLE 15.4-1 (Page 1) TIME SE00ENCE OF EVENTS FOR INCIDENTS WHICH CAUSE Accident REACTIVITY AND p0WER DISTRIBUTION ANOMALIES Event ( Time (sec.) Uncontrolled Rod Initiation of uncontrolled 0.0 Cluster Control rod withdrawal from 10 -9 of Assembly Bank nominal power Withdrawal from a Suberitical or Low Power Startup Condition Power range high neutron 10.4 flux low setpoint reached Peak nuclear power occurs 10.5 Rod begin to fall into core 10.9 Peak heat flux occurs. 13.0 Minimum DNBR occurs 13.0 Peak average clad temperature- 13.4 occurs f Peak average fuel temperature 13.7 occurs J 0952T:6 15.4-40
TABLE 15.4-1 (Page 2) TIME SEOUENCE OF EVENTS FOR INCIDENTS WHICH CAUSE REACTIVITY AND POWER DISTRIBUTION ANOMALIES C Accident Event Time (sec.) Uncontrolled RCCA bank withdrawal at power
- 1. Case A Initiation of uncontrolled RCCA 0 withdrawal at a high reactivity 4
insertion rate (75 pcm/sec) s Power range high neutron flux 1.11 , high trip point reached Rods begin to fall into core 1.61 Minimum DNBR occurs 2.80
- 2. Case B Initiation of uncontrolled RCCA 'O withdrawal at a small reactivity insertion rate (1 pcm/sec) ,
Overtemperature AT reactor 104.1 trip signal initiated Rods begin to' fall into core 106.1 Minimum DNBR occurs 107.2 l 0952T:6 15.4-41
I l i i 1 TABLE 15.4-1 (Page 3) 1 TIME SEOUENCE OF EVENTS FOR INCIDENTS WHICH CAUSE REACTIVITY AND p0WER DISTRIBUTION ANOMALIES Accident Event Time (sec.) Startup of an Initiation of pump startup 2.0 inactive reactor coolant loop at Power reaches P-8 trip 9.1 an incorrect setpoint temperature { Rods begin to drop 9.6 Minimum DNBR occurs 10.4 , Rod Cluster Control Assembly Ejection
- 1. Beginning-of- Initiation of rod ejection 0.0 Life, Full Power
, Power range high neutron flux 0.05 setpoint reached Peak nuclear power occurs 0.14 Rods begin to fall into core 0.55 Peak fuel average temperature 2.34 occurs Peak clad temperature occurs '2.41
~
1 ' l . 0952T:6 15.4-42 e.
TABLE 15.4-1 (Page 4) TIME SEQUENCE OF EVENTS FOR INCIDENTS WHICH CAUSE REACTIVITY AND POWER DISTRIBUTION ANOMALIES I Accident Event Time (sec.) Peak heat flux occurs 2.42 Peak fuel center temperature 2.91 occurs i 2. End-of-Life, Initiation of rod ejection 0.0 Zero Power j l Power range high neutron flux 0.18 low setpoint reached ( End-of-Li f e , Zero Power (cont.) Peak nuclear power occurs Rods begin to fall into core 0.21 0.68 } Peak clad temperature occurs 1.56 d 4 Peak heat flux occurs 1.56 Peak fuel average temperature 1.88 occurs Peak fuel center temperature. 2.84 occurs.
.0952T:6- *15.4-43
--_. e
TABLE 15,4-g l PARAMETERS USED IN THE ANALYSIS OF THE ROD CLUSTER CONTROL ASSEMBLY EJECTION ACCIDENT BEGINNING BEGINNING END OF ENO OF TIME IN LIFE OF CYCLE OF CYCLE CYCLE CYCLE Power level, Percent 102 0 102 0 Ejected rod worth, Percent AK 0.23 0.75 0.23 0.90 Delayed neutron fraction, Percent 0.50 0.50 0.44 0.44 , 1 Feedback reactivity weighting 1.30. 2.07 1.30 3.55 Trip reactivity, Percent AK 4.0 2.0 4.0 2.0 Fq before rod ejection 2.50 - 2.50 - F after rod ejection 5.3 11.5 5.9 q 20.0 Number of operational pt.mps 4 2 4 2 Max. fuel pellet average 4113 3519 3951 3502 temperature, *F Max. fuel center temperature, *F 5006 4101 4893 3993 Max. clad average tempera- 2270 2621 2175 2635 ture, "F - Max. fuel store energy, cal /gm 180 150 172 149 Pr.rcent fuel melt <20 0 <10- 0 0952T:6 15.4-44 e
l 2 10 ,, ; , C a 2 .. I g .. i C 1 2 10 == .. = u. ( C ~; -- 2 O
~~
c 4 m 100: : :
.u. ::
m -: " w .. g .. C ..
- a. ..
m -1 m 10 55 . 55 a u .. C .. g .. C .. 10-2 , 0 5 10 15 20 25 30 I TIME (SECONOS) l
~
FIGURE 15.4.1-1
)
Neutron Flux Tranment for Uncorrtrolled Rod Witndrawed from a Sucentical Condition l
- 15.4-45
- -v w
( 1.2 1.0 * ~ E_. C E z - 8 I
~
5
;- . 6" -
W ( ~ 5 CK b. a
- 4 .
E ' 2-
$ I N
0 4.0 - 3.5 - 3.0 - - r e 2.5 -
- 2.0 - -
1.5 ' ~ 1.0 0 10 20 30 40 50 60 L TIME (SECONDS) FIGURE 15.6-1 Inadvertent Opening of a Pressurizer Safety Valve L-15.6-5 w - - - - . . - . . ,e +
2400' 2250 .. ( ' t 5 2000* 2 5 1750-N h u a.
~
5 1500- - " E is m E m 1250 -
- 1000 620 F-600 - - .
W 5 0 580< - .. k u E 560 " .. W U e . v 540 - - " t 520 1 0 10 20 30 4 50 i 60 TlHE(SECONOS) FIGURE 15.6-2 i Inadvertent Opening of a Pressurizer Safety Valve 15.6-6 l 1 _ ___ _ _ - . . ~ , -, ,_-,
TABLE 15.2-1 (Page 6) TIME SEOUENCE OF EVENTS FOR INCIDENTS WHICH CAUSE A DECREASE IN HEAT REMOVAL BY THE SECONDARY SYSTEM Accident Event Time (sec) Loss of Normal Feedwater Main feedwater flow 10 Flow stops Low-low steam generator 63 water level trip Rocs begin to drop 65 Peak wa'ter level in 67 pressurizer occurs - Four steam generators 123
'begin to receive auxiliary feedwater f rom two motor driven auxiliary feedwater_ pumps Core decay heat decreases - 1000 to auxiliary _feedwater har.t removal capa:ity L
l 1 0965T:1/6 '15.2-32 I
C l T 2.4 2.2 - - .. 2.0 - -
+ =
5 1. 8 - - -
- 1. 6 - - ..
- 1. 4 - --
1.2 - 0 1 2 3 4 5 TIME (SECOMOS) FIGURE 15.3.2-4 , l DNBR vs. Time for Four Loops in Operation. Four Pumps Coast'ng Down 15.3-22
i ( l.2
~~T 5
>=
35 1. 0 - (
==
5, .8 ,, W 5
~
.6 .
1.2 .
~~T
= 1.0 - .
h 8 .75- ., 5 u
= .50-- .,
a 3 d .25- -
$.J
- e 0
.25 - '
.40 .
0 2 4 6 8 10 TIME (SECONOS) FIGURE 15.3.3-1 ! Flow Transients for Four l. oops in Operation, One Locked Rotor 15.3-23
~ =
2700 2600 .
- 2500 - - ..
2 2 [ g 2400- - .. N
$ 2300- - --
m - M 2200 . 2100 . 2000 0 2 4 6 8 10 TINE (SECONOS) FIGURE 15.3.3 Peak Reactor Coolant Pressure for Four Loops - in Operation, One Locked Rotor - i ' 15.3-24 i I , _ _ _ . ,, . - , ,-,
( 7 5__ 1.2 1.0 g . 5 ( 5 W= 8
.6--
D I
=
g .4- . a. ( Or y M z
.2- .
O
- 1. 2 7 1.0 ._ ..
5
=5 2.
.g ,
NOT CHANM(L .. ( *g 5
"I ;
W
.6 -
E .4 - p t 4
.s , , .
1.2 2, 1.0 . ,
~
E 5 .8 . AVERAGE CHANNEL
-s ,,
Wg - g ;- .6- - . W 5 .4 - .
.2 ,
0 2 4 6 8 10 TIME (SECONOS) FIGURE 15.3.3-3 Nuclear Power, Average And Hot Channel Hsat Flux Transients for Four Loops 15.3-25 in Operation, One Locked Rotor
l 3000 (
= - 2500 - -
S > 2
$ 2000 - - --
ar - W 5 1500 - -
. 1000 - , --
500 0 2 4 6 8 10 TIME (SECONOS) b FIGURE 15.3.3-4 Maximum C!ad Temperature at Hot Spot for
/
Four Loops in Operation, One Locked Rotor ( 15.3-26. l 1
l l t !. 15.4 REACTIVITY AND POWER DISTRIBUTION ANOMALIES A number of faults have been postulated which could result in reactivity and power distribution anomalies. Reactivity changes could be caused by control C rod motion or ejection, boron concentration changes, or addition of cold water to the Reactor Coolant System. Power distribution changes could be caused by control red motion, misalignment, or ejection, or by static means such as fuel assembly mislocation. These events are discussed in this section. Detailed analyses are presented for the most limiting of these events. Discussions of the following incidents are_ presented in Section 15.4:
- 1. Uncontrolled Rod Cluster Control Assembly Bank Withdrawal from a Suberitical or Low Power Startup Condition
- 2. Uncontrolled Rod Cluster Control Assembly Bank Withdrawal at Power f
( 3. 4. Rod Cluster Control Assembly Misoperation Startup of an Inactive Reactor Coolant Pump at an Incorrect Temperature
- 5. Spectrum.of Red Cluster Control Assembly Ejection Accidents Items 1, 2, 3, and 4 are considered to be ANS Condition II events, and Item 5 an ANS Condition IV event. Item 3 entails both Condition II and III events.
Section 15.0.1 contains a discussion of ANS classifications. 15.4.1 UNCONTROLLED ROD CLUSTER CONTROL ASSEMBLY BANK WITHDRAWAL FROM A SUBCRITICAL OR LOW POWER STARTUP CONDITION 15.4.1.1 Identification of Causes and Accident Description A rod cluster control assembly (RCCA) withdrawal accident is defined as an uncontrolled addition of reactivity to the reactor core caused by withdrawal of RCCA's resulting in a power. excursion. Such a transient could be caused by { l 0952T:6 15.4-1:
a malfunction of the reactor control or rod control systems. This could occur s with the reactor either suberitical, hot zero power or at power. The "at power" case is discussed in Section 15.4.2. Although the reactor is normally brought to power from a subcritical condition by means of RCCA withdrawal,- initial startup procedures with a clean core call for boron dilution. The maximum rate of reactivity increase in t'1e case of boron dilution is less than that assumed in this analysis. The RCCA drive mechanisms are wired into preselected bank configurations whi-h ,, are not altered' during reactor life. These circuits prevent the RCCA's from being automatically withdrawn in other than their respective banks. Power supplied to the banks is controlled such that no more than two banks can be withdrawn at the same time and in their proper withdrawal sequence. The RCCA~ drive mechanisms are of the magnetic latch type and coil actuation is sequenced to provide variable speed trave.l. The maximum reactivity _ insertion rate analyzed in the detailed plant analysis is that occurring with the
- simultaneous withdrawal of the combination of two sequential control banks
[ having the maximum combined worth at maximum speed. This event is classified as an ANS Condition II incident (an incident of mocerate frequency) as defined in Section 15.0.1. The neutron flux response to a continuous reactivity insertion is charac-terized by a very fast rise terminated by the reactivity feedback effect of the negative Doppler coefficient. This self limitation of the power excursion of primary importance since it limits the power to a tolerable level during the delay time for protective action. Should a continuous RCCA withdrawal accident occur, the transient will be terminated by the following automatic 1 features of the Reactor Protection System:
- 1. Source Range High Neutron Flux Reactor Trip - actuated when either of two '
independent source range. channels in,dicates a neutron flux level above a preselected manually adjustable setpoint. This trip function may be manually bypassed only after an intermediate range flux channel indicates 0952T:6 15.4-2 N
a flux level above a specified level. It is automatically reinstated when both intermediate range channels indicate a flux level below a specified
- level.
( 2. Intermediate Ranoe High Neutron Flux Reactor Trio - actuated when either of two independent intermediate range channels indicates a flux level above a preselected manually adjustable setpoint. This trip function may be manually bypassed only after two of the four power range channels are reading above approximately 10 percent of full power ano is automatically reinstated when three of the four channels indicate a power level below this value.
- 3. Power Rance High Neutron Flux Reactor Trip (Low Setting) - actuated when two out of the four power range channels indicate a power level above approximately 25 percent of full power. This trip function maybe manually bypassed when two of the four power range channels indicate a power level above approximately 10 percent of full power and is automatically
( reinstated only after three of the four channels indicate a power level below this value.
- 4. Power Range High Neutron Flux Reactor Trip (High Setting) - actuates when two out of the four power range channels indicate a power level above a-preset setpoint. This trip function is always active, j 5. High Nuclear Flux Rate Reactor Trio - actuated when the positive rate of change of neutron flux on two out of four nuclear power range channels indicate a rate above the preset setpoint. This trip function is always active. ,
In addition, control rod stops on high intermediate range flux level (one of two) and high power range flux level (one out of four) serve to discontinue red withdrawal and prevent the need to actuate the intermediate range ~ flux level trip and the power range flux level trip, respectively. l 0952T:6- 15.4-3 e.
15.4.1.2 Analysis of Effects and Consecuences l Method of Analysis The analysis of the Uncontrolled RCCA bank withdrawal from subcritical accident is performed in three stages: first an average core nuclear power transient calculation, then an average core heat transfer calculation, and finally the DN8R calculation. The average core nuclear calculation is performed using spatial neutron kinetics methods (TWWKLE (Reference 1) to determine the average power generation with time including the various total core feedback cffects, i.e., Doppler reactivity and moderator reactivity. The average heat flux and temperature transients are determined by performing a fuel rod transient heat transfer calculation in FACTRAN (Reference 2). The average heat flux is next used in THINC for transient DNBR calculation. This accident is analyzed using the Improved Thermal Design Procedure as described in WCAP-8567. Plant characteristics and initial conditions are discussed in Section 15.0.3. In order to give conservctive results for a startup accident, the following assumptions are made:
- 1. Since the magnitude of the power peak reached during the initial part of the transient for any given rate of reactivity insertion is strongly dependent on the Doppler coefficient, conservatively low values as a function of power are used. See Section 15.0.4 and Table 15.0.3-2.
- 2. Contribution of the moderator reactivity coefficient is negligible during the initial part of the transient because the heat transfer time between the fuel and the moderator is much longer than the neutron flux response time. However, after the initial neutron flux peak, the succeeding rate of power increase is affected by the moderator reactivity coefficient. A highly conservative value is used in the analysis to yield the maximum peak heat flux.
- 3. The reactor is assumed to be at hot zero power. This assumption is more conservative than that of a lower initial system temperature. The higher 0952T:6 15.4-4
initial system temperature yields a large fuel-water heat transfer + coefficient, larger specific heats, and a less negative (smaller absolute magnitude) Doppler coefficient, all of which tend to reduce the Doppler feedback effect thereby increasing the neutron flux peak. The initial effective multiplication factor is assumed to be 1.0 since this results in the worst nuclear power transient.
- 4. Reactor trip is assumed to be initiated by power range high neutron flux (low setting). The most adverse combination of instrument and setpoint errors, as well as delays for trip signal actuation and rod cluster control assembly release, is taken into account. A 10 percent increase is assumed for the power range flux trip setpoint raising it from the nominal value of 25 percent to 35 percent. Since the rise in the neutron flux is so rapid, the effect of errors in the trip setpoint on the actual time at which the rods are released is negligible. In addh ion, the reactor trip insertion characteristic is based on the assumption . hat the highest worth ,
rod cluster control assembly is stuck in its fully wi;hdrawn position. ( See Section 15.0.5 for rod cluster control assembly i sertion characteristics.
- 5. The maximum positive reactivity insertion rate ass'med is greater than that for the simultaneous withdrawal of the combination of the two sequential control banks having the greatest c',mbined worth at maximum speed (45 inches / minute).
- 6. The most limiting axial and radial power shapes, associated with having the two highest combined worth banks in their high worth position, is assumed in the DNB analysis.
l
- 7. The initial power level was assumed to be below the power level expected for any shutdown condition (10 cf nominal power). This combination of
/ highest reactivity insertion rate and lowest initial power produces the highest peak heat flux.
- 8. Two reactor coolant pumps are assumed to be in operation. This is
! conservative with respect to DNB. 0952T:6 -15.4-5
i l 1 Results Figures 15.4.1-1 through 15.4.1-3 show the transient behavior for the uncontrolled RCCA bank withdrawal incident, with the accident terminated by reactor trip at 35 percent of nominal power. The reactivity insertion rate used is greater than that calculated for the two highest worth sequential contro,1 banks, both assumed to be in their highest incremental worth region. Figure 15.4.1-1 shows the neutron flux transient. ~ The energy release and the fuel temperature increases are relatively small. The thermal flux response, of interest for DNB considerations, is shown on Figure 15.4.1-2. The beneficial effect of the inherent thermal lag in the fuel is evidenced by a peak heat flux much 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 15.4.1-3 shows the response of the average fuel and cladding temperature. The average fuel temperature increases to a value lower than the nominal full power value. The minimum DNBR at all times remains above the limit value. The calculated sequence of events for this accident is shown on Table 15.4-1. 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 Conclusions In the event of a RCCA withdrawal accident from the subcritical' condition, the core and the Reactor Coolant System are not adversely affected, since the corbination of thermal power and the coolant temperature result in a DNBR greater _than the limit value. Thus, no fuel or clad damage is predicted as a result of DNB. f 0952T:6 _15.4-6
15.4.2 UNCONTROLLED ROD CLUSTER CONTROL ASSEMBLY BANK WITHDRAWAL AT POWIR 15.4.2.1 Identification of Causes and Accident Descriotion ( Uncontrolled rod cluster control assembly -(RCCA) bank withdrawal at power results in an increase in the core heat flux. Since the heat extraction from the steam generator lags behind the core power generation until the steam generator pressure reaches the relief or safety valve setpoint, there is a ret increase in the reactor coolant temperature. 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 protection System is designed to terminate any such transient before the DNBR falls below the limit value. This event is classified as an ANS Condition II incident (an incident of moderate frequency) as defined in Section 15.0.1. ( The automatic features of the Reactor Protection System which prevent core damage following the postulated accident include the following:
- 1. Power range neutron flux instrumentation actuates a reactor trip if two-of-four channels exceed an overpower setpoint.
- 2. Reactor trip is actuated if any two-out-of-four AT channels exceed an overtemperature AT setpoint. This setpoint is automatically varied with axial power imbalance, coolant temperature and pressure to protect against DNB.
- 3. Reactor trip is actuated if any two-out-of-four. AT channels exceed an overpower AT setpoint. This setpoint is automatically varied with axial power imbalance to ensure that_the allowcsie heat generation rate (kw/ft) is not exceeded.
- 4. A high pressurizer _ pressure reactor trip actuated from-any two-out-of-four pressure channels wnich is_ set at a fixed point. This set pressure is less than the set pressure for the pressurizer safety valves.
0952T:6 15.4-7
~
i
- 5. A high pressurizer water level reactor trip actuated from any two-out-of-three level channels when the reactor power is above approximately 10 per-cent (Permissive 7).
