ML20236W768
| ML20236W768 | |
| Person / Time | |
|---|---|
| Site: | Trojan File:Portland General Electric icon.png |
| Issue date: | 10/31/1987 |
| From: | Davidson S, Dzenis E WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML20236W729 | List: |
| References | |
| NUDOCS 8712080220 | |
| Download: ML20236W768 (214) | |
Text
{{#Wiki_filter:- _ - - .
,3 :- i; .< /if .r
, ? 'p 1,'
yw > , c
, .j . p.
J' 'a
.-). ,f i i
.f,', ; . { ', ,
,)
cv . 5 i i h'. ENCLOSURE-
. TO' L'
;t. C A 1 6 1- ..;.
- - NUCLEAR?FdEL UPGRADE /ECCS
- REANALYSIS.
f
; PLANT SAFETY EVALUATION
- FOR THE.~
TROJAN NUCLEAR l PLANT: FUELLUPGRADE
+-
I i
)
l I l
.s 8712000?20 871120 g PDP ADOCK 05000344 P PDR r .: ..e
__ ___ -_____ - _ - . . ]
i ENCLOSURE
- to LCA 161 .
Nuclear Fuel Upgrade !
/ECCS Reanalysis' {
( . l PLANT SAFETY EVALUATION FOR THE 1 1 TROJAN NUCLEAR PLANT FUEL UPGRADE OCTOBER .1987 ; i
- S. L Davidson, Editor i
Approved: E- WS E. A. Drenis, Managef Core Operations ; i I
]
1 l l WESTINGHOUSE ELECTRIC CORPORATION Commercial Nuclear Fuel Division P. O. Box 3912 Pittsburgh, Pennsylvania 15230
\
- I , \
l 870lL6-471020
..__2._ _ .- )
b PLANT SAFETY EVALUATION for'the TROJAN NUCLEAR' PLANT FUEL UPGRADE TABLE OF CONTENTS TITLE PAGE SECTION 1.0 Introduction and Summary 1-1 2.0 Design Features 2.1 Introduction 2-1 2.2 Reconstitutable Top Nozzle 2-1
~
2.3 Axial Blankets 2-2 2.4 Extended Burnup 2-2 3.0 Nuclear Design-3.1 Introduction and Summary 3-1 3.2 Methodology 3 3.3 Design Evaluation - Physics Characteristics 3-2 and Key Safety Parameters 3.4 Design Evaluation - Power Distribution and Peaking 3-3 Factors 3.5 Technical Specification Changes Relative to Nuclear 3-4 Design 3.6 Nuclear Design Evaluation Conclusion 3-4 4.0 Thermal and Hydraulic Design i 4.1 Introduction and Summary 4-1 ! 4.2 Calculational Methods 4-1 4.3 Hydraulic Compatibility- 4-2 / 4.4 Effects of Fuel Rod Bow on DNBR 4-2 l 4.5 DNBR Effect on the. Upgraded Fuel 4-2 lE 4.6 Fuel Temperature for Safety Analysis 4-3
. ima-enon 3$
L_ _ - _
n . i! , c.
. TABLE OF' CONTENTS (Cont.)l SECTION TITLE ' PAGE r . 5.0 Accident Analysis-
'5.1 Non-LOCA Accidents 1' 5.2 LOCA Accidents- 5-10 1
6.0 . References 6-1 i APPENDIX Impact of Extended Fuel Bur ~up on Radiological Consequence.of Accidents 4 i l 9 j i 1 1
' 570sL.4-471020 jjj
. J
' LIST'0F FIGURES
~
TITLE PAGE FIGURE 2.1 Bottom Nozzle to Thimble Tube Connection 2-4 3.1 Cycle 11 Quarter Core Fuel Loading and- 3-7 Burnup (MWD /MTU) Distribution 3.2 Cycle 11 Normalized Power Density Distribution 2-8 Near Beginning of Life, Unrodded Core, Ho't Full ; Power, Equilibrium Xenon
.3.3 Cycle 11 Normalized Power. Dens'ity Distribution- 3-9 Near Beginning 'of Life, . Bank D at Insertion Limit, Hot Full Power, Equilibrium Xenon 3.4 Cycle 11 Normalized Power Density Distribution 3-10 Near Middle of Life, Unrodded Core, Hot Full Power,' Equilibrium Xenon 3.5 Cycle 11 Normalized Power Density Distribution 3-11 Near End of Life, Unrodded Core, Hot Full -{
Power, Equilibrium Xenon 3.6 Cycle 11 Normalized Power Density Distribution 3-12 C Near End of Life, Bank D at Insertion Limit Equilibrium Xenon 3.7 versus Axial Core 3-13 .] Maximum F0 .PREL Height During Cycle 11 Normal Core Operation 4 l l l J smu-ume gy - _ -_____-________a
LIST OF TABLES TITLE PAGE. TABLE' 3.1 Representative Nuclear Design Parameters and Cycling 3-5 Scheme 3.2 Range of Key Safety Parameters 3-6 4.1 Trojan Thermal' and Hydraulic Design Parameters 4-4 DNBR Margin Sunnary 4-8 ' . 4.2 I
$708LS-471020 - y
\
f .;. . .j ;. I N -- r .,; s l' . 1
.7 l' =
1.0 INTRODUCTION
AND'
SUMMARY
i f l i 1 1 l l l 1 b
,f N i
. S:
,j ,
.s I.