1 In addition to the above listed reactor trips, there are the following RCCA withdrawal blocks:
- 1. High neutron flux (one-out-of-foer power range)
- 2. Overpower AT (two-out-of-four) f
- 3. Overtemperature AT (two-out-of-four)
The manner in which the combination of overpower and overtemperature AT trips provide protection over the full range of Reactor Coolant System conditions is described in FSAR Chapter 7,. Figure 15.0.3-1 presents allowable reactor coolant loop average temperature and AT for the design power distribution and flow as a function of primary coolant pressure. The boundaries of operation defined by the overpower AT trip and the overtemperature AT trip are represented as " protection lines" on 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 a given DNBR can be represented as a line. The DNB lines represent the locus of conditions for which the DNBR equals'the limit value (1.47 for the thimble cell and 1.49 for the typical cell). All points below and to the left-of a DNB line for a given pressure have a DNBR greater than the_ limit.value. The diagram shows that DNB is prevented for all cases if the area enclosed with the maximum protection i lines is not traversed by the applicable DNBR line at any' point. The area of permissible operation (power, pressure and temperature) is bounded by the combination of reactor trips: high neutron flux (fixed setpoint); high pressure (fixed setpoint); low pressure (fixed setpoint); overpower and over-temperature AT (variable setpoints). 1 i 0952T:6- 15.4-8
15.4.2.2 Analysis of effects and Consecuences Method of Analysis The transient is analyzed by the LOFTRAN Code (Reference 3). This ccde simulates the neutron kinetics, Reactor Coolant System, pressurizer, pressurizer. relief and safety valves, pressurizer spray, steam generator, and steam generator safety valves. The code computes pertinent plant variables including temperature, pressures, and power level. The core limits as illustrated in Figure 15.0.3-1 are used as input to LOFTRAN to determine the minimum DNBR during the transient. This accident is analyzed with the Improved Thermal Design Procedure as described in WCAP-8567. Plant characteristics and initial conditions are discussed in Section 15.0.3.
- 1. Initial reactor power, pressure, and RCS temperatures are assumed to be at
( their nominal valoes. Uncertainties in initial conditions are included in the limit DNBR as described in WCAP-8567.
- 2. Reactivity Coefficients - Two cases are analyzed:
- a. Minimum Reactivity Feedback. A least negative moderator coefficient of reactivity is assumed corresponding to the beginning of core life.
A variable Doppler power coefficient with core power is used in the analysis. A conservatively small (in absolute magnitude) value is assumed.
- b. Maximum Reactivity Feedback. A conservatively large positive moderator density coefficient and a large (in absolute magnitude) negative Doppler power coefficient are assumed.
L 0952T:6 15.4-9
. ... ____--- J
l 1 I
- 3. The reactor trip on high neutron flux is assumed to be actuated at a conservative value of 118 percent of nominal full power. The AT trips include all adverse instrumentation and setpoint errors; the delays for trip actuation are assumed to be the maximum values.
- 4. The RCCA trip insertion characteristic is based on the assumption that the highest worth assembly is stuck in its fully withdrawn position.
- 5. The maximum positive reactivity insertion rate is greater than that for the simultaneous withdrawal of the two control banks having the maximum combined worth at maximum speed.
The effect of RCCA movement on the axial core power distribution is accounted for by causing a decrease in overtemperature AT trip'setpoint proportional to the decrease in margin to DNB. No single active failure in any plant sys em will adversely affect the consequences of the accident. Results Figures 15.4.2-1 through 15.4.2-3 show the transient response for a rapid RCCA withdrawal incident starting from full power. Reactor trip on high neutron flux occurs shortly after the start of the accident. Since this is rapid witn respect to the thermal time constants of the plant, small changes in T,yg and pressure result and margin to DNB is maintained. The transient response for a slow RCCA withdrawal from full power is shown in Figures 15.4.2-4 through 15.4.2-6. Reactor trip on overtemperature AT I occurs after a longer period and the rise in temperature and pressure is consequently larger than for rapid RCCA withdrawal. Again, the minimum DNBR is greater than the limit value. Figure 15,4.2-7 shows the minimum DNBR as a function of reactivity insertion rate from initial full power operation for minimum and maximum reactivity 0952T:6 15.4-10
. , - . . . . . - - - - . . - - - - . . . ~ . - - - - . . . . . . ~-- - -
l
~
I l feedback. It can be seen that two reactor trip channels provide protection over the whole range of reactivity. insertion rates. These are the high , neutrun flux and overtemperature AT channels. The minimum DNBR is never j less than the limit value. t l 1 Figures 15.4.2-8 and 15.4.2-9 shows 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, except as the initial power is decreased, the range over which the t i overtemperature AT trip is effective is increased. In neither case does the j DNBR fall below the limit value. The shape of the curves of minimum DNB ratio versus reactivity insertion rate in the reference figures is due both to reactor core and coolant system s transient response and to protection system action in initiating a reactor ) trip. 1 I Referring to Figure 15.4.2-8, for example, it is noted that:
- 1. For high reactivity insertion rates (i.e., above - 2 x 10~4 oK/sec) reactor trip is initiated by the high neutron flux trip for the minimum reactivity feedback cases. 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 j system fluid. Thus, the reactor is tripped prior to significant.incresse in heat flux or water temperature with resultant high minimum DNB ratios
~
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 DNB ratio during the transient thus decreases with decreasing insertion rate. I
. 2. The overtemperature AT reactor trip' circuit initiates a reactor trip when measured coolant loop'AT exceeds a setpoint based on measured Reactor Coolant System average temperature and pressure. '.It is important i ~
to note that the average temperature centribution to the circuit is
'0952T:6 115.4-11 1
-i i
..-,i . ._, -- - ._ - , _ , , . . .-. ,
1
; lead-lag compensated in order to decrease the effect of the thermal capacity of th6 Reactor Coolant System in response to power increases.
- 3. With further decrease in reactivity insertion rate, the overtemperature AT and high neutron flux trips become equally effective in terminating the transient (e.g. , at - 2 x 10 6K/see reactivity insertion rate).
For reactivity insertion rates between - 2 x 10 ~4 6K/see and - 6 y 10 6K/see the effectiveness of the overtemperature AT trip increases (in terms of increased minimum DNB ratio) due to the fact that J with lower insertion rates the power increase rate is slower, the rate of rise of average coolant temperature is slower and the system lags and delays become less significant. ~
- 4. For reactivity insertion rates less than - 6 x 10-5 6K/sec, the rise in the reactor coolant temperature is sufficiently high so that the steam generator safety valve setpoint is reached prior to trip. Opening of these valves, which act as an additional heat load on the Reactor Coolant System, sharply decreases the rate of increase of Reactor Coolant System average temperature. This decrease in rate of increase of the average coolant system temperature during the transient is accentuated by the lead-lag compensation causing the overtemperature AT trip setpoint to be
; reached later with resulting lower minimum DNB ratios.
Figures 15.4.2-7, 15.4.2-8, and 15.4.2-9 illustrate minimum DNBR calculated for minimum and maximum reactivity feedback. Since the RCCA withdrawal at power incident is an overpower transient, the. fuel temperatures rise during the transient until after reactor trip occurs. For high reactivity insertion rates, the overpower transient is fast with respect to the fuel rod thermal time constant, and the core heat flux lags behind the neutron flux' response. Due-to this lag, the peak core heat flux does not exceed 118 percent of its nominal value (i.e., the high neutron flux trip setpoint assumed in the analysis). Taking into account'the effect of the RCCA withdrawal on~ the axial core power distribution. .the peak fuel centerline ' ~ temperature will still remain below.the fuel melting temperature. - 0952T:6 15.4-12
For slow reactivity insertion rates, the core heat flux remains more nearly in equilibrium with the neutron flux. The overpower transient is terminated by the overtemperature AT reactor trip before a DNB condition is reached. The peak heat flux again is maintained below 118 percent of its nominal value. Taking into account the effect of the RCCA withdrawal on the axial core power distribution, the peak fuel centerline temperature will remain below the fuel melting temperature. Since DNB does not occur at any time during the RCCA withdrawal at power transient, the ability of the primary coolant to remove heat from the fuel rod is not reduced. Thus, the fuel cladding temperature does not rise significantly above its initial value during the transient. The calculated secuerce of events for this accident is shown on Table 15.4-1. With the reactor tripped, the plant eventually returns to a stable condition. The plant may subseqJently be Cooled down further by following normal plant shutdown procedures. 15.4.2.3 Conclusions The high neutron flux and overtemperatur6 AT trip channels provide adequate protection over the entire range of possible reactivity insertion rates, i.e., the minimum value of DNBR is always larger than the limit value. 15.4.3. ROD CLUSTER CONTROL ASSEMBLY MISOPERATION 15.4.3.1 Identification of Causes and Accident Description
~
Rod cluster control assembly misoperation accidents include: A. One er more dropped RCCAs withir the same group B. A dropped RCCA bank C. Statically misa11gned RCCA i 0952T:6 15.4-13 n - ,
G. Withdrawal of a single RCCA. Each RCCA has a 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. Full length RCCAs are always moved in preselected banks, and the banks are
- always moved in the same preselected sequence. Each bank of RCCAs is divided into two groups. The rods comprising a group operate in parallel through multiplexing thyristors. The two groups in a bank move sequentially such that the first group is always within one step of the second group in the bank. A d& finite 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 mechanism. Since the stationary gripper, movable gripper, and lift coils associated with the four RCCAs of a rod group are driven in parallel, any single failure which would cause rod withdrawal would affect a minimum of one group. Mechanical failures are in tne direction of insertion, or immobility.
i The dropped RCCA, dropped RCCA bank, and statically misaligned RCCA events are classified as ANS Condition II incidents (incidents of moderate frequency) as defined in subsection 15.0.1. However, the single RCCA withdrawal incident is classified as an ANS Condition III event, as discussed below. Nn single electrical or mechanical failure in the rod control system could cNasetheaccidentalwithdrawalofasingleRCCAfromthe-insertedbankat fu'il power operation. The operator could withdraw a single RCCA in the con-trol bank since this feature is necessary in order to retrieve an assembly should one be accidentally dropped. The event analyzed must result from multiple wiring failures or multiple serious operator errors, and subsequent - k and repeated operator disregard of event indication. The probability of such a combination of conditions is so low that the limiting consequences may
~ , include slight fuel damage.
0952T:6 15.4-14 f
Thus, coasistent with the philosophy and format of ANSI N18.2, the event is classified as a Condition III event. By definition, " Condition III occurrences include incidents, any one of which may occur during the lifetime of a particular plant", and "shall not cause more than a small fraction of C fuel elements in the reactor to be damaged ...". This selection of criterion is in accordance with General Design Criterion (GDC) 25 which states, "The protection system shall be designed to assure that specified acceptable fuel design limits are not exceeded for any sinole mal-function of the reactivity control systems, such as accidental withdrawal (not ejection or dropout) of control rods." (Emphases have been added). It has been shown that single failures resulting in RCCA bank withdrawals do not violate specified fuel design limits. Moreover, no single malfunction can result in the withdrawal of a single RCCA. Thus, it is concluded that criterion established for the single rod withdrawal at power is appropriate and in accordance with GDC 25. - ( A dropped RCCA or RCCA bank is detected by: A. Sudden drop in the core power level as seen by the nuclear instrumen-tation system B. Asymmetric power distribution as seen on out-of-core neutron detectors or core exit thermocouples C. Rod at bottom signal D. Rod deviation alarm ,
- E. Rod position indication.
Misaligned RCCAs are detected by: A. Asymmetric power distribution as seen on out-of-core neutron detectors or core exit thermocouples l l C - i 0952T:6 15.4-15 i
B. Rod deviation alarm C. Rod position indicators The resolution of the rod position indicator channel is +5 percent of span
.(+7.2 inches). Deviation of any RCCA from its group by twice this distance (10 percent of span, or 14.4 inches) will not cause power distributions worse than the design limits. The deviation alarm alerts the operator to rod devia-tion with respect to the group position in excess of 5 percent of span. If the rod deviation alarm is not operable, the operator is required to take action as required by the Technical Specifications.
If one or more rod position indicator channels should be out of service, detailed operating instructions shall be followed to assure the alignment of the non-indicated RCCA. The operator is also required to take action as required by the Technical Specifications.. In the extremely unlikely event of simultaneous electrical failures which could result in single RCCA withdrawal, rod deviation and rod control failure would both be displayed on the plant annunciator, and the red position indica-tors would indicate the relative positions of the assemblies in the bank. The urgent failure alarm also inhibits automatic rod motion in the group in which it occurs. Withdrawal of a single RCCA by operator action, whether deliberate or by a combination of errors, would result in activation of the same alarm and the same visual indications. Withdrawal of a single RCCA results in both positive reactivity insertion tending to increase core power, and an increase in local power censity in the core area associated with the RCCA. Automatic protection for this event is provided by the overtemperature AT reactor trip; although, due to the increase in local power density, it is not possible in all cases to provide assurance that the core safety limits will not be violated. 15.4.3.2 3 2alysis of Effects and Consecuences 15.4.3.2.1 Dropped RCCAs., Dropped'.RCCA Bank, and Statically Misaligned RCCA' 0952T:6 15.4-16
15.4.3.2.1.1 Method of Analysis A. One or more dropped RCCA from the same group For evaluation of the dropped RCCA event, the transient system response is calculated using the LOFTRAN(3) code. The code simulates the neutron kinetics, Reactor Coolant System, pres-surizer, pressurizer relief and safety valves, pressurizer spray, steam generator, and steam generator safety valves. The code computes pertinent plant variables including temperatures, pressures, and power level. Statepoints are calculated and nuclear models are used to obtain a hot channel factor consistent with the primary system conditions and reactor power. By incorporating the primary conditions from the tran-sient and the hot chacnel factor from the nuclear analysis, the DNB design basis is shown to be met using the THINC code. The transient C response, nuclear peaking factor analysis, and DNB design basis con-firmation are performed in accordance with the methodology described in Reference 4. B. Statically Misaligned RCCA Steady state power distributions are analyzed using the computer codes as described in the McGuire FSAR Section 4.4. The peaking factors are then used as input to the THINC code to calculate the DNBR. 15.4.3.2.1.2 Results A. One or more dropped RCCAs (
\. Single or multiple RCCAs within the same group result in a negative reactivity insertion which may be detected by the power range negative neutron flux rate trip circuitry. If detected, the reactor is tripped C
0952T:6 15.4-17
within approximately 2.5 seconds following the drop of the RCCAs. The core is not adversely affected during this period, since power is decreasing rapidly. Following reactor trip, normal shutdown proce-dures are followed. The operator may manually retrieve the RCCA by following approved operating procedures. For those dropped RCCAs which co not result in a reactor. trip, power may be reestablished either by reactivity feedback or control bank withdrawal. Following a dropped red event in manual rod control, the plant will establish a new equilibrium condition. The equilibrium process without control system interaction is n.onotonic, thus removing power overshoot as a concero, and establishing the automatic rod con-trol mode of operation as the limiting case. For a dropped RCCA event in the automatic rod control mode, the Rod Control System detects the drop io power and initiates control bank withdrawal. Power overshoot may occur due to this action by the auto-matic red controller after which the control system will insert the control bank in order to restore noninal power. Figures 15.4.3-1 and 15.4.3-2 show a typical transient response to a dropped RCCA (or RCCAs) in automatic control. In all cases the minimum DNBR remains above the limit value. B. Dropped RCCA Bank A dropped RCCA bank typically results in a reactivity insertion greater than 500 pcm which will be detected by the power range nega- . tive neutron flux rate trip circuitry. The reactor is tripped within approximately 2.5 seconds following the drop of an RCCA. The core is not adversely affected during this period, since the power is decreas-
- ing rapidly. Following reactor trip, normal shutdown procedures may subsequently be followed to further cool down the plant.
t 0952T:6 15.4-18
l 1 C. Statically Misaligned RCCA The most severe misalignment situations with respect to DNBR at signi-( ficant power levels arise from cases in which one RCCA is fully inser-ted, or where bank D is fully inserted with one RCCA fully withdrawn. Multiple independent alarms, including a bank insertion limit alarm, alert the operator well before the postulated conditions are approached. The bank can be inserted to its insertion limit with any one assembly fully withdrawn without the DNBR falling below the limit value. Any action required of the operator to maintain the plant in a stabilized condition will be in a time frame in excess of ten minutes following the incident. The insertion limits in the Technicai Specifications may vary from time to time depending on a number of limiting criteria. It is pre-ferable,'therefore, to analyze the misaligned RCCA case at full power - for position of the control bank as deeply inserted as the criteria on minimum DNBR and power peaking factor will allow. The full power C insertion limits on control bank D must then be chosen to be above that positior. and will usually be dictated by other criteria. Detailed results will vary from cycle to cycle depending on fuel arrangements. . For this RCCA misalignment, with bank D inserted to its full power insertion limit and one RCCA fully withdr2wn, DNBR does not fall below the limit value. This case is analyzed assuming the initial reactor ( power, pres:ure, and RCS temperature are at their nominal values as given in Table 15.0.3-3, but with the increased radial peaking. factor associated with the misaligned RCCA.
' DNB calculations have not been performed specifically for RCCAs mis-sing from other banks; however, power shape calculations have been performed as required for the RCCA ejection analysis. Inspection of the power shapes shows that the DNB and peak kw/ft situation is less severe than the bank D case discussed above assuming insertion limits on the other banks equivalent to a b'ank D full-in. insertion limit.
0952T:6 .15.4-19 t-
For RCCA misalignments with one RCCA fully inserted, the DNSR d c not fall below the limit value. This case is analyzed assuming the ini-tial reactor power, pressure, and RCS temperatures are at their nomi-nal values as given in Table 15.0.3-3 but with the increased radial peaking factor associated with the misaligned RCCA. { A DNB does not occur for the RCCA misalignment incident and thus the ability of the primary coolant to remove heat from the fuel rod is not reduced. The' peak fuel temperature corresponds to a linear heat gene-ration rate based on the radial peaking factor penalty associated with the misaligned RCCA and the design axial power distribution. The resulting linear heat generation is well below that which would cause fuel melting. Following the identification of a RCCA misalignment condition by the opera-tor, the operator is required to take action as required by the plant Tech-nical Specifications and operating instructions. 15.4.3.2.2 Single RCCA Withdrawal 15.4.3.2.2.1 Method of Analysis. Power distributions within the core are calculated by the computer codes as described in FSAR Table 4.1-2. The peak-ing factors are then used by THINC to calculate the minimum DNBR for the event. The case of the worst rod withdrawn from bank D inserted at the insertion limit, with the reactor initially at full power, was analyzed. This incident is assumed to occur at beginning-of-life since this results in the minimum value of moderator temperature. coefficient. This assumption maximizes the power rise and minimizes the tendency of increased moderator temperature - to flatten the power distribution. 15.4.3.2.2.2 Results of the Analy g . For the single rod withdrawal event, two cases have been considered as follows: i 0952T:6 15.4-20
l A. If the reactor is in the manual control mode, continuous withdrawal of a single RCCA results in both an increase in core power and coolant temperature, and an increase in the local hot channel factor in the ( area of the withdrawing RCCA. In terms of the overall system response, this case is similar to those presented in subsection 15.4.2; however, the increased local power peaking in the area of the withdrawn RCCA results in lower minimum DNBRs than for the withdrawn bank cases. Depending on initial bank insertion and location of the withdrawn RCCA, automatic reactor trip may not occur sufficiently fast to prevent the minimum core DNBR from falling below the limit value. Evaluation of this case at the power and coolant conditions at which the overtemperature AT trip would be expected to trip the plant shows that an upper limit for the number of rods with a DNBR less than the limit value is 5 percent. B. If the reactor is in the automatic control mode, the multiple failures that rescit in the withdrawal of a single RCCA will result in the ( immobility of the other RCCAs in the controlling bank. The transient will then proceed in the same manner as case A described above. For such cases as above, a reactor trip will ultimately ensue, although not sufficiently fast in all cases as to prevent a minimum DNB ratio in the corr of less than the limit value. Following reactor trip, normal shutdown procedures may be followed. 15.4.3.3 Conclusions For cases of dropped RCCAs or dropped banks, for which the reactor is tripped by the power range negative neutron flux rate trip, there is no_ reduction in the margin to core thermal limits, and consequently the DNB design basis is met. It is shown for all cases which do not result in reactor trip that the - DNBR remains greater than the limit value and, therefore, the DNB design is met. L 0952T:6 15.4-21 y 'm=
l For all cases of any RCCA fully inserted, or bank D inserted to its rod inser-tion limits with any single RCCA in that bank fully withdrawn (static mis-l alignment), the DNBR remains greater than the limit value. For the case of the accidental withdrawal of a single RCCA, with the reactor l in the automatic or manual control mode and initially operating at full power j with bank D at the insertion limit, an upper bound of the number of fuel rods- ! experiencing DNB is 5 percent of the total fuel rods in the core. 15.4.4 STARTUP OF AN INACTIVE REACTOR C001. ANT PUMP AT AN INCORRECT TEMPERATURE 15.4.4.1 Identification of Causes and Accident Descriotion If the plant is operating with one pump out of service, there is reverse flow through the inactive loop due to the pressure difference across the reactor vessel. The cold leg temperature in an inactive loop is identical to the cold leg temperature of the active loops (the reactor core inlet temperature). If the reactor is operated at power, and assuming the secondary side of the steam generator in the inactive loop is not isolated, there is a temperature drop across the steam generator in the inactive loop and, with the reverse flow, the hot leg temperature of the inactive loop is lower than the reactor core inlet temperature. Starting of an idle reactor coolant pump without bringing the inactive loop hot leg temaerature close to the core. inlet temperature would result in the injection of cold water into the core, which would cause a reactivity insertion and subseouent power increase. - This event is classified as an ANS Condition II incident (an incident of moderate frequency) as defined in Section 15.0.1. . t l Should the startup of an inactive reactor coolant pump accident occur, the transient will be terminated automatically by a reactor trip on low coolant loop flow when the power range neutron flux (two out of four channels) exceeds j the P-8 setpoint, which has been previously reset for three' loop operation. i l i 0952T:6 15.0-22
, 15.4.4.2 Analysis of Effects and Consecuences
- l Method of Analysis This transient is analyzed by three digital computer codes. The LOFTRAN Code (Reference 3) is used to calculate the loop and core flow, nuclear power and core press.ure and temperature . transients following the startup of an idle pump. FACTRAN (Reference 2) is used to calculate the core heat flux transient based on core flow and nuclear power from LOFTRAN. The THINC Code is then used to calculate the DNBR during the transient based on system conditions (pressure, temperature, and flow) calculated by LOFTRAN and heat flux as '
calculated by FACTRAN, This accident is analyzed with the Improved Thermal Design Procedure as described in WCAP-8567. Plant characteristics and initial conditions are discussed in Section 15.0.3. _ ( In order to obtain conservative results for the startup of an inactive pump accident, the following assumptions are made:
- 1. Initial reactor power, pressure, and RCS temperatures are assumed to be at their nominal N-1 loop operation values. Uncertainties in initial conditions are included in the limit DNBR as described in WCAP-8567.
- 2. Following initiation of startup of the idle pump, the inactive loop flow reverses and accelerates to its nominal flow value in approximately 20-seconds. This value is faster than the expected startup time, and is conservative for this analysis.
- 3. A conservatively large moderator density coefficient.
-1
- 4. A. conservatively small (absolute value) negative Doppler power. coefficient. ,
- 5. The initial reactor coolant loop flows are at the _ appropriate values for one pump out of service.
0952T:6 15.4-23 A
i l 4
- 6. The reactor trip is assumed to occur on low coolant flow when the power range neutron flux exceeds the P-8 setpoint. The P-8 setpoint is conser-vatively assumed to be 84 percent of rated po,ver which corresponds to the nominal setpoint plus 9 percent for nuclear instrumentation trrors.
Results The results following the startup of an idle pump with the above listed assumptions are shown in Figures 15.4.4-1 through 15.4.4-5. As snown in these l curves, during the first part of the transient, the increase in core flow with j cooler water results in an increase in nuclear power and a decrease in core
- average temperature. The minimum DNBR during the transient is considerably l greater than the limit value.
Reactivity addition for the inactive loop startup accident is due to the decrease in core water temperature. Duri.ng the transient, this decrease is due both to the increase in reactor coolant flow and, as the inactive loop flow reverses, to the colder water entering the core from the hot leg side (colder temperature side prior to the start of the transient) of the steam generator in the inactive loop. .Thus, the reactivity insertion rate for this-i transient changes with titre. The resultant core nuclear power transienti, computed with consideration of both moderator and; Doppler reactivity feedback effects, is shown in Figure 15.4.4-1. i The calculated sequence of events for this accident is shown in Table 15.4-1. The transient results illustrated in. Figures 15.4.4-1 through 15.4.4-5 indicate that a stabilized plant condition, with the reactor tripped, is . approached rapidly. Plant cooldown may subsequently be achieved by following ' normal shutdown procedures. 15.4.4.3 Conclusions The transient results show that the core is not adversely affected. There is considerable margin to the limit DNBR value; thus, no fuel or clad damage is predicted. 0952T:6: - 15.'4-24 : ~
-7 .w+ - , ~ y
15.4.5 SPECTRUM OF RCD CLUSTER CONTROL ASSEMBLY EJECTION ACCIDENTS 15.4.5.1 Identification of Causes and Accident Descriotion This accident is defined as the mechanici.1 failure of a control rod mechanism pressure housing resulting in the ejection of a rod cluster control assembly (RCCA) and drive shaft. The consequence of this mechanical failure is a rapid positive reactivity insertion together with an adverse core power distribu-tion, possibly leading to localized fuel rod damage. 15.4.5.1.1 Dasign Pracautioris and Protection Certain features are intended to preclude the possibility of the rod ejection accident, or to limit the consequances if the accident were to occur. .These include a sound, conservative mechanical design of the rod housings, together with a thorough quality control (testing) program during asserably, and a nuclear design which lessens the potential ejection worth of RCCAs, ana ( minimizes the number of assemblies inserted at high power levels. Mechanical Desion Mechanical design and quality control procedures intended to preclude the possibility nf a RCCA drive mechanism housing failure are listed below:
- 1. Each full-length control rod drive mechanism housing is completely assembled and shop tested at 4100 psi.
2. The meenanism housings are individually hydrotested after they are attached to the head adapters in the reactor vessel head, and checked during the hydrotest of the completed reactor coolant system. C 0952T:6 15.4-25
- 3. Stress levels in the mechanism are not affected by anticipated system transients at power, or by the thermal movement of the coolant loops.
Moments induced by the design-basis earthquake can be accepted within the allowable primary working stress range specified by the ASME Code, Section III, for Class I components.
- 4. The latch mechanism housing and rod travel housing are each a single length of forged Type 304 stainless steel. This material exhibits excellent notch toughness at all temperatures which will be encountered.
A significant margin of strengtn in the elastic range together with the large energy absorption capability.in the plastic range gives additional assurance that gross failure cf the housing will not occur. The joints between the latch mechanism housing and head adapter, and between the latch mechanism housing and rod travel housing, are threaded joints reinforced by canopy type rod welds. Administrative regulations re. quire periodic inspections of these (and other) welds. Nuclear Design Even if a rupture of a RCCA drive mechanism housing is postulated, the operation of a plant utilizing chemical shim is such that the severity of an ejected RCCA is inherently limited. In general, the reactor is operated with the RCCA's inserted only far enough to permit load' follow. Reactivity changes caused by core depletion and xenon transients are compensated by boron changes. Further, the location and grouping of control RCCA banks are selected during the nuclear design to lessen the severity of a RCCA ejection accident. Therefore, should a RCCA be ejected from its normal position during full power operation, only a minor reactivity excursion, at worst, could be expected to occur. However, it may be occasionally desirable to operate with larger than normal : insertions. For this reason, a rod insertion limit is defined as a function of power level. Operation with the RCCAs abcVe this limit guarantees
-0952T:6 15.4-26
acequate shutdown capability and acceptable power distribution. The position of all RCCAs is continuously indicated in the control room. An alarm will occur if a bank of RCCAs aoproaches its insertion limit or if one RCCA ( deviates from its bank. Operating instructions require boration at low level alarm and emergency boration at the low-low alarm. Reactor Protection The reactor protection in the event of a rod ejection accident has been described in Reference 5. The protection for this accident is provided by high neutron flux trip (high and low setting) and high rate of neutron flux increase trip. Effects on Adjacent Housinos Disregarding the remote possibility of the occurrence of,a RCCA mechanism housing failure, investigations have shown that failure of a housing due to ( either longitudinal or circumferential cracking would not cause damage to adjacent housings. However, even if damage is postulated, it would not be expected to lead to a more severe transient since RCCAs are inserted in the core in symmetric patterns, and control rods immediately adjacent to worst ejected rods are not in the core when the reactor is critical. Damage to an adjacent housing could, at worst, cause that RCCA not to fall on receiving a trip signal; however, this is already taken into account in the analysis by assuming a stuck red adjacent to the ejected rod. 15.4.5.1.2 Limiting Criteria This event is classified as an ANS Condition IV incident. See Section 15.0.1 for a discussion of ANS classifications. Due to the extremely low probability of a RCCA ejection accident, some fuel damage could be considered an - acceptable consequence. l 1 l 0952T:6 15.4-27 ' l
Comprehensive studies of the threshold of fuel failure and of the threshold or significant conversion of the fuel thermal energy to mechanical energy, have been carried out as part of the SPERT project by the Idaho Nuclear Corporation (Reference C). Extensive tests of zirconium clad U02 I"'I " d5 "*P'"**~ tive of those in Pressurized Water Reactor type cores have demonstrated failure thresholds in the range of 240 to 257 cal /gm. However, other rods of a ( sightly different design have exhibited failures as low as 225 cal /gm. These results differ significantly from the TREAT (Reference 7) results, which ir.dicated a failure threshold of 280 cal /gm. Limited results have indicated that this threshold decreases by about 10 percent with fuel burnup. The clad failure mechanism appears to be melting for zero burnup rods and brittic fracture for irradiated reds. Also important is the conversion ratio of thermal to mechanical energy. This ratio becomes marginally d?tectable above 300 cal /gm for unirradiated rods and 200 cal /gm for irradiated rods; catastrophic failure, (large fuel dispersal, large pressure rise) event for irradiated rods, did not occur below 300. cal /gm. In view of-the above experimental results, criteria are applied to ensure that there is little or no possibility of fuel dispersal in the coolant, gross lattice distortion, or severe shock waves. These criteria are:
- 1. Average fuel pellet enthalpy at the hot spot below 225 cal /gm for unirradiated fuel and 200 cal /sm for irradiated fuel.
- 2. Average clad temperature at the hot spot below the temperature at which clad embrittlement may be expected (2700*F).
- 3. Peak reactor coolant pressure less than that which could cause stresses to exceed the faulted condition stress limits.
- 4. Fuel melting will be limited to less than ten percent of the fuel volume
- at the hot spot even if the average fuel pellet enthalpy is below the limits of criterion 1 above. j 0952T:6 15.4-28
15.4.5.2 Analysis of Effects and Consecuences Method of Analysis ( The calculation of the RCCA ejection transient is performed in two stages, first an average core channel calculation and then a hot region calculation. The average core calculation is performed using spatial neutron kinetics methods to determine the average power generation with time including the various total core feedback effects, i.e., Doppler reactivity and moderator reactivity. Enthalpy and temperature transients in the hot spot are then determined by multiplying the average core energy generation by the hot channel factor and performing a fuel rod transient heat transfer calculation. The power distribution calculated without feedback is pessimistically assumed to persist throughout the transient. A detailed discussion of the method of analysis can be found in Reference [5]. ( Average Core Analysis Thi spatial kinetics computer code, TWINKLE (Reference 1), is used for the average core transient analysis. This code solves the two group neutron dift!sion theory kinetic equation in one, two or three spatial dimensions (rectangular coordinates) for six delayed neutron groups and up to 2000 spatial points. The computer code includes a detailed multiregion, transient fuel-clad-coolant heat transfer model for calculation of pointwise Doppler and moderator feedback effects. In this analysis, the code is used as a one dimensional axial kinetics code since it allows a more realistic representa-tion of the spatial effects of axial moderator feedback and RCCA movement. However, since the radial dimension is missing, it is still necessary to employ very conservative methods (described in the following) of calculating the ejected rod worth and het channel factor. . Further description of TWINKLE appears in Section 15.0.7. G . 0952T:6 15.4-29 Y r 2
Hot Spot Analysis In the hot spot analysis, the initial heat flux is equal to the nominal times the design hot channel factor. During the transient, the heat flux hot channel factor is linearly increased to the transient value in 0.1 second, the time for full ejection of the rod. Therefore, the assumption is made that the hot spot before and after ejection are coincident. This is very conservative since the peak after ejection will occur in or adjacent to the assembly with the ejected rod, and prior to ejection the power in this region will neces-sarily be depressed. The hot spot analysis is performed using the detailed fuel and cladding transient heat transfer computer code, FACTRAN (Reference 2). This ccmputer code calculates the transient temperature distribution in a cross section of a netal clad UO 2 fuel rc,d, and the heat flux at the surface of the rod, using as input the nuclear power versus time and the lecal coolant conditions. The zirconium water reaction is explicitly represented, and all material proper-ties are represented as functions of temperature. A conservative pellet radial power distribution is used within the fuel rod. FACTRAN uses the Dittus-Boelter or Jens-Lottes correlation to determine the film heat transfer before DNB, and the Bishop-Sandberg-Tong correlation (Reference 8) to determine the film boiling coefficient after DNB. The BST correlation is conservatively used assuming zero bulk fluid quality. The DNB ratio is not calculated, instead the code is forced into DNB by specifying a conservative DNB heat flux. The gap heat transfer coefficient-can be calcu-lated by the code; however, it is adjusted in order to force the full power steady-state temperature distribution to agree with the fuel heat transfer design codes. Further description of FACTRAN appears in Section 15.0.7. g 0952T:6 '15.4-30'
. i er ii - i r m . - . _ . .. . - _ _ _ _ _ . _ . - - . -. - - - - - - - - - -
System Overpressure Analysis Because safety limits for fuel damage specified earlier are not exceeded, ( there is little likelihood of fuel dispersal into the coolant. The pressure surge may therefore be calculated on the basis of conventional heat transfer from the fuel and prompt heat generation in the coolant. The pressure surge is calculated by first performing the fuel heat transfer calculation to determine the average and hot spot heat flux versus time. Using this heat flux data, a THINC calculation is conducted to determine the volume surge. Finally, the volume surge is simulated in a plant transient computer code. This code calculates the pressure transient taking into account fluid transport in the reactor coolant system and heat transfer to the steam generators. No credit is taken for the possible pressure reduction caused by the assumed failure of the control rod pressure housing. 15.4.5.2.1 Calculation of Basic ParaNeters ( Input parameters for the analysis are conservatively selected on the basis of values calculated for this type of core. The more important parameters are discussed below. Table 15.4-2 presents the parameters used in this analysis.
; Ejected Red Worths and Hot Channel Factors
- The values for ejected rod worths and hot channel factors are calculated using either three dimensional static methods or by a synthesis method employing one dimensional and two dimensional calculations. Standard nuclear design codes are used in the analysis. No credit is taken fc.- the flux flattening effects of reactivity feedback. The calculation is performed for the maximum allowed bank insertion at a given power level, as determined by the rod insertion limits. Adverse xenon distributions are considered in the calculation.
Appropriate margins are added to the ejected rod worth and hot channel factors - to account for any calculational uncertainties, including an allowance for ! nuclear power peaking due to densification. I 0952T:6 15.4-31 l
Power distributions before and after ejection for a " worst case" can be found C in Reference 5. During plant startup physics testing, ejected rod worths and power distributions are measured in the zero and full power redded configurations and compared to values used in the analysis. It has been founti that the ejected rod worth and power peaking factors are consistently overpredicted in the analysis. Reactivity Feedback Weichting Factors The largest temperature rises, and hence the largest reactivity feedbacks, occur in channels where the power is higher than average. Since the weight of a region is dependent on flux, these regions have high weights. This means that the reactivity feedback is larger than that indicated by a simple channel analysis. Physics calculations have been carried out for temperature changes with a flat temperature distribution, and with a large number of axial and radial temperature distributions. Reactivity changes were compared and effective weighting factors determined. These weighting factors take the form of multipliers which when applied to single channel feedbacks correct them to effective whole core feedbacks for the appropriate flux shape. In this analysis, since a one-dimensional (axial) spatial kinetics method is employed, axial weighting is not necessary if the initial condition is made to match the ejected rod configuration. In addition, no weighting is applied to the moderator feedback. A conservative radial weighting factor is applied to the transient fuel temperature to obtain an effective fuel temperature as a function of time accounting for the missing spatial dimension. These weighting factors have also been shown to be conservative compared to three dimensional analysis (Reference 5). Moderator and Doppler Coefficient The critical boren concentrations at the beginning of life and end of life are adjusted in the nuclear code in order to obtain moderator density coefficient curves which are conservative ccmpared to actual design conditions for the plant. As discussed above, no weighting factor is applied to these results. 0952T:6 15.4-32
L l The Doppler reactivity defect is determined as a function of power level using l a one dimensional steady-state computer code with a Doppler weighting factor of 1.0. The Doppler defect used is given in Section 15.0.4. The Doppler weighting factor will increase under accident conditions, as discussed above. Delayed Neutron Fraction, B Calculations of the effective delayed neutron fraction (S,ff) typically yield values no less than 0.70 percent at beginning of life and 0.50 percent 4 at end of life for the first cycle. The accident is sensitive to S if the ejected rod worth is equal to or greater than 6 as in zero power transients. In order to allow for future cycles, pessimistic estimates of 6 of 0.50 percent at beginning of cycle and 0.44 percent at end of cycle were used ir. the analysis. Trip Reactivity Insertion - ( The trip reactivity insertion assumed is given in Table 15.4-2 and includes the effect of one stuck RCCA. These values are reduced by the ejected rod reactivity. The shutdown reactivity was simulated by dropping a rod of the required worth into the core. The start of rod motion occurred 0.5 seconds after the high neutron flux trip point was reached. This delay is assumed to consist of 0.2 seconds for the instrument channel to produce a signal, 0.15 seconds for the trip breaker to open and 0.15 seconds for the coil to release the rods. A curve of trip rod insertion versus time was used which assumed that insertion to the dashpot does not occur until 3.3 seconds after the start of fall. The choice of such a conservative insertion rate means that there is over one second after the trip point is reached before significant shutdown reactivity is inserted into the core. This is a particularly important conservatism for hot full power accidents. The minimum design shutdown available for this plant at HZP may be reached only at end of life in the equilibrium cycle. This value includes an allowance for the worst stuck rod, adverse xenon distribution, conservative L 0952T:6 15.4-33 l
~ .
l Doppler and moderator defects, and an allowance for calculational uncertain-ties. Physics calculations for this plant have shown that th'e effect of two stuck RCCAs (one of which is the worst ejected rod) is to reduce the shutdown by about an additional one percent Ak, Therefore, following a reactor trip resulting from an RCCA' ejection accident, the reactor will be suberitical when the core returns to HZP. Depressurization calculations have been performed for a typical four-loop ( plant assuming the maximum possible size break (2.75 inch diameter) located in the reactor pressure vessel head. The results show a rapid pressure drop and a decrease in system water mass due to the break. The safety injection system is actuated on the coincidence of low pressurizer pressure and level within one minute after the break. The reactor coolant pressure continues to drop and reaches saturation (1100 to 1300 psi depending on the system temperature) in about two to three minutes. Due to the large thermal inertia of primary ! and secondary system, there has been no significant decrease in the reactor coolant system temperature below no-load by this time, and the depressuriza-
; tion itself has caused an increase in shutdown margin by about 0.2 percent ak'due to the pressure coefficient. The cooldown transient could not absorb the available shutdown margin until more than ten minutes after the break.