.. i
1.0 INTRODUCTION
AND
SUMMARY
0 , . . It is planned to refuel and coerate the Trojan Nuclear Plant with Westinghouse i upgraded. fuel Lfeatures. As a result, future core loadings will have fuel assemblies consisting of Reconstitutable Top Nozzles (RTNs)i and axial blankets-with the capability.of achieving extended discharge burnups wh'ile satisfying both a . higher peaking- factor and higher LOCA peaking factor limit than -l previous reload de' signs. The upgraded fuel features are'a part of the VANTAGE-5 design features generically approved by the NRC,via their review of the
" VANTAGE 5 Reference Core Report," WCAP-10444-P-A, Reference'1. A brief summary of the VANTAGE 5 design features and major advantages.of the upgraded fuel design are given below.:
Reconstitutable Top Nozzle (RTN) - A mechanical disconnect feature facilitates the. top nozzle removal. Changes in the design of both the top and
-bottom nozzles increase burnup margin by providing additional plenum space and l room for fuel rod growth.- ;
l ' Axial Blanket - The axial blanket consists of natural UO2 pellets at each end of the fuel stack to reduce neutron leakage and to improve uranium utilization. For the upgraded reload cores, low leakage loading patterns (burned radial blankets) are shown to further. improve uranium utilization and provide additional pressurized thermal shock margin. Extended Burnuo - The fuel upgrade will be capable of achieving a lead rod . average burnups up to the maximum lead rod burnup permitted by the approved )
- Westinghouse Topical, WCAP-10125-P-A, Reference 2. l l The Trojan Plant Safety Evaluation (PSE) is to serve as a reference safety i i ' evaluation / analysis report for the region-by-region reload transition from the l present Trojan core (Cycle 10) to an entire core containing the above upgraded fuel features. The analyses were performed at a core thermal power level of 3411 Megawatt thermal (MWt) with the following conservative assumption made in the safety evaluation
- a nuclear enthalpy rise hot channel factor (FAH) of 1.62, the LOCA analyses performed at an uprated core thermal power level of 1-1 4013P:6-671022
WVW ' 3 .
+
r" y '} i ll- . . . . < , 1 " .y 9 J j;
- v l
1 i j
< 102% of 3558 MWt with.an increase. in the maximum t.0CA10 to 2.50 with 3rd l 11.ine' segment to l'.5, and a ~.1.5% plant total steam generator tuce plugging level.
i l ThePSE.utilizesthestanoard'reloaddesignmethodEdescribed'in~ Reference 3 ) and will .be'used as' a basic. reference document:in' support of future Trojan Reload Safety Evaluations,(RSEs) for: upgraded fuel reloads. Section 2.0 through 5.0 of the PSE' summarize the' Mechanical,: Nuclear. Thermal and-Hydraulic. . and the Accident Evaluation, respectively. !
- , 1
]
The radiological consequences due to the-extended-burnup operation are' described in the appendix, Consistent with the Westinghouse. standard reload methodology, Reference 3, p'arameters'are chosen to' maximize the applicability of the PSE evaluations for
~
future cycles. The objective of subsequent. cycle specific RSEs will be to verify that applicable safety limits are satisfied based on the reference evaluation / analyses established in this safety evaluation. In order to demonstrate early performance of the VANTAGE 5 design product features that included the RTNs, axial blankets, and extended burnup
~
L capability, four demonstration assemblies (17x17) were loaded into the V. C. Summer Unit 1 Cycle 2 core and began power. production in Decemoer of 1984. l; These assemblies completed one cycle of irradiation in October of 1985 with an average 'burnup of 11,357 MWD /MTU. Post-irradiation examinations showed all 4 demonstration assemblies were of good mechanical integrity. No mechanical damage or wear was evident on.any of the VANTAGE 5 conconents including the RTN and axial blankets. All four demonstration assemblies were reinserted into V. C. Summer i for a second cycle of irradiation. This cycle was completed in March of 1987, at which time the demonstration assemblies achieved an average burnup of about 30,000 MWD /MTU. The observed behavior of ll the four demonstration assemblies at the end of 2 cycles of irradiation was as good as that observed above at the end of the first cycle of irradiation. The four assemblies were reinserted for a third cycle of irradiation. 1-2 40138 6-4P1022 L. --. . _ _ _ _ _ _ _ _ . _ _ _ - _ - - _ _ - _ _ _ _ _
.!;~
l' , .. f-:. n l ,
-The results of evaluation /analysi.s described herein lead te the following-conclusions:
- 1. The Westinghouse'fue' assemblies containing upgraded fuel features for the Trojan plant are mechanically compatible with the current' fuel assemblies, control' rods, and reactor internals interf aces. The g
upgraded' fuel assemblies satisfy the current design bases.for the
. Trojan plant.