~ The addition of borated (2000 ppm) safety injection flow starting one minute after the break is much more.than sufficient to ensure that the core remains suberitical during the cocidown. j Reactor protection As discussed in Section 15.4.5.1.1, reactor protection for a rod ejection is provided by high neutron flux trip (high and low setting) and high rate of neutron flux increase trip. These protection functions are part of the , reactor trip system. No single failure of the reactor trip system will negate the protection functions required for the rod ejection accident, or adversely affect the consequences of the accident. 4 0952T:6 15.4-34
Results Cases are presented for both'beginning and end of life at zero and full power. ( 1. Beginning of Cycle, Full Power Control bank D was assumed to be inserted to its insertion limit The worst ejected rod worth and hot channel factor were conservatively calculated to be 0.23 percent Ak and 5.3 respectively. The peak hot spot clad average temperature was 2270'. The peak hot spot fuel center temperature reached melting, conservatively assumed at 4900*F. However, melting was restricted to less than 10 percent of the pellet.
- 2. Beginning of Cycle, Zero. power For this* condition, control bank D was assumed to be fully inserted and banks B and C were at their insertion limits. The worst ejected rod is
( located in control bank D and has a worth of 0.75 percent Ak and a hot channel factor of 11.5. The peak hot spot clad temperature reached 2621'F, the fuel center temperature was 4101*F.
- 3. End of Cycle, Full Power. .
Control bank D was assumed to be inserted to its insertion limit. The ejected rod worth and hot channel factors were conservatively calculated to be 0.23 percent ak and 5.9 respectively. This resulted in a peak clad temperature of 2175'F. The peak hot spot fuel temperature reached
- nelting conservatively assumed at 4800 F. However, melting was restricted to less than 10 percent of the pellet.
C 0952T:6 15.4-35
r
- 4. End of Cycle, Zero Power The ejected rod worth and hot channel factor for this case were obtained assuming control bank D to be fully. inserted and bank C at its insertion limit. The results were 0.90 percent ak and 20.0 respectively. The peak clad and fuel center temperatures were 2685'F and 3993*F. The Doppler weighting factor for this case is significantly higher thar. for the other cases due to the very large transient hot channel factor.
A summary of the cases presented above is given in Table 15.4-2. The nuclear power and hot spot fuel and clad temperature transients for the worst cases are presented in Figures 15.4.5-1 through 15.4.5-4. (beginning-of-life full power and end-of-life zero power.) The calculated sequence of events for the worst case rod ejection accidents, as shown in Figures 15.4.5-1 through 15.4.5-4, is presented in Table 15.4-1. For all cases, reactor trip occurs very early in the transient, after whien the nuclear power excursion is terminated. As discussed previously in Section 15.4.5.2.2, the reactor will remain suberitical following reactor trip. The ejection of an RCCA constitutes a break in the Reactor Coolant System, located in the reactor pressure vessel head. Following the RCCA ejection, the operator would follew the same ' emergency instructions as for any other loss of coolant accident to recover from the event. Fission product Release It is assumed that fission products are released from the gaps of all rods entering DNB. In all cases considered, less than 10 percent of the rods entered DNB based on a detailed three dimensional THINC analysis (Reference 5). 1 l l 0952T:6 15.4-36
.. j
Pressura Surce A detailed calculation of the pressure surge for an ejection worth of one dollar at beginning of life, hot full power, indicates that the peak pressure does not exceed that which would cause stress to exceed the faulted condition stress limits (Reference 5). Since the severity of the present analysis does not exceed the " worst case" analysis, the accident for this plant will not result in an excessive pressure rise or further damage to the Reactor Coolant System. Lattice Deformations A large temperature gradient will exist in the region of the hot spot. Since the fuel rods are free to move in the vertical direction, differential expansion between separate rods cannot produce distortion. However, the temperature gradients across individual rods may produce a differential expansion' tending to bow the micpoint of the rods toward the hotter side of ( the rod. Calculations have indicated that this bowing would result in a negative reactivity effect at the hot spot since Westinghouse cores are under-moderated, and bowing will tend to increase the under-moderation at the hot spot. Since the 17 x 17 fuel design is also under-moderated, the same effect would,be observed. In practice, no signifjcant bowing is anticipated, since the structural rigidity of the core is more than sufficient to withstand the forces produced. Boiling in the hot spot region would produce a net flow away from that region. However, the heat from the fuel is released to the water relatively slowly, and it is considered inconceivable that crossflow will-be sufficient to produce significant lattice forces. Even if massive and rapid boiling, sufficient to distort the lattice, is hypothetically postu-lated, the large void fraction in the hot spot region would produce a reduc-tion' in the total core moderator to fuel ratio, and a large reductien in this ratio at the hot spot. The net effect would therefore be a negative feed-back. It can be concluded that no conceivable mechanism exists for a net positive feedcack resulting from lattice deformation. In fact, a small ! negative feedback may restrlt. The effect is conservatively ignored in t'he analysis.
~
0952T:6 15.4-37
I f i 15.4.5.3 Conclusions
; Even on a pessimistic basis, the analyses indicate that the described fuel and clad limits are not' exceeded. It is concluded that there is no danger of sudden fuel dispersal into the coolant. Since the peak pressure does not '
exceed that which would cause stresses to exceed the faulted condition stress limits, it is concluded that there is no danger of further consequential damage to the Reactor Coolant System. The analyses have demonstrated that the upper limit in fission product release as a result of a number of fuel rods entering DNB amounts to ten percent. l I 4
- i.
- i C
0952T:6. 15.4-38' - g ~~ .-. m a , e.. ~se--
15.
4.6 REFERENCES
\
i
- 1. Risher, D. H., Jr. and Barry, R. F. , " TWINKLE - A Multi-Dimensional
( Neutron Kinetics Computer Code," WCAP-7979-A (Proprietary) and WCAP-8028-A (Non-Proprietary), January 1975.
- 2. Hargrove, H. G., "FACTRAN - A Fortran-IV Code for Thermal Transients in a U02 Fuel Rod," WCAP-7908, June 1972.
- 3. Burnett, T. W. T. , et al . , "LOFTRAN Code Description," WCAP-7907, June 1972.
- 4. Morita, T., et al., " Dropped Rod Methodology for Negative Flux Rate Trip Plants," WCAP-10297-P-A, June 1983.
- 5. Risher, D. H., Jr., "An Evaluation of the Rod Ejection Accident in Westinghouse Pressurized Water Reactors Using Spatial Kinetics Methods,"
( 6. WCAP-7588, Revision I-A, January 1975. Taxelius, T. G. (Ed), " Annual Report - SPERT Project, October,1968, September, 1969," Idaho Nuclear Corporation IN-1370, June 1970.
- 7. Liimataninen, R. C. and Testa, F. J., " Studies in TREAT of Zircaloy-2-Clad, UO2-Core Simulated Fuel Elements," ANL-7225, January -
June 1966, p.177, . November 1966.
- 8. Bi shop, A. A. , Sandberg, R. O. , and Tong, L. S. , " Forced - Convection " eat Transfer at High Pressure After the Critical Heat Flux," ASME 65-HT-31, August 1965.
4 L 0952T:6 15.4-39
TABLE 15.4-1 (Page 1) TIME SEQUENCE OF EVENTS FOR INCIDENTS WHICH CAUSE l REACTIVITY AND POWER DISTRIBUTION ANOS'ALIES (_ Accident Event Time (sec.) Uncontrolled Rod Initiation of uncontrolled 0.0 Cluster Control rod withdrawal from 10 -9 of - Assembly Bank nominal power Withdrawal from a Subcritical or Low Power Startup Condition Power range high neutron 10.4 flux low setpoint reached Peak nuclear power occurs 10.5 Rod begin to fall into core 10.9 Peak heat flux occurs. 13.0 Minimum DNBR occurs 13.0 Peak average clad temperature -13.4 occurs
-Peak average fuel temperature 13.7 occurs l
i 0952T:6 '15.4-40
l l 1
'I i
\
TABLE 15.4-1 (Page 2) TIME SEOUENCE OF EVENTS FOR INCIDENTS WHICH CAUSE REACTIVITY AND POWER DISTRIBUTION ANOMALIES C Accident Event Time (sec.) Uncontrolled RCCA bank withdrawal at power
- 1. Case A Initiation of uncontrolled RCCA 0 withdrawal at a high reactivity
, insertion rate (75 pcm/sec) Power range high neutron flux 1.11 high trip point reached Rods begin to-fall t'nto core 1.61 Minimum DNBR occurs 2.80
- 2. Case B Initiation of uncontrolled RCCA 0 withdrawal at a small reactivity
-insertion rate (1 pcm/sec)
( Overtemperature AT reactor-trip signal initiated 104.1 Rods begin to fall into core 106.1-Minimum DNBR occurs 107.2 l1 i l 0952T:6 -15.4-41 i
i TABLE 15.4-1 (Page 3) t TIME SE0VENCE OF EVENTS FOR INCIDENTS WHICH CAUSE REACTIVITY AND POWER DISTRIBUTION ANOMALIES Accident Event Time (sec.) Startup of an Initiation of pump startup 2.0 inactive reactor coolant loop at Power reaches P-8 trip 9.1 an incorrect setpoint temperature Rods begin to drop 9.6 Minimum DNBR occurs 10.4 Rod Cluster Control Assembly Ejection
- 1. Beginning-of- Initiation of roc ejection 0.0 Life, Full Power Power range high neutron flux 0.05 setpoint reached i
Peak nuclear power occurs 0.14 Rods begin to fall into core 0.55 l Peak' fuel average temperature 2.34 occurs I l Peak clad temperature occurs 2.41 - l . { 0952T:6- 15.4-42 . r
C TABLE 15.4-1 (Page 4) TIME SEOUENCE OF EVENTS FOR INCIDENTS WHICH CAUSE REACTIVITY AND POWER DISTRIBUTION ANOMALIES
-C Accident Event Time (sec.)
Peak heat flux occurs 2.42 Peak fuel center temperature 2.91 occurs
- 2. End-of-Life, Initiation of rod ejection 0.0 Zero Power Power range high neutron flux 0.18 low setpoint reached
. End-of-Life, Peak nuclear power occurs 0.21 Zero Power (cont.)
Rods begin to fall into core 0.68 Peak clad temperature occurs 1.56 Peak heat flux occurs 1.56 Peak fuel average temperature. 1.88' occurs
~
i
. Peak fuel center temperature 2.84 occurs L .
0952T:6 -15.4-43
l ( l TABLE 15.4-2 l PARAMETERS USED IN THE ANALYSIS OF THE ROD CLUSTER CONTROL ASSEMBLY EJECTION ACCIDENT BEGINNING BEGINNING END OF END OF TIME IN LIFE OF CYCLE OF CYCLE CYCLE CYCLE Power level, Percent 102 0 102 0 Ejected rod worth, Percent AK 0.23 0.75 0.23 0.90 Delayed neutron fraction, Percent 0.50 0.50- 0.44 0.44 Feedback reactivity weighting 1.30L 2.07 1.30 3.55 Trip reactivity, Percent AK 4.0 2.0 4.0 2.0 F before rod ejection 2.50 - 2.50 q - F after rod ejection 5.3 11.5 q 5.9 20.0 Number of operational pumps 4 2 4 2 Max. fuel pellet average 4113 3519 3951 3 502 temperature, "F Max. fuel center temperature, "F 5006 4101- 4893 3993 . Max. clad average tempera- 2270 2621 '2175- 2685 ture, 'F Max. fuel store energy, cal /gm 180 150 172 149 Percent fuel melt <20 0 <10 0 , 0952T:6 15.4-44
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C, r~~ w w w N N ~ ~ n n J J ' 19NG NfNINIH Figure 15.4.2 7 Minimum DNBR vs. Reactnrity insertion Rate: Roo Withcrawal frem 100% Power 15.4-54
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b f b O O O O ' ~ core I4 EAT FLUM (FRACTION OF NOMINAtt NUCLEAR POWER (FRAcil0N OF NOMINAt l E E E E E .C CR g C g g g c g I l ! l l l ; , , If. : \
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w [ 2000 -
- l l 1900 . . _ _ _
I O 50 100 150 200 TIME (SECONDS) i FIGURE 15.4.3-2 PRESSURI!J2 PRESStRI hNSir: AND ' CORI AVERAGE TIMPIRATURI TRANSIr:T FOR DROF?D RCD CLUSTER CONTRCL AS$ZMBLY 15.4-58
a a 4 ( 1.2 ( 7 E 1.0 - - R
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l .2 . ~ 0 0 5 10 15 20 25 30 . TIME (SECONO3) E t ,1ecae 15... 441 Imprecer Start. o cf an inactve Reacter Coctant Pump C 15.4-59
( 1.2 _ 1. 0 - - .. 5
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= 5 j .6- - " M 3
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1907 CHANNEL E W
.2 , ,
AVfAAGE CManntL 0 ' O 5 10 15 20 25 30 Tile (SECONOS} FIGURE 15.4.4-2 improper Startup of an inactive Reactor Coolant Pump 15.4-60 9
l ( i l 2.00 ( 2 1.75 - - 1.50 - ,,, 5 l B 1.25 . .
~
l E 3 1.00 .. E .75 . . y e u .50 .
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0 0 5 10 55 20 25 30 TIME (SEC305) i i FIGURE 15.4.4-3 Imcrecer Startue of an inac:m i Reacter Ccolant Pumo l l ! 15.4-61 l l l I 1
/
2600 . 2500 -
- 2400 -
G t. g 2300 -- -- R I c E 2200 -- - 5 2 E 2100 - -
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c 2000 -- 1900 700 ' 675 -
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w 650 - -- E M g 625 W . W 600 -- -
?=
575 -- ( m 550 -- J L 1 525 -- - 0 5 10 15 20 25 30 i l TIME (SECONOS) FIGURE 15.4.4-4 Imorcoor Startuo of an inac:ive Reactor Ccolant Pump 1 i 15.4-62 i l
( ( 3.00 ' l { f 2.75 - - 2.50 " { g
- 1 E
2.25 " " 2.00 - . l 1.75 ' 0 2.5 5.0 7.5 10.0 12.5 15.0 TIME (SECONOS) b FIGURE 15.4.4-5 tmorcoer Startuo of an inactive Reactor Coolant Pumo C 15.4-63
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o 2.0 -- .. 2 1 9 U
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W ,, f E ( 0 - - 0 2 4 6 8 10 TIME (SECONDS) FIGURE 15.4.5-1 Nuclear Power Transient. l 80L HFP Rod Election Accident l 15.4-64
l 6000 i C 5000 - - ( e cz* ~u~
.m$ F CENTER e- 4000 - ..
55
>6 M- -
FUEL AVERAGE .. m Ww I5 - U 3 2000 . . ge CLAD OUTER ( ce
'b u 1000 . .
0 , , , , 0 2 4 6 8 10 TIME (SECONDS) FIGURE 15.4.5-2 Hot Soot Fuel and Cad Temoerature vs Time, 80L HFP Rod Election Accident 15.4-65 l t l l _ -- -- . . , ., , .- .-
102 ; .
.=.
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TIME (SECOMOS)- FIGURE 15.4.5-3 Nuclear Power Transient.
^
EOL HZP Rod Ejection Accident ^ 15.4-66
C 1 i ( 6000 SMO - - "" _w 4000 - - '"
- FUEL CENTER y FUEL AvtRAGE
- 3000 --f a
$ CLAD OUTER 220 - ,
~
1000 " ~
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0 0 2 4 6 8 10 TIE (SECOMOS) b FIGURE 15.4.5-4 l Hot Spot Fuel and Clad Temperature vs. Time, EOL HZP Rod Ejection Accident C. 15.4-67 t
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15.5 I_NCREASE IN REACTOR COOLANT INVENTORY Discussion and analysis of the following event is presented in this section: Inadvertent Operation of Emergency Core Cooling System During Power Operation This event, considered to be ANS Condition II, causes an increase in reactor coolant inventory. Section 15.0.1 contains a discussion of ANS classifi-cations. 15.5.1 INADVERTENT OPERATION OF EMERGENCY CORE COOLING SYSTEM DURING POWER OPERATION 15.5.1.1 Identification of Causes and Accident Description Spurious Emergency Core Cooling System (ECCS) operation at power could be ( caused by operator error or a false electrical actuation signal. A spurious signal may originate from any of the safety injection actuation channels as described in the McGui're FSAR Section 7.3. Following the actuation signal, the suction of the coolant charging pumps is diverted from the volume control tank to the refueling water storage tank. The valves isolating the baron injectior tank from the charging pumps and the valves isolating the boron injection tank from the injection header then automatically open. The charging pumps then force concentrated (2000 ppm) boric acid solution from the boron injection tank, through the header and injection line, and into the cold leg of each loop. The safety injection pumps also start autocatically but provide no flow when the Reactor Coolant System is at normal pressure. The passive injection system and the low head system also provida no flow at normal Reactor Coolant System pressure. A Safety Injection System (SIS)- signal normally results in a reactor trip followed by a turbine trip. However,.it cannot be assumed that any single fault that actuates the SIS will also produce a reactor trip. If a' reactor ! 0956T:1/6 15.5-1
trip is generated by the spurious SIS signal, the operator should determine if the spurious signal was transient or steady state in nature. The operator must also determine if the safety ir.jection signal should be blocked. For a spurious occurrence, the operator would stop the safety injection and maintain the plant in the hot shutdown condition. If the ECCS actuation instrumenta-tion must be repaired, future plant operation will be in accordance with the Technical Specifications. If the Reactor Protection System does not produce an immediate trip as a result of the spurious SIS signal, the reactor experiences a negative reactivity excursion due to the injected baron causing a decrease in reactor power. The power mismatch causes a drop in T,yg and consequent coolant shrinkage. Pressurizer pressure and water level drop. Load will decrease due to the effect of reduced steam pressure on load after the turbiae throttle valve is fully open. If automatic rod control is used, these effects will be lessened until the rods have moved out of the core. The transient is eventually terminated by the Reactor Protection System low pressure trip or by manual trip. The time of trip is affected by initial operating conditions including core burnup history which affects initial boron concentration, rate of change of boron concentration, Doppler and moderator coefficients. Recovery from this second case is made ir. the same manner as described for the case where the SIS signal results directly ir, a reactor trip. The only difference is the lower T,yg and pressure associated with the power mismatch during the transient. The time at which reactor trip occurs is of little concern for this transient. At lower loads, coolant contraction will be slower, resulting in a longer time to trip. This event is classified as a Condition II incident (an incident of moderate frequency) as defined in Section 15.0.1.
'1 1
0956T:1/6 15.5-2 l
15.5.1.2 Analysis of Effects and Consecuences Method of Analysis The spurious operation of the Safety Injection Syrum is analyzed by employing the detailed digital computer program LOFTRAN (Reference 1). The code simu-lates the neutron kinetics, Reactor Coolant System, pressurizer, pressurizer relief and safety valves, pressurizer spr' y,a steam generator, steam generator safety valves, and the effect of the Safety Injection System. The program computes pertinent plant variables including temperatures, pressures, and power level. Because of the power and temperature reduction during the transient, operating conditions do not approach the core limits. Analysis of several cases has shown that the results are relatively independent of time to trip. A typical transient is presented representing minimum reactivity feedback. ( Results with maximum reactivity feedback are similar except that the transient is slower. For calculational simplicity, zero injection line purge volume was assumed in this analysis;.thus the boration transient begins immediately when the appropriate valves are opened. This accident is analyzed with the Improved Thermal Design Procedure as described in WCAP-8567. Plant characteristics and initial conditions are discussed in Section 15.0.3. The major assumptions are as follows:
- 1. Initial Operating Conditions Initial reactor power, pressure, and RCS temperatures are assumed to be at their nominal values. Uncertainties in initial conditions are included in the limit DNBR as described in WCAP-8567.
0956T:1/6 15.5-5
- 2. Moderator and Doppler Coefficients of Reactivity A least negative moderator temperature coefficient was used. A low (absolute value) Doppler power coefficient was assumed.
- 3. Reactor Control The reactor was assumed to be in manual control.
- 4. Pressurizer Heaters Pressurizer heaters were assumed to be inoperable in order to increase the rate of pressure drop.
- 5. Baron Injection At time zero two charging pumps inject 2000 ppm borated water into the cold leg of ee;h loop.
- 6. Turbine Load Turbine load was assumed constant until the governor drives the throttle valve wide open. Then turbine load drops as steam pressure drops.
- 7. Reactor Trip Reactor trip was initiated by low pressurizer pressure.
- Results Figures 15.5-1 through 15.5-3 show the transient response to inadvertent ^
operation of ECCS during power operation. Neutron flux starts decreasing immediately due to boron injection,'but steam flow does not decrease until 0956T:1/6 15.5-4 1
J later in the transient when the turbine throttle valve goes wide open. The mismatch between load and nuclear power causes T,yg, pressurizer water level, and pressurizer pressure to drop. Wnen the low pressure trip setpoint is reached the reactor trips and control rods start moving into the core. DNBR increases throughout the transient. The calculated sequence of events is shown on Table 15.5-1. After reactor trip, pressure and temperature slowly rise since the turbine is tripped and the reactor is producing some power due to delayed neutron fissions and decay heat. Recovery from this accident is discussed in Section 15.5.1.1. 15.5.1.3 Conclusiom Results of the analysis show that spurious safety injection without immediate reactor trip presents no hazard to the integrity of the Reactor Coolant System. . ( DNB ratio is never less than the initial value. Thus, there will be no cladding damage and no release of fission products to the Reactor Coolant-System. If the reactor does not trip immediat31y, the low pressure reactor trip will be actuated. This trips the turbine and prevents excest, cooldown, thereby expediting recovery from the incident. i i 15.
5.2 REFERENCES
- 1. Burnett, T. W. T. , et al. , "LOFTRAN Code Description," WCAP-7907, June
; 1972.
1
\
l 0956T:1/6 15.5 i
TABLE 15.5-1 TIME SE00ENCE OF EVENTS FOR INCIDENT WHICH RESULTS IN AN INCREASE IN REACTOR COOLANT INVENTORY Accident Event Time (sec) Inadvertent Actuation Spurious SI signal generated; , of ECCS During Power two Charging Pumps begin Operation injectir.g borated water 0 i Turbine throttle valve wide open, load begins to drop with steam pressure 49 Low pressurizer pressure reactor trip setooint ret.:hed- 106 Control rod motion begins 108 0956Til/6 '15.5-6 l
i ( 1.2
=. 1.0 - --
=_
( 3
=
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~i 5 560 - - --
m !!E BW o 550 - - -- 540 " _ 530 ' O 25 50 75 100 125 150 175 200 TIME (SECONDS) FIGURE 15.5-1 Inacvertent Operation of ECCS During Power Operation 15.5-7 , ___ __ - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - ~ ~ - - - ' - - - - ~ - - - - - - - ' - '
1 2400l 2300-- " 2
@ 2200 . -
U
$ 2100- - "
2 O 2 2000. . " E e a. 1900- - " 1800 1200 : g 1100..- ( 52 ( m
~
5 1000A -- g 900- - " W 600 1 I
.x f.
y 700 . "
=
h n. 600- - " 500- - " 400 0 f 25 50 75 100 125 150 175 200-( TIME (SECONDS) FIGURE 15.5-2 Inaavertent Operation of ECCS During Power Operation' 15.5-8
6.0 5.5 - - 5.0 -- , 4.5 - - - o 4.0 " " 3.5 - - 3.0 - - - 2.5 " " d' 2.0 1.2 ' ' 1.0 -- 5 ' g .8- - E E" .6- - 3 e ,4 . I m
.2- - --
l 0- : . l 0 25 50 75 100 125 150 175 200 TIME (SECONOS) FIGURE 15.5-3 Inaevertent Coeration of ECCS During Power Operation 15.5-9
15.6 DECREASE ?N REACTOR COOLANT INVENTORY Discussion and analysis of the following event is presented in this section: Inadvertent opening of a pressurizer safety or relief valve. 15.6.1 INADVERTENT OPENING OF A PRESSURIZER SAFETY OR RELIEF VALVE 15.6.1.1 Identification of Causes and Accident Description An accidental depressurization of the Reactor Coolant System could occur as a result of an inadvertent opening of a pressurizer relief or safety valve. Since a safety valve is sized to relieve approximately twice the steam , flowrate of a relief valve, and will therefore allow a much more rapid depressurization upon opening, the most' severe core conditions resulting from an accidental depressurization of the Reactor Coolant System are associated ( with an inadvertent opening of a pressurizer safety valve. Initially, the event results in a rapidly decreasing Reactor Coolant System pressure until this pressure reaches a value ccrresponding to the het leg saturation pressure. At this time, tnt pressure decrease is slowed considerably. The pressure continues to decrease throughout the transient. The effect of the pressure decrease would be to decrease power via the moderator density feedback, but the reactor centrol systec (if ir. the automatic mode) functions to mainten tM pow 2r and average coolant ter;perature until reactor trip cccers. Pressurizer lave) increases initially due to expansion caused by depressurization and then decreases following reactor trip. The reactor may be tripped by the following Reactor Protection System signals:
- 1. Overtemper'ature AT
- 2. Pressurizer low pressure 0942T:1/6 15.6-1
l l An inadvertent opening of a pressurizer safety valve is classified as an ANS Condition II event, a fault of moderate frequency. See Section 15.0.1 for a discussion of Condition II events. 15.6.1.2 Analysis of Effects and Consecuences Method of Analysis The accidental depressurization transient is analyzed by employing the detailed digital computer code LOFTRAN (Reference 1). The code simulates the neutron kinetics, Reactor Coolant System, pressurizer, pressurizer relief and safety valves, pressurizer spray, steam generator, and steam generator safety valves. The code computes pertinent plant variables including temperatures, pressures, and power level. This accident is analyzed with the Improved Thermal Design Procedure as described in WCAP-8567. Plant characteristics and initial conditions are discussed in Section 15.0.3. - In order to give censervative results in calculating the D'IBR during the transient, the following assteptions are mace:
- 1. Initial reactor power, pressure, and RCS tampernums irre assumaa to be at their nominal values. Ur.cer:Ainties in initial conditions are included in the lirtit DNBR as descrit+d in !: CAP-8557.
- 2. A least negative moderator coefficient of re4.ctivity is assumed. The spatial effect of void due to local or >ub:ocied boiling is not considered in the analysis with respect to reactivity feedback or core power shape.
- 3. A large (absolute value) Doppler coefficient of reactivity such that the resultant amount of positive feedback is conservatively high in order to retard any power decrease due to moderator reactivity feedback. .
0942T:1/6 15.6-2
Normal reactor control systems are not required to function; however, the rod control system is assumed to be in the automatic mode in order to hold the core at full power longer and thus delay the wip. This is a worst-case assumption; if the reactor were in manual control, an earlier trip could occur en low pressurizer pressure. The Reactor Protection System functions to trip the reactor on the appropriate signal. No single active failure will prevent the Reactor protection System from functioning properly. Results The svstem response to an inadvertent opening of a pressurizer safety or relief valve is shown in Figures 15.6-1 and 15.6-2. Figure 15.6-1 illustrates the nuclear power transient following the depressurization. Nuclear power is maintained at the initial value until reactor trip occurs on low pressurizer pressure. The pressure decay transient and average temperature transient following the accident are given in Figure 15.6-2. Pressure drops more rapidly while core heat generation is reducea via the trip, and would then slow once saturation temperature is reached in the hot leg. The DNBR decreases initially, but increases rapidly following the trip, as shown in Figure 15.6-1. The DNGR remain.; above the limit valte throughout the transient. The calculated sequence of events for the inadvertent opening of a pres-surizer safety valve incident is shown on Table 15.6-1. 15.6.1.3 Conclusions The results of the analysis show that the pressurizer low pressure and the overtemperature AT Reactor Protection System signals provide adequate protection against the RCS depressurization event. No fuel or clad damage is predicted for this accident. i l 15.
6.2 REFERENCES
I
- 1. Burnett, T. W. T. , et al . , "LOFTRAN Code Description," WCAP-7907, June 1972. 'l 0942T:1/6 15.6-3
l 1 o l \ l l LOCA ANALYSIS ( L 0587L:6 B-1
.. .. . . - . .. . , - - . - . _ ~ - . _ - . . . - - . - . .-. . ,
i I ^ TABLE OF CONTENTS l i j Section Description Page 15.6 Decrease in Reactor Coolant Inventory 15,6-7 l , 15.6.5 Loss,of Coolant Accidents Resulting from 15.6-7 4 a Spectrum of Postulated Piping Breaks 1 Within the Reactor Coolant Pressure Boundary ( 4 . j -15.6.6 References 15.6-19 a i i o ( ' f 1 I 4 h 1 s
- l, d
4 i B-ii ~ 0587L:6' I
-- , , -,, ,, , -, h v, , ..n. . + <r. , - -
1 i LIST OF TABLES Section Uescription Page 15.6.5-2 Large Break LOCA - Time Sequence of Events, 15.6-21 Imperfect Mixing 15.6.5-3 Large Break LOCA - Time Sequence of Events, 15.6-22 Perfect Mixing 15.6.5-4 Large Break LOCA Results, Imperfect Mixing 15.6-23 15.6.5-5 Large Break LOCA Results, Perfect Mixing 15.6-24 15.6.5-6 Safety Injection Pumped Flow 15.6-25 15.6.5-7 ' Nodal Representation of Core Transients Large Break 15,6-26 1 T C
'B-iti 0587L:6
m LIST OF FIGURES Number Title Page 15.6.5-1 Sequence of Events for Large Break LOCA Analysis 15.6-27 15.6.5-2 Code Interface Description for Large Break Model 15.6-28
-15.6.5-3 Fluid Quality - 0.4I 15.6-29
( . 15.6.5-4 Fluid Quality - 0.61 15.6-30 15.6.5-5 Fluid Quality - 0.81 15.6-31 15.6.5-6 Fluid Quality - 0.6P 15.6-32 15.6.5-7 Fluid Quality - 0.8P 15.6-33 15.6.5-8 Mass Velocity - 0.41 15.6-34 15.6.5-9 Mass Velocity - D.61 15.6-35 15.6.5-13 Mass Velocity - 0.8I 35.6-35 15.E.5-11 Mass Velocity - 0.6P 15.6-37 15.6.5-12 Mass velocity - 0.8P 15.6-38 15.6.5-13 Hea.t Transfer coefficient - 0.41 15.6-39 15.6.5-14 . Heat Transfer Coefficient - 0.61 15.6-40 , 15.6.5-15 Heat Transfer Coefficient - 0.8I 15.6-41 15.6.5-16 Heat Transfer Coefficient - 0.6P -15.6-42 B-iv 0587Li6
i LIST OF FIGURES (Continued) Number Title Page 15.6.5-17 Heat Transfer Coefficient - 0.8P 15.6-43 15.6.5-18 Core Pressure - 0.41 15.6-44 15.6.5-19 Core Pressure - 0.6I 15.6-45 15.6.5-20 Core Pressure - 0.81 15.6-46 15.6.5-21 Core Pressure - 0.6P 15.6-47 15.6.5-22 Core Pressure - 0.8P 15.6-48 15.6.5-23 Flowrate at Lower Half and Midplane at Core - 0.4I ( 15.6-49 15.6.5-24 Flowrate at Lower Half and Midplane at Core - 0.61 15.6-50 15.6.5-25 Flowate at Lower Half and N!dplace at Core - 0.81 15.6-51 15.6.5-26 Flowrate at Lower Half and M16.plar.e at Core - 0.6P 15.6-E2 15.6.5-27 F',ocate at Lower Palf and Mid: nano at Core - 0.8P 15.6-53 15.6.5-28 Flowrate at Upper Half and Top of Core - 0.4I 15.6-54 15.6.5-29 Flowrate at Upper Half and Top of Core - 0.61 15.6-55 15.6.5-30 Flowrate at Upper Half and Top of Core - 0.8I 15.6-56 15.6.5-31 Flowrate at Upper Half and Top of Core - 0.6P 15.6-57 15.6.5-32 Flowrate at Upper Half and Top of Core - 0.8P 15.6-58 L B-v 0587L:6
I t i l LIST OF FIGURES (Continued) Number Title Page 15.6.5-33 Void Fraction in Lower Half of Core - 0.4I 15.6-59 15.6.5-34 Void Fraction in Lower Half of Core - 0.6I 15.6-60 1 15.6.5-35 Void Fraction in Lower Half of Core - 0.8I 15.6-61
; 15.6.5-36 Void Fraction in Lower Half of Core - 0.6I 15.6-62 15.6.5-37 Void Fraction in Lower Half of Core - 0.8P 15.6-63 15.6.5-38 Void Fraction in Upper Half of Core - 0.41 15.6-64 15.6.5-39 Void Fraction in Upper Half of bore - 0.6I 15.6-65 15.6.5-40 Void Fraction in Upper Half of Core - 0.8I 15.6-66 C- .
15.6.5-41 Void Fraction in Upper Half of Core - 0.6P 15.6-67 15.6.5-42 Vc,id Fraction in Upper half of Core - 0.8P 15.6-68 15.6.5-43 Pea' s clad Temperature - Nodes 7 and 13 - 0.41 15.6-69 15.6.5-44 Feat Clad Temperature - Nodes 12 and 14 - 0.6I 15.6-70 15.6.5-45 Peak Clad Temperature - Nodes 10 and 13 - 0.8I 15.6-71 C 15.6.5-46 Peak Clad Temperature - Nodes 10 and 7 - 0.6P 15.6-72 15.6.5-47 Peak Clad Temperature - Nodes 8 and 7 - 0.8P' 15.6-73 15.6.5-48 Fluid Temperature - Neder 7 and 13 - 0.4I 15.6-74 B-vi 0587L:6
l l l i LIST OF FIGURES (Continued) Number Title Page 15.6.5-49 Fluid Temperature - Nodes 12 and 14 - 0.6I 15.6-75 15.6.5-50 Fluto Temperature - Nodes 10 and 13 - 0.8I 15.6-76 15.6.5-51 Fluid Temperature - Nodes 10 and 7 - 0.6P 15.6-77 15.6.5-52 Fluid Temperature - Nodes 8 and 7 - 0.8P 15.6-78 15.6.5-53 Reflood Transient - 0.41 15.6-79 15.6.5-54 Reflood Transient - 0.6I 15.6-80
~
15.6.5-55 Reflood Transient - 0.8I 15.6-81 ( 15,6.5-56 Reflued Transient - 0.6P 15 4-E IS.6.3-57 Reflood Trensia:.t + 0.F.P 15.6-83 15.6.5-58 Reflood Inlet Velocity - 0.4I 15.6-84 15.6.5-59 Reflood In?.et Velocity - C.6I 15.6-85 15.6.5-60 Reficod Inlet Velocity - 0.81 15.6-86 15.6.5-61 Reflood Inlet Velocity - 0.6P 15.5-87 l 15.6.5-62 Reflood Inlet Velocity - 0.8P 15.6-88 - I 15.6.5-63 Accumulator Flowrates - 0.41 15.6-89 15.6.5-64 Accumulator Flowrates - 0.61 15.6-90 B-vii 0587L:6
LIST.0F FIGURES (Continued) Number Title Page-15.6.5-65 Accumulater Flowrates - 0.8I 15.6-91 15.6.5-66 Accumulator Flowrates - 0.6P 15.6-92 15.6.5-67 Accumulator Flowrates - 0.8P 15.5-93 . 15.6.5-68 SI + Accumulator Flow - 0.41 15.6-94 1 15.6.5-69 SI + Accumulator Flow - 0.61 15.6-95
- 15.6.5-70 SI + Accumulator Flow - 0.81 15.6-56 15.4.5~71 SI + Accumulator Flow - 0.6P 15.6-97 15.6.5-72 SI + Ac:umulater Flow - 0.8P 15.6-98 I
I I 4 l i \
'I B-viii 0587L:6 !
1
L 15.6 DECREASE IN REACTOR COOLANT INVENTORY Discussion and analysis of the following event is presented in this section: _ loss of coolant accidents resulting from a spectrum of
~
postulated piping breaks within the reactor coolant pressure boundary. 15.6.5 LOSS OF COOLANT ACCIDENTS RESULTING .FROM A SPECTRUM OF POSTULATED PIPING BREAKS WITHIN THE REACTOR COOLANT PRESSURE BOUNDARY 15.6.5.1 Identification of Causes and Frequency Classification A LOCA is the result of a pipe rupture of the RCS pressure boundary. For the analyses reported here, a major pipe break (large break) is defined as a rupture with a total cross-sectional area equal to or grea-ter than 1.0 square feet (ft2). This svent is considered an ANS Con-dition IV event, a limiting fault, in that it is not expected to occur during the lifetime of the plant but is postulated as a conservative desisc basis (see Section 15.0.1). 1 A minor pipe creak (sr.all beer.k), as considered in this section, is defined as a rupture of the reactor coolant pressure boundary with e i total cross-sectienal area less than 1.0 ftE in which the normally cee*ating charging system flow is not sufficient to sustein pressurizer level and pressure. This is considered a Condition III event, in that it is an infrequent fault which ray occur during the life cf the plant. The Acceptance Critaria for the LOCA are described in 10CFR50.46 as follows:
- 1. The calculated peak fuel element clad temperature is below the ~
requirement of 2200'F.
- 2. The amount of. fuel element cladding ti.:t reacts chemically with water or steam does not exceed 1 percent of the total amount of Zircaloy in,the reactor.
0587L:6 15.6-7 r 'm- - _ i ._ __ ___ _.__._.____ i __ _ _ _ _ . _ _ _ . . _ _ _ _ _ _ _ _ _ . _ _ _ _ . _ _ _..__._____]
- 3. The clad temperature transient is terminated at a time when the core geometry is still amenable to cooling- The lor:lized cladding oxi-dation limits of 17 percent are not exceeded during or after quenching.
- 4. The core remains amenable to cooling during and after the break.
- 5. The core temperature is reduced and decay heat is removed for an extended period of time, as required by the long lived radioactivity remaining in the core.
These criteria were established to provide signficant margin in Emer-gency Core Cooling System (ECCS) performance following a LOCA. In all cases, small breaks (less than 1.0 ft2) yield results with more margin to the Acceptance Criteria limits than large breaks. i j 15.6.5.2 ' Sequence of Events and Systems Doerations Snould a major break occur, depressuritatior, of the RCS results it. a pressure decrease in the pressurizer. The reactor trip signal sub-sequently occurs whe i the pressurizer low pressure trip setpoint is reached. A safety injecticn signal is gsnerated wt.en the appropriate setooint is reached. Thess counterueasures will '.imit t,he consequences of tht accident in two ways:
- a. Reactor trip and borated water injectica complement void formation in causing rapid reduction of power to a residual level correspond-ing to fission product decay heat. However, no credit is taken in the LOCA analysis for boron content of the injection water. In addition, the insertion of control rods to shut down the reactor is neglected in the large break analysis,
- b. Injection of borated water provides for heat transfer from the core and prevents excessive clad temperatures.
~
0587L:6 15.6-8
Descriotion of Larce Break LOCA Transient The sequence of events following a large break LOCA is presented in ( Figure 15.6.5-1. Before the break occurs, the unit is in an equilibrium condition, i.e., the heat generated in the core is being removed via the secondary sys-tem. During blowdown, heat from fission product decay, hot internals, and the vessel continues to be transferred to the reactor coolant. At the beginning of the blowdown phase, the entire RCS contains subcooled liquid which transfers heat from the core by forced convection with some fully developed nucleate boiling. Thereafter, the core heat transfer is based on local conditions with transition boiling and forced convection to steam as the major heat transfer mechanisms. The heat transfer between the RCS and.the secondary system may be in either direction depending on the relative temperatures. In the case of continued heat addition to the second&ry, secondary system pressure increases and the main steam safety valves may detuate to limit the pressure. Makeup water to the secondary side is automatitally prcvided - by tne Avx111ary Feedwater System. The safety injection signal setuates a feedwater i:,clation signal which isnlates normal feedwater flow b) closing the main feedwater isolation valves and also initiates emergency feedwater flow by starting the auxiliary feedwater pumps. The secondary flow aids in the reduction of RCS pressure. When the Reactor Coolant System pressure falls below approximately 1250 psia, the upper head injection accumlators begin to inject borated water directly into the reactor upper head region. This water is di.rected from the upper head directly to all but eight peripheral assembies in the core via the RCC guide tubes and UHI support columns. This flow provides additional core cooling during the blowdown phase of the transient. A detailed description of the interactions of UHI water and those effects on the blowdown and subsequent reflood transients is given in Reference 2. L 0587L:6 15.6-9 4
d When the RCS depressurizes to approximately 430 psia, the accumulators begin to inject borated v.ater into the reactor coolant loops. Since the loss of offsite power is e,ssumed, the reactor coolant pumps are assumed to trip at the inception of the accident. The effects of pump coastdown are included in the blowdown analysis. ( The blowdown phase of the transient ends after the RCS pressure (initi-ally assumed at 2280 psia) falls to a value approaching that of the containment atmosphere. Prior to or at the end of the blowdown, the mechanisms that are responsible for the bypassing of emergency core cooling water injected into the RCS are calculated not to be effective. At this time (called end-of-bypass) refill of the reactor vessel lower d-plenum begins. Refill is complete when emergency core cooling water has filled the lower plenum of the reactor vessel which is bounded by the bottom of the fuel rods (called bottom of core recovery time). The reflood phase of the transient is defined as the time period lasting from the end-of-refill until the reactor vessel has been filled with water to the extent that the core tamperature rise has been terminated. From the later stage of blowdown and than the beginning-of-reflood, the , safety injection accumulator tanks rapidly discharge borated cooling water into the RCS, contributing to the filling of the reactor vessel downcomer. The downcomer water elevation head provides the driving force required for the refloocing of the reactor core. The low head and high tead safety injection pumps aid in the filling of the downcomer and subsequently supply water to maintain a full downcomer and complete the reflooding process. Pumped safety injection flows are provided in Table 15.6.5-6. Continued operation of the ECCS pumps supplies water during long term
~
cooling. Core temperatures have been reduced to long-term steady state levels associated with dissipation of residual heat generation. After-the water level of the refueling water storage tank reaches a minimum allowable value, coolant for long-term cooling of the core is obtained by switching to the cold leg-recirculation phase of operation in'which 0587L:6 15.6-10
# p
; spilled borated water is drawn from the engineered safety features sumps by the low head safety injection (residual heat removal) pumps and j returned to the RCS cold legs. The Containment Spray System continues
( to operate to further reduce Containment pressure. Approximately 15 hours after initiation of the LOCA, the ECCS is realigned to supply water to the RCS hot legs in order to control the boric acid concentration in the reactor vessel. Description of Small Break LOCA Transient Ruptures of small cross-sections will cause expulsion of the coolut at a rate which can be accommodated by the charging pumps which would maintain an operational ~ water level in the pressurizer permitting the operator to execute an orderly shutdown. The coolant which would be released to the containment contains the fission products existing at equilibrium. ( The maximum break size for which the normal makeup systems can maintain the pressurizer level is obtained by comparing the calculated flow from the Reactor Coolant System through the postulated break against the charging pump makeup flow at normal Reactor Coolant System pressure, i.e., 2250 psia. A makeup flow rate from one centrifugal charging pump is typically adequate to sustain the pressurizer level at 2250 psia for a break through a 0.375 inch diameter hole. This break results in a - loss of approximately 17.25 lb/sec. Shouldalargerbreakoccur,depressurizationoftheReactorCoolant System causes fluid to flow to the Reactor Coolant System from'the pressurizer, resulting in a pressure and level decrease in the-pressurizer. Reactor trip occurs when the pressurizer low pressure trip setpoint is reached. The Safety Injection-System is actuated when the appropriate setpoint is reached. The consequences of the accident are limited in two ways: I i l l 0587L:6 15.6-11 l
- 1. Reactor trip and borated water injection complement void formation in causing _ rapid reduction of nuclear power to residual level corresponding to the delayed fission and fission product decay.
- 2. Injection of borated water ensures sufficient flooding of the core '
to prevent excessive clad temperatures. l l Before the break occurs, the plant is in an equilibrium condition, i.e., the heat gei.erated in the core is being removed via the secondary system. During blowdown, heat from decay, hot internals, and the vessel continues to be transferred to the Reactor Coolant System. The heat transfer between The Reactor Coolant System and the secondary system may be in either direction depending on the relative temperature. In the case of continued. heat addition to the secondary, system pressure increases and steam dump.may occur. Makeup to the secondary side is automatically provided by the auxiliary feedwater pumps. The safety injection signal stops normal feedwater flow by closing the main feedwater line isolation valves and initiates emergency feedwater flow by starting auxiliary feedwater pumps. The secondary flow aids in the reduction of Reactor Coolant System pressure. When the Reactor Coolant System depressurizes to the upper head accumulator setpoint pressure, the upper head accumulator begins injecting borated water into the reactor vessel upper head. A description of the operation of upper head injection during small break transients can be found in WCAP-8479, - Revision 2 (Reference 1). When the RCS depressurizes to approximately 430 psia, the cold leg accumulators begin to inject water into the coolant loops. The reactor coolant pumps are assumed to be tripped at the initialization of_the accident and effects of pump coastdown are included in the blowdown analyses. 0587L:6 -15.6-12
l 4 15.6.5.3 Core and System Performance 15.6.5.3.1 Mathematical Model The requirements of an acceptable ECCS evaluation model are presented in Appendix K of 10CFR50.' l Laroe Break LOCA Evauation Model The analysis of a large break LOCA Transient is divided into three phases: 1)' blowdown, 2) refill, and 3) reflood. There are three distinct transients analyzed in each phase, including the thermal-hydraulic transients in the RCS, the pressure and temperature transient within the Containment, and the fuel and clad temperature transient of the hottest fuel rod in the core. Based on these considerations, a system of interrelated computer codes has been - developed for the analysis of the LOCA. The description of the various aspects of the LOCA analysis methodology is given in Reference 1 or 2. These documents describe the major phenomena modeled, the interfaces among the computer codes, and the-features of the codes which ensure compijance with the' Acceptance Criteria. The differences between the approved non-UHI Westinghouse Appendix K Model and the model used for these analyses are reported in WCAP-8479, Revision 2 (Reference 1). The thermal analyses reported in-this section were performed with an upper head fluid temperature of T cold. The UHI accumulator pressure setpoint ensures UHI activation prior to upper head fluid flashing. The SATAN-VI, POWLOCTA, WREFLOOD, l LOTIC, and LOCTA-IV codes which are used in the LOCA analysis are described in detail in References 1 through 6. These codes are used to assess the core heat transfer geometry and to determine if the core ~
~
remains amenable to cooling throughout and subsequent.to the blowdown, refill, and reflood phases of the LOCA. The SATAN-VI computer. code analyzes the thermal-hydraulic transient in the RCS during blowdown ~and refill while the POWLOCTA calculates the average channel'axialT L 0587L:6 .15.6-13 1
'l l
temperature distribution during this period of the transient. The WREFLOOD computer code is used to calculate the thermal-hydraulic transient during the reflood phases of the accident. The LOTIC computer code is used to calculate the centainment pressure transient during all three phases of the LOCA analysis. Similarly, the LOCTA-IV computer code is used to compute the thermal transient of the hottest fuel rod during the three phases. Fuel parameters input to the LOCTA code were taken from a new version of the PAD code. SATAN-VI'is used to calculate the RCS pressure, enthalpy,' density, and j the mass and energy flow rates in the RCS, as well as steam generator energy transfer between the primary and secondary systems as a function of time during the blowdown phase of the LOCA. SATAN-VI also calculates
~
the accumulator water mass and internal pressure and the pipe break mass and energy flow rates that are assumed to be vented to the containment during blowdown. At the end of the blowdown and refill phases, these data are transferred to the WREFLOOD code. Also at the end-of-blowdown and refill phases, the mass and energy release rates during blowdown are transferred to the LOTIC code for use in the determination of the containment pressure response during these phases of the LOCA. Additional SATAN-VI output data from the end-of-blowdown, including the core pressure, and the core power decay transient, are input to the LOCTA-IV code. With input from the SATAN-VI and POWLOCTA codes, WREFLOOD uses a system l thermal-hydraulic model to determine the core flobding rate, the coolant j pressure and temperature, and the quench front height during the reflood phases of the LOCA. WREFLOOD also calculates the mass and energy flow addition to the containment through the break. Since the mass flow rate to the containment depends upon the core flooding rate and the local _ core pressure, which is a function of the containment backpressure, the transient pressure computed by the LOTIC code is input to the WREFLOOD . code. WREFLOOD is also linked to the LOCTA-IV-code in that thermal-hydraulic parameters from WREFLOOD are used by LOCTA-IV in its calculation of the fuel temperature. 0587L:6 15.6-14
< n 4 A . w.
i l LOCTA-IV is used throughout the analysis of the LOCA transient to calculate the fuel cladding temperature and metal-water reaction of the hottest rod in the core. Dynamic steam cooling is included in the l LOCTA-IV calculation as described in Reference 7. Schematic representation of the computer code interfaces is given in Figure 15.6.5-2. Tha large break analysis was performed with the NRC-approved Westinghouse UHI ECCS Evaluation Model, Reference 2, as updated to correspond to the 1981 version of the non-UHI Evaluation Model. Small Break LOCA Evaluation Model The WFLASH program used in the analysis of the small break LOCA is an extension of the FLASH-4 code [8] develped at the Westinghouse Bettis Atomic Power Laboratory. The WFLASH program permits a detailed spatial representation of the RCS. The RCS is nodalized into volumes interconnected by flowpaths. The broken loop is modeled explicitly with the intact loops lumped into a second loop. The transient behavior of the system is determined from the governing conservation equations of mass, energy, and momentum applied through the system. A detailed descripton of WFLASH is given in Reference 9. . The use of WFLASH in the analysis involves, among other things, the representation of the reactor core as a heated control volume with the associated bubble rise model to permit a transient' mixture height calculation. The multinode capability of the program enables an explicit and detailed spatial representation of various system components. In particular, it enables a proper calculation of the behavior of the loop seal during a loss of coolant transient. 0587L:6 15.6 .
Clad thermal analyses are performed with the LOCTA-IV code (Reference 6) which uses the RCS pressure, fuel rod power history, steam flow past the uncovered part of the core, and mixture height history from the WFLASH hydraulic calculations as input. 15.6.5.3.2 Inout Parameters and Initial Conditiens The bases used to select the numerical values that are input parameters to the analysis have been conservatively determined from extensive sensitivity studies (refer. to Reference 1). In addition, the requirements of Appendix K regarding specific model features were met by selecting models which provide a significant overall conservatism in the analysis. The assumptions made pertain to the conditions of the reactor and associated safety system equipment at the time that the LOCA occurs and include such items as the core peaking factors, the containment pressure, and the performance of the ECC$.. ' Decay heat generated tnroughout the transient is also conservatively calculated as required by Appendix K, 10CFR50.46. The worst break (CD = 0.6) was run with a variation in UHI accumulator 3 volume delivery ,for the perfect (1011 ft ) and imperfect mixing case 3 (790 ft ) assumptions. The delivered volume considered in the analysis encompasses the volume delivery band associated with UHI delivery uncertainties. Cases presented herein provide the results of a conservative application of this range of values. The imperfect mixing case was arilyzed conservatively at a low delivery volume since the upper head drains earlier in the transient and I subsequently voids the lower plenum and core. The imperfect mixing case i was also run at a higher pressure (1300) than the perfect mixing case i 0587L:6 15.6-16
. , - a -,n.
l i l l j (1200) to allow for uncertainty in accumulator setpoint pressure. Similarly, the high pressure for the imperfect mixing case represents the most conservative case since the smaller accumulator volume would be delivered in a shorter amount of time and earlier in the blowdown C transient, thereby providing for a longer core heatup time. 15.6.5.3.3 Results Laroe Break Results Based on the results of the LOCA sensitivity studies, Reference li the limiting large break was found to be the double ended cold leg guillotine (DECLG). Therefore, only the DECLG break is considered in the large break ECCS performance analysis. Calculations were performed for a range of Moody break discharge coefficients. The results of these calculations are summarized in Tables 15.6.5-2 through 15.6.5-5. ( Figures 15.6.5-3 through 15.6.5-77 present the parameters of principal interest from the large break ECCS analyses. For all cases analyzed, transients of the following parameters are presented: Figures 15.6.5-3 The following quantities are presented at through the clad burst location and at the hot spot Figures 15.6.5-17 (location of maximum clad temperature), both on the hottest fuel rod (hot rod):
- 1. fluid quality
- 2. mass velocity
- 3. heat transfer coefficient The heat transfer coefficient shown is ~
calculated by the LOCTA IV code. Figures 15.6.5-18 The system pressure shown is the calculated thruugh pressure in the core. Core flowrates and Figures 15.6.5-42 core void fraction are also presented. 0587L:6 If.6-17
l I Figures 15.6.5-43 These figures show the hot spot clad through temperature transient and the clad Figures 15.6.5-52 temperature transient at the burst location. The fluid temperature shown is also for the hot spot and burst location. The nodal notation of the figures is defined in Table 15.6.5-7. Figures 15.6.5-53 These figures show the core reflood through transient. Figures 15.6.5-62 Figures 15.6.5-63 These figures show the Emergency Core through
- Cooling System flowrates for all cases Figures 15.6.5-72 analyzed. Both UHI and cold leg accumulators are included in the figures.
As described earlier, the cold leg accumulator delivery during blowdown is discarded until the end of bypass is calculated. Cold leg accumulator flow, however, is established in refill-reflood
, calculations. The cold leg accumulator {
flow assumed is the sum of that injected in the intact cold legs. The maximum clad temperatures calculated for a large break is 2175'F, which is less than the Acceptance Criteria limit of 2200*F of 10CFR50.46. The maximum local metal water reaction is well below the embrittlement limit of 17 percent as required by 10CF50.46. The total core metal-water reaction is less than 0.3 percent for all breaks, as compared with the 1 percent criterion of 10CFR50.46, and the clad temperature transient is terminated at a time when the core geometry is still amenable to cooling. As a result, the core temperature will continue to drop and the ability to remove decay heat generated in the fuel for an extended period of time will be provided. 0587L:6 15.6-18
l l Small Break Results As noted previously, the calculated peak cladding temperature resulting from a small break LOCA is mucn less than that calculated for a large break. Since the major change in input parameters due to the smaller optimized fuel rod is a slight increase in core flow area, the small break ECCS results with optimized fuel would not be greatly different than the standard 17x17 small break ECCS results. The maximum calculated peak cladding temperature for this plant configuration with standard 17x17 fuel was 1499'F for a 6-inch diameter break. This is much less than both the worst case large break peak clad temperature and the acceptance criteria limit of 2200'F in 10 CFR 50.46. A fuel specific analysis is not presented due to the 700*F difference between the standard fuel analysis and the acceptance criteria. 15.
6.6 REFERENCES
- 1. Young, M. Y., " Westinghouse Emergency Core Cooling System Evaluation Model Application to Plants Equipped with Upper Head Injection,"
WCAP-8479, (Westinghouse Proprietary), and WCAP-8480, January 1975.
- 2. Bordelon, F. M., Massie, H. W., and Zorden, T. A., " Westinghouse ECCS Evalution Model-Summary," WCAP-8339, (Non-Proprietary) July 1974.
- 3. Bordelon, F. M., et al., " SATAN-VI Program: Comprehensive Space Time Dependent Analysis of Loss of Coolant," WCAP-8302, (Proprietary) June 1974, and (Non-Proprietary) June 1974.
- 4. Kelly, R. D., et al., " Calculated Model for Core Reflooding (Proprietary) June -1974, and (Non-Proprietary) June 1974.
- 5. Hsieh, T., and Raymond, M., "Long Term Ice Condenser Containment LOTIC Code Supplement 1," hCAP-8355 Supplement 1, May 1975, WCAP-8354 (Proprietary), July 1974.
L 0587L:6 15.6-19
- 6. Bordelen, F. M., et al., "LOCTA-IV Program: Loss of Coolant Transient Analysis," WCAP-9220 (Proprietary) February 1979, and WCAP-9221 (Non-Proprietary) February 1978.
- 7. Eicheldinger, C., " Westinghouse ECCS Evaluation Model, February, 1978 Version," WCAP-9220 (Proprietary) February 1979, and WCAP-9221 (Non-Proprietary) February 1978.
- 8. Porsching, T. A., et al., " FLASH-4: A Fully Implicit FORTRAN-IV Program for the Digital Simulaition of Transients in a Reactor Plant," WAPD-TM-84, Bettis Atomic Power Laboratory, March 1969.
- 9. Esposito, V. J. , Kesavan, K. , and Maul, B. A. , "WFLASH - a FORTRAN IV Computer Program for Simulation of Transients in a Multi-Loop PWR," WCAP-8200, Revision 2, (Proprietary) July 1974, and WCAP-8261, Revision 1, (Non-Proprietary) July 1914.
- 10. Rahe, E. P., Westinghouse Letter to Thomas, C. O., NRC, Letter Number NS-EPR-2673, October 27, 1982,
Subject:
" Westinghouse Revised PAD Code Thermal Safety Model," WCAP-8720, Addendum 2 (Proprietary) .
L g 0587L:6 15.6-20
l I TABLE 15.6.5-2 l LARGE BREAK LOCA TIME SEQUENCE 0.: EVENTS, IMPERFECT MIXING DECL CD = 0.8 DECL CD = 0.6 DECL CD = 0.4 (Sec) (Sec) (Sec) START 0.0 0.0 0.0 Rx Trip Signal , 0.58 0.59- 0.60 S. I. Signal 4.02 4.14 4.33 Acc. Injection (CL) 14.9 17.5 24.3 End of Blowdown 79.2 '76.5 91.6 Bottom of Core -Recovery 79.2 76.5 91.6 Acc. Empty (CL) 141.9,, 1,45.8 160.2 Pump Injection 29.02 29.14 29.33 End of Bypass 48.9 40.4 61.7' i c 0587L:6 15.6-21 i
TABLE 15.6.5-3 LARGE BREAK LOCA TIME SEQUENCE OF EVENTS' PERFECT MIXING C9 = 0.8 C9 = 0.6 DECLG DECLG (sec) (sec) Start 0.0 0.0 Reactor Trip Signal 0.58 0.59 Safety Injection Signal 4.02 4.15 Cold Leg Accumulator Injection 14.9 17.0 End of Blowdown 105.0 113.5 Bottom of Core Recovery- 105.0 113.5 l Cold Leg Accumulator Empty 150.0 153.5 Pump Injection 29.02 29.15 End of Bypass 54.9 55.5' [
~
t 0587L:6- 15.6-22 ; l
TABLE 15.6.5-4 LARGE BREAK LOCA RESULTS, IMPERFECT MIXING i CD = 0.8 CD = 0.6 CD = 0.4 Results DECLG DECLG DECLG Peak Ciad Temperature, 'F 2103 2175 2187 Peak Clad Location, ft 7.5 7.5 7.25 Local Zr/H 2O Reaction, (max)% 4.7 5.6 6.5 Local Zr/H2 O Reaction Location, ft 7.5 7.5 7.5 Total Zr/H 2O Reaction, % <0.3 <0.3 <0.3 Hot Rod Burst Time, sec 72.7 66.6 80.5 Hot Rod Burst Location, ft 6.5 7.00 5.75 Core Power, MWt, 102% of -
,3411 -3411 3411 Peak Linear Power, kw/ft, 102% of 11.71 11.71 12.30 Peaking Factor (at License Rating) 2.15 2.15 2.26 C
I 0587L:6 > 15.6-23
TABLE 15.6.5-5 LARGE BREAK LOCA RESULTS - PERFECT MIXING CD = 0.8 CD = 0.6 Results DECLG DECLG Peak Clad Temperature, *F 2150 2189 Peak Clad Location, ft 5.75 5.75 Local Zr/H 2O Reaction, (max)*4 3.8 3.9 Local Zr/H2 O Location, ft 6.5 6.0 Total Zr/H 2O Reaction, % <0.3 <0.3 Hot Rod Burst Time, sec 67.1 64.6 Hot Rod Burst Location, ft 6.0 6.5 Core Power, MWt, 102% of 3411 3411 Peak Linear Power, kw/ft, 102*; of 12.30 12.30 Peaking Factor (at License Rating) 2.26 2.26 Accumulator Water Volume (Cold Leg,-Nominal 3 1078 ft Per Accumulator - Setpoint Value) Accumulator Water Volume Delivered (UHI, 3 790 ft Imperfect Mixing Nominal
- Delivered Value) 3 1011 ft Perfect Mixing L
0587L:6 15.6-24
i
- j. TABLE 15.6.5-6 f
SAFETY INJECTION PUMPED FLOW Pressure SI Flow (psia) (lb/sec) I
- 14.7 493.2
;- 34.7 437.2-i 54.7 378.6 74.7 315.5 114.7 198.4-214.7 100.2 614.7 81.2
, 1014.7 . 58.5 l 3014.7 0.0 ( r j , i- . L
'0587L:6 15.6-25 4
.l-,
.*-r- - - a n,- - < , , . , , .,, w ~ $- - -- v
1 1 1 TABLE 15.6.5-7 NODAL REPRESENTATION OF CORE TRANSIENTS - LARGE BREAK - Elevation from Elevation from Node Bottom of Core (ft.) " Node" Bottom of Core (ft.) 1 0.0 11 6.75' 2 1.5 12 7.00 3 3.0 13 7.25 4 4.0 14 7.50 5 5.0 15 7.75 6 5.5 15 8.0
- 7 5.75 17 9.0 8 6.0 18 10.5 .
i 9 6.25 19 12.0 10 6.50
- Applicable to the nodas in Figures 15.6.5-3 through 15.6.5-52 l
C 0587L:6 15.6 , m - -,-e r * ,
l
. l A i:IAx 000*JR$
l .tia: T:4 TRe> ' :usest: r3 errrgevirge pectgggg) (
} UNI ACCUv0LA T;A '*JECTIch 8 _ <
L
. F J:9fD !aFETY ik.ECTIC% !11?'AL IMl-l CONT. F AESS. OR '.0 i P( !UR IZER PE!!$. ) i 0
w
' /? 80 $1FETY I4JECTl0M !!3R$ (L!!UNIMG CFF$iTE #C"!4 '?tlLLELE)
D !.",L.3.. LEG &COUMJ' LATCR INJE: TION W
'J M t TERitatiTED n
( Y 0;?'..h':.M?tELT EMD CF 1 F'15 i g_.At Pi..e theT:1TiCN (t; J:*rr . ?!ITE 80.6 *,?
. . i)
- g. - 'Ua'!0 U.FI TY .NEC'4tg !!sths (1130MihC LOSS CF Or85iTE <0-Es)
~
E F E*0 0F !L:40 Cam 1 '"* T ( BOTTOM 0F CORE RECOVERY t L E CONTa iMuf nr ME AT t!=0 val SYSTEM in lTI A TION f a$$UM14G LCSS OF 0FFSITE F0a!R) F g COL LEC ACCUMULATCA3 EMPTY . 0 - 0 D
!".ni QJ!n;4ED C &
L 0 g..;;3 . *3 n _ . - . . f? 4EC 9Ct.at;CN CN *n37 LCw LEvf L ALanu (MANUAL ACTICu)
~-
G
?
E a
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C 0 0 G I f a ;nggg $ no agg,'LL PERIOD FOR THE PE4FECT MIIING CASES T Figure 15.6.5-1 Sequence of Events for Large Break Loss-of-Coolant Analysis 0587L:6 15.6-27 I l
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l.4000 UHl N00EL 5 ATAN (060.5 IMPERF H!I FOR DAP OATA P[R S AT AN DECK 3-82.is M00(L A5 IR ANSPOSE D AND A(Vl5[0
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- o 2 ? 8 8 2 IINE ISECB FIGllllE 15.6.5-34 Void Fraction in Lower llalf of Core - 0.61
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1 o o ATTACIDENT 2B l l RELOAD SAFETY EVALUATION i MCGUIRE NUCLEAR STATION UNIT 1 CYCLE 2 j November, 1983 . Edited by: D. J. Petrarca e 1 I Approved: l [G.~ Arlotti, Manager ~ Fuel Licensing and Program Support - Nuclear Fuel Division i e
~l
'l 0827L:6 -
h
TABLE OF CONTENTS e Title Pace s
1.0 INTRODUCTION
AND
SUMMARY
1 1.1 Introduction 1 1.2 General Description - 2 1.3 Conclusions 2 2.0 REACTOR DESIGN 4 2.1 Mechanical Design 4 2.2 Nuclear Design 5 2.3 Thermal and Hydraulic Design ' 6 6 3.0 POWER CAPABILITY AND ACCIDENT EVALUATION 7 3.1 Power Capability 7 3.2 Accident Evaluation 7 3.2.1 Kinetic Parameters 8 3.2.2 Control Rod Worths 8 3.2.3 Core Peaking Factors 8 4.0 TECHNICAL _ SPECIFICATION CHANGES 10
5.0 REFERENCES
11 APPENDIX A - Technical Specification Page Changes-,
-.i 0827L:6
LIST OF TABLES d Table Title Page 1 Fuel Assembly Design Parameters 12 2 Kinetic Characteristics 13
, 3 Shutdown Requirements and Margins 14 i .
4 Control Rod Ejection Accident Parameters 15 LIST OF FIGURES
- Fioure Title Pace Core Loading Pattern and Source and 1
16 Burnable Absorber Locations 11 V
- CS27L:6 . .
1.0 INTRODUCTION
AND
SUMMARY
1.1 INTRODUCTION
s This report presents an evaluation for McGuire Unit 1, Cycle 2, which demonstrates that the core reload will not adversely affect the safety of the plant. This evaluation was performed utilizing the methodology described in WCAP-9273, " Westinghouse Reload Safety Evaluation Methodology"II) . McGuire Unit 1 is operating in Cycle IA" with all Westinghouse 17x17 low parasitic (STD) fuel assemblies. For Cycle 2 (expected startup early 1984) and subsequent cycles, it is planned to refuel the McGuire Unit I core with Westinghouse 17x17 optimized fuel assembly (OFA) regions. In the OFA transition licensing submittal (2) to the NRC, approval was
. requested for the transition from the STD fuel design to the OFA design and the associated proposed changes to the McGuire Units 1 and 2 l .
Technical Specifications. The licensing submittal justifies the compatibility of the OFA design with the STD design in a transition core as well as a full 0FA core. The OFA transition licensing submittal (2) contains mechanical, nuclear, thermal-hydraulic, and accident evaluations which are applicable to the Cycle 2 safety evaluation. All of the accidents comprising the licensing bases (2,3) which could potentially be affected by the fuel reload have been reviewed for the Cycle 2 design described herein. The results of new analyses are included in the above mentioned licensing submittal and in this evaluation, and the justification for the applicability of previous results for the remaining analyses is presented.
' " Safety Evaluation for the W. B. McGuire Unit 1 Mid-Cycle 1 Core Redesign, April 1983.
s . 0827L:6 1 l 4
1.2 GENERAL DESCRIPTION 2 The McGuire Unit 1 Cycle 2 reactor core will be comprised of 193 fuel , assemblies arranged in the core loading pattern configuration shown in Figure 1. During the Cycle IA/2 refueling, 60 STD fuel assemblies will be replaced with 60 Region 4 optimized fuel assemblies. A summary of the Cycle 2 fuel inventory is given in Table 1. Nominal core design parameters utilized for Cycle 2 are as follows: Core Power (MWt) 3411 System Pressure (psia) 2250 Core Inlet Temperature (*F) 559.2 Thermal Design Flow (gpm) 386,000 Average Linear Power Density (kw/ft) 5.44
, (based on 144" active fuel length)
1.3 CONCLUSION
S From the evaluation presented in this report, it is concluded that the Cycle 2 design does not cause the previously acceptable safety limits to be exceeded. This conclusion is based on the following:
- 1. Cycle 1A burnup is between 14767 and 15533 MWD /MTU.
- 2. Cycle 2 burnup'is limited to 10200 MWD /MTU.
- 3. The analyses and proposed Technical Specification changes submitted in support of the OFA transition licensing submittal (2) are approved by the NRC prior to Cycle 2 startup.
0827L:6 .2
- 4. There is adherence to plant operating limitations given in the Tech-nical Specifications with the exception of the proposed changes submitted in support of the OFA transition licensing submittalI2) and the Technical Specification changes given in Appendix A. With these revisions, there is adherence to all plant operating limita-tions in the. Technical Specifications.
e i 0827L:6 3-5
2.0 REACTOR DESIGN 2.1 MECHANICAL DESIGN The new Region 4 fuel assemblies are Westinghouse OFAs. The mechanical description and justification of their compatibility with the
. Westinghouse STD design in a transition core is presented in the OFA transition licensing submittal.(2)
The OFAs and Core Components are designed to be handled by existing handling tools. The control rods, burnable absorber rods, and source rods are compatible with both the STD and 0FA designs. The standard design thimble plugging devices used in Cycle 1 are not mechanically compatible with the OFA design. Dually compatible thimble plugging devices have been designed for use with both th'e STD and 0FA designs. These dually compatible thimble plugging devices were used in Cycle 1A to replace the burnable absorber rods' removed. Both thimble plugging device designs will be used in Cycle 2. Table I presents a comparison of pertinent design parameters of the various fuel regions. The Region 4 fuel has been designed according to the fuel performance model(4) . The fuel is designed and operated so that clad flattening will not occur, as predicted by the Westinghouse clad flattening model(5) For all fuel regions, the fuel rod internal pressure design basis, which is discussed and shown acceptable in Reference 6, is satisfied. Westinghouse has had considerable experience with Zircaloy clad fuel. This experience is described in WCAP-8183, " Operational Experience with Westinghouse Cores."I ) Operating experience for Zircaloy grids has also been obtained from six demonstration 17x17 0FAs and four - demonstration 14x14 0FAs. This experience is summarized in the OFA transition licensing submittal.(2) l l 0827L:6 4 o . . .
[ 2.2 NUCLEAR DESIGN i The Cycle 2 core loading is oesigned to meet a Fg (z) x P ECCS limit of 1 2.15 x K(z). i Relaxed Axial Offset Control (RAOC) will be employed in Cycle 2 to enhance operational flexibility. RAOC makes use of available margin by expanding the allowable.AI band, particularly at reduced power. The RAOC methodology and application is fully described in Reference 8. The analysis for Cycle 2 indicates that no change to the safety parameters is required for RAOC operation. Adherence to thegF limit is obtained by using the FgSurveillance Technical Specification, also described in Reference 8. F g surveillance replaces the previous F surveillance by comparing a xy o measured gF , increased to account for expected plant maneuvers, to the Fg limit. This provides a more convenient form of assuring plant operation below thegF limit while retaining the. intent of using a l measured parameter to verify operation below Technical Specification i limits. F surveillance g is only a change to the plant's surveillance..
- requirements and as such has no impact on the results of the Cycle 2 analysis or safety parameters.
i 1 i Table 2 provides a summary of changes in Cycle 2 kinetics character . istics compared with the current limit based'on previously submitted accident analyses. , Table 3 provides the control rod worths and requirements at the most limiting condition during the cycle (end-of-life) for the standard j burnable absorber design. .The required shutdown margin is based on f- previously submitted accident analysis. The available shutdown margin - - exceeds the minimum required.' . The loading pattern contains 64 burnable absorber (BA) rods _ located in L 16 BA rod assemblies. Location of-the BA rods are shown'in Figure 1. P 0827L:6 5
- -- - ~ , e -E , t 4 -4
2.3 THERMAL AND HYDRAULIC DESIGN The thermal hydraulic methodology, DNBR correlation, core DNB limits, and safety analyses used for Cycle 2 are consistent with the OFA transition licensing submittal (2) No significant variations
. in thermal margins will result from the Cycle 2 reload.
The thermal-hydraulic methods used to analyze axial power distributions generated by the RAOC methodology are similar to those used 1.n the Constant Axial Offset Control (CAOC) methodology. Normal operation power distributions are evaluated relative to the assumed limiting normal operation power' distribution used in the accident analysis. Limits on allowable operating axial flux imbalance as a function of power level from these considerations were found to be less restrictive than those resulting from LOCA F g considerations. The Condition II analyses were evaluated relative to the axial power distribution assumptions used to generate DNB core limits and resultant Overtemperature Delta-T setpoints (including the f(AI) function). No changes in these limits are required for RAOC operation. 0827L:6 6
3.0 POWER CAPABILITY AND ACCIDENT EVALUATION 3.1 p0WER CAPABILITY The plant power capability has been evaluated considering the con-sequences of those incidents examined in the FSAR(3) using the previously accepted design basis. It is concluded that the core reload will not adversely affect the ability to safely operate at the design power level (Section 1.0) during Cycle 2. For the overpower transient, the fuel centerline temperature limit of 4700 F can be accommodated with margin in the Cycle 2 core. The time dependent densification model(9) was used for fuel temperature evaluations. The LOCA limit at rated power can be met by maintaining qF (z) at or below 2.15 x K(z). 3.2 ACCIDENT EVALUATION The effects of the reload on the design basis and postulated incidents analyzed in the FSAR(3) were examined. In all cases, it was found that the effects were accommodated within the conservatism of the initial assumptions used in the previous applicable safety analysis, or - the safety analysis performed in support of the OFA transition licensing submittal (2) , A core reload can typically affect accident analysis input parameters in the following areas: core kinetic characteristics, control . rod worths, and core peaking factors. Cycle 2 parameters in each of these three areas were examined as discussed in the following subsections to ascertain whether new accident analyses (in addition to the OFA analyses) were required. l i l L 0827L:6 7 1
3.2.1 KINETICS PARAMETERS i Table 2 is a summary of the Cycle 2 kinetics parameters current limits along with the associated Cycle 2 calculated values. The current limits reflect parameters used in the previously applicable safety analyses (FSAR). However, the Cycle 2 values were used in the analyses performed in support of the OFA transition licensing submittal.(2) Since the analytical basis used in the OFA transition licensing submittal (2) 33 the same for both STD and 0FA fuel, the exceptions to the current limits in Table 2 are incorporated in the new analyses for both fuel types. 3.2.2 CONTROL ROD WORTHS Changes in control rod worths may affect differential rod worths, shut-down margin, ejected rod worths, and trip reactivity. Table 2 shows that the maximum differential rod worth of two RCCA control banks moving
~
together in their highest worth region for Cycle 2 meets the current e limit. As noted in the OFA transition licensing submittal,(2) Table 3 shows that the Cycle 2 shutdown margin requirement has been changed from 1.6%ap.to 1.3%Ap. The reduced shutdown margin was shown to be acceptable by the results of the OFA transition safety analyses.(2) Table 4 is a summary of current limit control rod ejection analysis parameters and the correspondino Cycle 2 values. The current limits reflect parameters used in the previously applicable safety analysis ( FSAR) . However, the Cycle 2 values were used in the analysis performed in support of the OFA transition licensing submittal.(2) Thus, the Cycle 2 ejected rod worths have been shown to be acceptable. 3.2.3 CORE PEAKING FACTORS Peaking factors for the dropped RCCA incidents were evaluated based on - the NRC approved dropped rod methcdology described in Reference 10. Results show that DNB design basis is met for all dropped rod events initiated from full power. 0827L:6- 8
The OFA transition control red ejection analysis used peaking factors as shown for Cycle 2 on Table 4. Thus, the peaking factors following control red ejection for Cycle 2 are acceptable. The peaking factors for steamline break have been evaluated and are within the bounds of the , limits of the OFA transition licensing submittal (2) analysis. 'I J G 4 # 9 J $ 1 1 1 a 6 4 0827L:6 9
4.0 TECHNICAL SPECIFICATION CHANGES To ensure plant operation consistent'with the design and safety evaluation conclusion statements made in this report and to ensure that these conclusions remain valid, several technical specifications changes will be needed for Cycle 2. These changes are summarized below. (1) Technical Specification changes outlined in the OFA transition licensing submittal.(2) (2) Technical Specification changes given in Appendix A. 4 0827L:6' ' 10
l 1
5.0 REFERENCES
- 1. Bordelon, F.M. et. al., " Westinghouse Reload Safety Evaluation Methodology", C~P-9273, March 1978.
- 2. Dukit Power Company Transmittal to NRC, " Safety Evaluation for McGuire Units 1 and 2 Transition to Westinghouse 17x17 Optimized Fuel Assemblies."
- 3. "McG;Jire Final Safety Analysis Report."
- 4. M1111tr,J.V.,(Ed.), " Improved Analytical Model used in Westing-house Fuel Rod Design Computations", WCAP-8785, October 1976.
- 5. George, R. A. , (et. al. ), "Revistd Clad Flattening Model", WCAP-8381 July 1974.
- 6. Risher, D. ii., (et. al.), " Safety Analysis for the Revised Fuel Rod Internal Pressure Design Basis," WCAP-8964, June 1977.
- 7. Skaritka, J., Iorii, J.A., " Operational Experience with Westinghouse Cores", WCAP-8183, Revision 12, August 1983.
- 8. Miller, R. W., (et al.), " Relaxation of Constant Axial Offset
, Control-FgSurveillance Technical Specification," WCAP-10217-A, June 1983.
- 9. Hellman, J.M. (Ed.), " Fuel Densification Experimental Results and Model for Reactor Operation", WCAP-8219-A, March 1975. *
- 10. Letter from NRC, C. O. Thomas to E. P. Rahe, Jr. , Westinghouse,
" Acceptance for Referencing of Licensing Topical Report WCAP-10297-(P), WCAP-10298 (NS-EPR-2545) Entitled Dropped Rod Methodology for Negative Flux Rate Trip Plants", March 31, 1983.
0827L:6 11
TABLE 1 MCGUIRE UNIT 1 - CYCLE 2 FUEL ASSEMBLY DESIGN PARAMETERS Region 1 2 3 4* Enrichment (w/o U-235)^ 2.108 2.601 3.106 3.20 Density (% Theoretical)* 94.53 94.56 94.84 94.5 Number of Assemblies 5 .64 64 60 Approximate Burnup at++ 14750' 17200 12000 0
, Beginning of Cycle 2 (MWD /MTU)
Approximate Burnup at++ 22850 26500 23000 10450 End of Cycle 2 (MWD /MTU)
+ All fuel region values are as-built except Region 4 values which are nominal.
++ Based on EOCIA = 15150 MWD /MTU ' '
*58 fuel assemblies'are of the optimized fuel design with zircalloy grids. 2
- planned demonstration fuel assemblies are optimized fuel with Intermediate flow mixer grids and 88 removable fuel rods in each, assembly'. Th?. removable .
rods contain fuel enriched at 3.10 w/o U-235, and the remaining rods are.
. 3.20 w/o U-235.
0827L:6 - 12
l j
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l TABLE 2 MCGUIRE UNIT 1 - CYCLE 2
- KINETICS CHARACTERISTICS l ,
OFA Cycle 2 fur *ent limit ( ) Transition Limits (2) Desion Minimum Moderator- 0 +5 < 70% of RTP +5 <70% of RTP Temperature Coefficient 0 > 70% of RTP 0 >70% of RTP (ptm/ F)* Doppler Temperature -2.9 to -1.4 -2.9 to -0.91 -2.9 to -0.91 Coefficient (pcm/*F)" . Least Negative Doppler- -10.2 to -6.7 -9.6 to -6.7 -9.6 to -6.7 Only Power Coefficient, Zero to Full Power, (pcm/% power)" Most Negative Doppler -19.4 to -12.6 -19.4 to -12.6 Only Power Coefficient, -
-19.4 to -12.6 Zero to Full Power (pcm/% . power)*
Minimum Delayed Neutron .44 .44 >.44 Fraction 6,ff, (%) . Minimum Delayed Neutron .55 .50 >.50 Fraction 6,ff, (%) [ Ejected Rod at BOL) Maximum Prompt Neutron 26 26 <26 Lifetime (u sec) Maximum Differential Rod 100 100 <100 Worth of Two Banks Moving Together (pcm/in)*
*pcm = 10 -5 3, I
0827L:6 :13
i l TABLE 3 END-OF-CYCLE SHUTDOWN REQUIREMENTS AND MARGINS MCGUIRE UNIT 1 - MID-CYCLE 1 REDESIGN AND CYCLE 2 Mid-Cycle 1 Control Rod Worth (%Ao) Redesign Cycle 2 All Rods Inserted 7.38 7.06 All Rods Inserted Less Worst Stuck Rod 6.39 5.94 (1) Less 10% 5.75 5.34 Control Rod Requirements Reactivity Defects (Doppler, T,yg, - 2.99 3.22 Void, Redistribution) Rod Insertion Allowance 0.50 0.50 (2) Total Requirements - 3.49 3.72
. Shutdown Margin f(1) - (2)] (%Ao) 2.26 1.62 Required shutdown Margin (%Ao) 1.60 1.30 6
S 0827L:6 14. t c
l TABLE 4 l MCGUIRE UNIT 1 - CYCLE 2 CONTROL ROD EJECTION ACCIDENT PARAMETERS Current OFA Transition HZP-BOC Limit Limit Cycle 2 Maximum ejected rod 0.86 0.75 <0.75 worth, u p Maximum Fq (ejected) 13.0 11.5 <11.5 HFP-BOC Maximum ejected rod 0.20 0.23 <0.23 worth, up Maximum Fg (ejected) 7.1 5.3 <5.3 HZP-EOC Maximum ejected rod 0.94 . 0.90 <0.90 worth, u p Maximum Fq (ejected) 19.7 20.0 <20.0 HFP-EOC Maximum ejected rod 0.25 0.23- <0.23 worth, M p Maximum Fg (ejected) 7.1 5.9 <5.9 t 0827L:6- _ 15 -
's --i-r- = .
..__..m -
i _ . . _ _ _ _ . .
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180* R P N M L K J H G F E D C B A 4 4 4 4 4 4 4 g 4 4 4 2 4 2 4 2 4 4 4 4 4 4 4 2 4 2 3 2 2 2 1 2 2 2 3 2 4 SS 3 4 3 3 3 2 3 3 3 2 3 3 3 4
, 4 l
4 ! 4 2 3 3 3 2 2 2 3 3 3 2 4 4 4 5 4 i 4 2 2 2 3 3 3 3 3 3 3 2 2 2 4 6 4 4 2 3 2 3 3 2 3 3 2 3 2 4 4 4
' 4 7 4 2 1 3 2 3 90, 2 1 2 3 2 3 1 2 4 8
4 4 2 270' 3 2' 3 3 2 3 3 2 3 2 4 4 4 4 9 4 2 2 2 3 3 3 3 3 3 3 2 2 2 4
-10 4 4 2 2 3 3 2 2 2 3 3 3 2 4 4 4 g 4
4 3 3 3 2 3 3 3 2 3 3 3 4 _g 4 2 3 2 2 2 1 2 2 2 3 2 4 33 13 4 4 4 2 4 2 4 2 4 4 4 4 4 4 4 14 4 4 4 4 4 4 4
-15 X region number O' Y BA's "
SS Secondary Source FIGURE 1 CORE LOADING PATTERN MCGUIRE UNIT 1, CYCLE 2 16 l
APPENDIX A TECHNICAL SPECIFICATION PAGE CHANGES (In addition to proposed changes submitted in support of the OFA transition licensing submittal (2)) 4 0827L:6 -
/
F10DIFICATIONS TO 3/4.2.1 AXIAL FLUX DIFFERENCE LIttITS e O e 9 e
3 /4.2 POWEF: DISTRIBUTION LIMITS 3/4.2.1 AXIAL FLUX DIFFERENCE (AFD) LIMITING CONDITION FOR OPERATION
- 3. 2.1 The indicated AXIAL FLUX DIFFERENCE (AFD) shall be maintained within the allowed operational space defined by Figure 3.2-1.
APPLICABILITY: MODE 1 ABOYE 50 PERCENT RATED THERMAL POWER ACTION:
- a. With the indicated AXIAL FLUX DIFFERENCE outside of the Figure 3.2-1 limits, 1.) Either restore the indicated AFD to within the Figure 3.2-1 limits within 15 minutes, or 2.) Reduce THERMAL POWER to less than 50t of RATED THERitAL POWER within 30 minutes and reduce the Power Range Neutron Flux - High Trip setpoints to less'than or equal to 55 percent of RATED THERitAL POWER within the next 4 hours.
- b. THERMAL POWER shall not be f.ncreased above 50t of RATED THERMAL POWER unless the indicated AFD is within the Figure 3.2-1
, limits. 9 4 0
]
POWER DISTRIBUTION LIf!!TS
. I SURVEILLANCE REQUIREMENTS 4.2.1.1 The indicated AXIAL FLUX DIFFERENCE shall be determined to be ;
within its limits during POWER OPERATION above 50 percent of RATED - THERitAL POWER by:
- a. Monitoring the indicated AFD for ea'ch OPERABLE execre channel:
- 1. At least once per 7 days when the AFD Monitor Alarm is OPERABLE, and
- 2. At least once per hour for the first 24 hours after re-storing the AFD Monitor Alarm to OPERABLE status.
d
- b. Monitoring and logging the indicated AXIAL FLUX DIFFERENCE for each OPERABLE excore channel at least once per hour for the first 24 hours and at least once per 30 minutes thereafter, when the AXIAL FLUX DIFFERENCE Monitor Alarm is inoperable.
The logged values of the indicated AXIAL FLUX DIFFERENCE shall i be assumed to exist during the interval preceding each log-ging.
, 4. 2.1. 2 The indicated AFD shall be considered outside of its limits when at least 2 OPERABLE excore channels are indicating the AFD to be outside the limits.
4 O g l O
l a: 55 $ c
~. a.
E553
- g
(-15,100) ____(6,100) 100 UNACCEPTABLE OPERATION UNACCEPTABLE OPERATION 80 ACCEPTABLE OPERATION
.- } - . ..
60 50 F31,50) (17,50) 40 . _ _ 20 0
-50 -40 -30 -20 -10 0 10 20 30 40 50 Flux Difference (aI)%
FIGURE 3.2-1 AXIAL FLUX DIFFERENCE LIMITS AS A FUNCTION OF RATED THERMAL' POWER
it0DIFICATI0tl5 TO 3/4.2.2 HEAT FLUX HOT CHAllflEL FACTOR LIMITS 9 e e e O 9 e D l 1 i
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PG,iER DISTRIBUTION LIMITS HEAT FLUX HOT CHANNEL FACTOR-F0(Z) j LIMITING CONDITION FOR OPERATION 3.2.2 F g(z) shall be limited by the following relationships: Fg (z) 1 [ 2.15 ) [K(Z)] for P > 0.5 Fg (z) 1 [ 2.15 ] [K(I)] for P 10.5 0.o H A [0WER where P = KAir.D [a ttr.RMAL POWr_.K and K(z) is the function obtained from Figure 3.2-2 for a given core height location. APPLICABILITY: MODE 1 ACTION: ,,
, With Fg (z) exceeding its limit:
- 1. Reduce THERMAL POWER at least 1 percent for each I percent F0 (z) exceeds the limit within 15 minutes and similarly reduce the Power Range Neutron Flux-High Trip Setpoints within the next 4 hours; POWER OPERATION may proceed for up to a total of ~72 hours; subsequent POWER OPERATION may proceed provided the Overpower AT Trip Setpoints (value of K4) have been reduced at least 1 percent (in AT span) for each 1 percent FQ (2) exceeds the lim.it.-
- b. Identify and correct the cause of the out of limit condi-tion prior to increasing THERMAL POWER; THERMAL POWER may then be increased provided FQ(z) is demonstrated through incore mapping to be within its limit.
em e
POWER DISTRIBUTIOri LIMITS SURVEILLAtlCE REOUIREMEtiTS 4.2.2.1 The provisions of Specification 4.0.4 are not applicable. 4.2.2.2 Fn(z) shall.be evaluated to determine if FQ (z) is within its limit by:
- a. Using the moveable incore detectors to obtain a power distri-bution map at any THERMAL POWER greater than 5 percent of RATED THERMAL POWER.
- b. Increasing the measured Fn(z) component of the power distri-bution map by 3 percent to account for manuf acturing tolerances and further increasing the value by 5 percent to account for measurement uncertainties.
- c. Satisfying the following relationship:
M Fg (z) 1 P x Wlz) f r P > 0.5
. Fg "(z) i g j* KO ) f r P 10.5 where (2) is the measured FQ (z) increased by the allow-ances for manufacturing tolerances and measurement uncertainty, 2.15 is the Fg limit, K(z)'.is given in Figure 3.2-2, P is i the relative THERMAL P0WER, and W(z) is the cycle dependent function that accounts for power distribution transients encountered during normal operation. This function is given in the Peaking Factor Limit Report as per Specification 6.9.1.12.
M
- d. Measuring Fn (z) according to the following schedule:
- 1. Upon achieving equilibrium conditions after exceeding by 10 percent or more of RATED THERMAL POWER, the THEPMAL POWER at which Fg(z) was last determined,* or~
- 2. At least once per 31 effective full power days, whichever ~
. occurs first.
*0uring power escalation at the beginning of each cycle, power level may .
be increased until a power level for extended operation has been achieved and a power' distribution map obtained. D
POWER DISTRIBUTION LIMITS
~
SURVEILLANCE REQUIREMEtiTS (Cont)
- e. With measurements indicating maximum F (z) over z q7j has increased since the previous determination of FgM (z) either of the following actions shall be taken:
- 1. F0 M(z) shall be increased by 2 percent over that specified in 4.2.2.2.c, or
- 2. F0 M(2) shall be measured at least once per'7 effec-tive full power days until 2 successive maps indicate that
***I*"*
O IZ) is not increasing. over : ggy ( )
. f. With the relationships specified in 4.2.2.2.c above not being satisfied:
- 1. Calculate the percent Fg(z) exceeds its limit by the
, following expression: l t . . i j aximum
.F O I )
- NIZ) i
-1 > x 100 . for P > 0.5-i over z 2.15 l g xK(z)
,i : - ,
F 0 1)*N(Z)
) f[*i [ 2.15
-1 ( x 100 -
I for P < 0.5 1 ( er x K(z) i i\ , 0.5 ,,/ 4
- 2. Either of the following actions shall be taken:
- a. Place .the core in an equilibrium condition where the limit in 4.2.2.2.c is satisfie'd. Power level may -
then be increased provided the AFD limits of Figure 3.2-1 are reduced 1%' AFD for each percent Fn(z) exceeded its limit, or .
- b. Comply with the requirements of Specification 3.2.2 for F Q(z). exceeding its limit by the peretnt cal-culated above e
G y
- g. The limits specified in 4.2.2.2.c, 4.2.2.2.e, and 4.2.2.2.f above are not applicable in the following core plane regions:
- 1. Lower core region 0 to 15 percent inclusive.
- 2. Upper core region 85 to 100 percent .nclusive.
4.2.2.3 When qF (z) is measured for reasons other than meeting the requirements of Specification 4.2.2.2 an overall measured Fg(z) shall be obtained from a power distribution map and increased by a percent to account for manufacturing tolerances and further increased by 5 percent to account for measurement uncertainty. i
- b I
f e 1 I t O f ,
3 /4. 2 POWER 1STRIBUTION LIMITS 1 S The specifications of this section provide assurance of fuel integrity during Condition I (Normal Operation) and II (Incidents of Moderate Frequency) events by: (a) maintaining the chiculated DNBR in the core at or above design during normal operation and in short term transients, and (b) limiting the fission gas release, fuel pellet temperature and cladding mechanical properties to within assumed design criteria. In addition, limiting the peak linear power density during Condition I events provides assurance that the initial conditions assumed for the LOCA analyses are met and the ECCS acceptance criteria limit of 2200*F is not exceeded. The definitions of certain hot channel and peaking factors as used in these specifications are as follows: Fg (z) Heat flux Hot Channel Factor, is defined as the maximum local heat flux on the surface of a fuel rod at core elevation I divided by the average fuel rod heat flux, allowing for manu-facturing tolerances on fuel pellets and rods. N F Nuclear Enthalpy Rise Hot Channel Factor is defined as the ratio of H the integral of linear power along the rod with the highest integrated power to the average rod power. 3/4.2.1 AXIAL FLUX DIFFERENCE (AFD) The limits on AXIAL FLUX DIFFERENCE assure that the Fq (z) upper bound envelope of 2.15 times the normalized axial peaking factor is not ex-ceeded during either normal operation or in the event of xenon redis-tribution following power changes. Provisions for monitoring the AFD on an automatic basis are derived from the plant process computer through the AFD Monitor Alarm. The computer determines the one minute average of each of the OPERABLE excore detec - tor outputs ar.d provides an alarm message immediately if the AFD for at least 2 of 4 or 2 of 3 OPERABLE excore channels are outside the allowed aI . cower operating space and the THERMAL- POWER is greater than 50 percent of RATED THERMAL POWER. 3/4.2.2 and 3/4.2.3 HEAT FLUX HOT CHANNEL FACTOR, RCS FLOWRATE AND NUCLEAR EhiALPY RISE HOT CHANNEL FACTOR The limits on heat flux hot channel factor, RCS flowrate, and nuclear enthalpy rise hot channel factor ensure'that 1) the' design limits on peak local power density and minimum DNBR are not exceeded and 2) in the event of a LOCA the peak fuel clad temperature will not exceed the 2200*F ECCS acceptance criteria. limit. l , i
1
^
l PCWER DISTRIBUTION LIMITS BASES (Cont) Each of these is measurable but will nomally only be detemined period-ically as specified in Specifications 4.2.2 and 4.2.3. This periodic surveillance is sufficient to insure that the limits are maintained provided:
- a. Control rods in a single group move together with no individual rod insertion differing by more than + 13 steps from the group demand position.
- b. Control rod groups are sequenced with overlapping groups as described in Specifiction 3.1.3.6.
- c. The control rod insertion limits of Specifications 3.1.3.5 and 3.1.3.6 are maintained.
- d. The axial power distribution, expressed in terms of AXIAL FLUX DIFFERENCE, is maintained within the limits.
Fyg will be mainta ned within its limits provided conditions a. tnrough d. above are maintained. As ,noted on Figure - 3.2-3, RCS flow and FfH may be " traded off" against one another to
~ ensure that the calculated DNBR will not be below the design DNBR value. The relaxation of FNH as a function of THERMAL POWER allows changes in the radial power shape for all permissible rod insertion limits.
R N 3 as calculated in Specification 3.2.3 and used in Figure 3.2-3, accounts ! for F g less than or equal to 1.4g. This value is used in the various accident analyses where Fg influences parameters other than DNBR, e.g. , peak clad tem-perature, and thus is the maximum "as measured" value allowed. When RCS flow rate and F" g are measured, no additional allowances are necessan prior to comparison with the limits of Figure ' 3.2-3. Measurement errors of 1.7% for RCS total fiow rate and 4% for FN have been allowed for in determination of the design DNBR value.
-1 I
l
The measurement error for RCS total flow rate is based uson performing a precision heat balance and using the result to calibrate the RCS flow rate ~ indicators. potential fouling o,f the feecvatar venturi wnich mign not be detected could bias the result from the precision heat balance in a non-consenative manner. Therefore, a penalty of 0.1% for undetected fouling of tne feecwater venturi is included in Figure 3.2-3. Any fouling whien might bias the RCS flow rate measurement greater than 0.1% can be detected by monitoring and trending various plant performance parameters. If detected, action shall be taken before perfoming subsequent precision heat balance measure =ents, i.e. , either 'the effect of the fouling snall be cuantified and cc. censated for in the RCS flow rate measurement or the venturi shall be cleaned to eliminate the fouling. The 12-hour periodic survei'llance of indicated RCS flow is sufficient .o detect only flow degradation wnich could lead to operation outside the accent-acle region of oceration shown on Figure 3.2-3. When an Fg measurement is taken, both experimental error and inanu-facturing tolerances must be allowed for. 5 percent is the appropriate allowance for a full core map taken with the incore detector flux cap-ping system and 3 percent is the apprppriate allowance for manufacturing tolerance.
~
The hot channel factor F0 M(z) is measured periodically and in-creased by a cycle and height dependent power factor, W(z), to provide assurance that the limit on the hot channel factor, Fn(z), is met. W(z) accounts for the effects of nomal operation transients and was detsmined from expected power control maneuvers over the full range of burnup conditions in the core. The W(z) function for nomal operation is provided in the Peaking Factor Limit Report per Specification 6.9.1.12
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i POWER DISTRIBUTION LIMITS BASES (Cont) j 3 /4.2.4 OUADRAUT POWER TILT RATIO The quadrant power tilt ratio limit assures that the radial power dis-tribution satisfies tne design values used in the power capability anal-ysis. Radial power distribution measurementi ~are made during startup testing and periodically during power operation. { The two hour time allowance for operation with a tilt condition greater than 1.02 but less than 1.09 is provided to allow identification and correction of a dropped or mistligned rod. In the event such action does not correct the tilt, the margin for uncertainty of Fp is rein-stated by reducing the power by 3 percent from RATED THER!kt POWER for each percent of tilt in excess of 1.0. 3 /a . 2. 5 DNB PARAMETERS The limits on the DNB related parameters assure that each of the para-meters are maintained within the normal steady state envelope of opera-tion assumed in the transient and accident analyses. The limits are consistent with the intial FSAR assumptions and have been analytically
- demonstrated adequate to maintain a design limit DNBR throughout each analyzed transient. -
The 12 hour periodic surveillance of these parameters through instrument readout is sufficient to ensure that the parameters are restored within their limits following load changes and other expected transient opera-tion. 9 l l
I THIS FIGURE DELETED Firare B 3/4 21 TYP1 CAL INDICATED AXIAL FLUX CIFFERENCE VERsus THERMAL PQWER
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O ADDITIONS TO 6.0
. ADMINISTRATIVE CONTROLS O
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PEAKING FACTOR LIl1IT REPORT 6.9.1.12 The W(:) function for normal operation shall be provided to the Director, Nuclear Reactor Regulations, Attention Chief of the Core Performance Branch, U. S. Nuclear Regulatory Commission, Washington, D.C. 20555 at least 60 days prior to cycle initial criticality. In the event that these values would be submitted at some other time during 4 ' core life, it will be submitted 60 days prior to the date the values would become effective unless otherwise exempted by the Commission. Any information needed to support W(:) will be by request from the NRC and need not be included in this report. 4 6 e
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ATTACHMENT 2C PEAKING FACTOR LIMIT REPORT FOR MCGUIRE UNIT 1, CYCLE 2 This Peaking Factor Limit Report is provided in accordance with Paragraph 6.9.1.12 of the McGuire Unit 1 Technical Specifications. The McGuire Unit 1. Cycle 2 W(z) functions for RAOC operation in the cycle burnup ranges _of 0-500 MWD /MTU, 500-1000 MWD /MTU, 1000-1500 MWD /MTU,- 1500-2000 MWD /MTU, 2000-3000 MWD /MTU, 3000-6000 MWD /MTU, 6000 MWD /MTU-EOL are shown in Figures 1 through 7 respectively. W(z) was calculated using the method described in Reference 1 The appropriate W(z) function is used to confirm that the heat flux hot channel factor, F g(z), will be limited to the Technical Specifications values of: Fn (z);c 215p [K(z)] for P > 0 5 and Fq (z) < _ 4.30 [K(z)] for P < 0.5 The appropriate W(z) function, when applied to a power distribution measured l under equilibrium conditions, demonstrates that the initial conditions assumed in the LOCA are met, along with the ECCS acceptance criteria of:10CFR50.46. (1) WCAP-10216-P-A, Relaxation of Constant- Axial Control - Fg urveillance S Technical Specification 9
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FIGURE 2 ( MCGUIRE UNIT 1, CYCLE 2 RAOC W(Z) l FOR CYCLE BURNUPS BETWEEN 500 AND 1000 MWD /MTV I [ l q
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