- 2. Changes in the nuclear characteristics due to the transition to upgraded fuel will be within the range normally seen from cycle to cycle due to fuel management effects.
- 3. The reload upgraded 'luel assemblies are hydraulically compatible with the fuel assemblies from previous reload cores.
- 4. The core design and safety analyses results documented in this report show the core's capability for operating safely for the rated Trojan design thermal power with a F AH f 1.62, Fg = 2.50, and steam generator tube plugging levels up to 11.5%.
- 5. Previously reviewed and licensed safety limits are met when the Trojan plant is reloaded with upgraded fuel as described in this section.
c Plant operating limitations given in the Technical Specifications will be satisfied with.the proposed changes, i I l l 1-3 4013F 6-471022 - = -_._L_.2-. ._
_. ,_, -. ~ --.
.7--.-
s.
,y L '.
h-t 2.0.. DESIGN FEATURES' e 4 4 ? l l \ [ u - - -- _- --- _ -- _ -
L i 1 2.0 DESIGN FEATURES 2.1' Introduction The Trojan upgraded fuel assemblies are Westinghouse 17x17 standard 12 foot assemblies. They differ from previous reload fuel assemblies in that they have'Reconstitutable' Top Nozzles'(RTN), modified bottom' nozzles, axial' blankets, and extended burnuo capabilities. In conjunction with the RTN, the upgraded fuel assemblies possess long tapered end plugs in the fuel rods to allow for easy removal and reinsertion of individual fuel rods.'
- 2.2 Reconstitutable Top Nozzle .
In the Reconstitutable Top Nozzle design, a stainless steel nozzle insert is mechanically connected to the nozzle adapter plate by means of a pre-formed circumferential bulge near the top of the insert. The insert engages a mating groove in the wall of the adapter plate thir61e tube thru-hole. The insert has four equally spaced axial slots which allow the insert to deflect inwardly 4 at the elevation of the bulce, thus permitting the installation or removal of the nozzle. The insert bulge is positively held in the adapter plate mating groove by placing a lock tube with a uniform inner diameter identical to that~ of the thimble tube into the insert. The lock tube is secured in place by two means. First, a top flare creates a tight fit. Second, six non yielding projections on the outer diameter which interface with the concave side of the insert preclude escape during core component trahsfer. Reconstitutable Top Nozzle attachment schematic details are given in Figure 2.4 of Reference 1. The full compierent of these joints comprises the structural connection (reconstitutable design feature) between the top nozzle and the remainder of the fuel assembly. The nozzle insert-to-adapter plate bulge joints replace the uppermost grid slaeve-to-adapter plate welded joints found in current fuel assemblies. The nozzle insert-to-thimble tube multiple 4-lobe bulge joint located in the lower portion of the insert represents the structural connection between the insert and the remainder of the fuel assembly below the elevation of the insert. The uppermost grid sleeve is connected to the thimble tube by similar 4-lobe bulge joints. 2-1 aos-mm
m.- n To remove the top nozzle, a tool is first -inserted through the lock tube and - expanded radially to engage zhe bottom and of the tube. An axial force is then exerted on the tool wnich overrides the local lock tube deformations and withdraws the locktube from the insert. After the lock tubes have been withdrawn, the nozzle is removed by. raising it off the upper slotted ends of the nozzle inserts which deflect inwardly under the axial lift load. With the top nozzle removed,' direct access is provided for fuel rod examination or replacement. Reconstitution is completed by the remounting of the nozzle and the insertion of new lock tubes. 2.3 Axial Blankets The upgraded fuel rods contain axial blankets and utilize a chamfered pellet physically different than the enriched standardized' pellets in the fuel stack. The axial blankets of natural uranium are at each end of the fuel pellet. stack. . Axial blankets reduce neutron leakage and improve fuel utilization. The physical differences in the natural pellet will help prevent accidental mixing with the enriched standardized pellets. Five reload cycles of operation with axial blankets have been completed on Westinghouse cores and no anomalies have been observed. 2.4 Extended Burnup Extended burnup capability fuel licensed in Reference 2 employs a reconstitutable bottom nozzle and a slightly longer fuel rod. This bottom nozzle is shorter and has a thinner top plate than previous reload assembly designs in order to accommodate the longer fuel rod. I The bottom nozzle reconstitution feature facilitates easy removal of the nozzle from the fuel assembly. The reconstitutable bottom nozzle design incorporates a thimole screw with a circular locking cup located around the screw head. The locking cup is crimped into mating detents (lobes) on the j l bottom nozzle. To remove the bottom nozzle, a counterclockwise torque is applied to the thimble screw until the detents of the cup are overpowered and the thimble screw removed, i 2-2
<n y .-.m 4
l' .y'. >- + . This reconstitutable design oermits remote unlocking, removing, and relocking of the thimble' screws as a new or the same bottom nozzle is re-attached to the fuel assembly. - See Figure E.1 for a comparison of the' conventional and reconstitutable bottom nozzie designs. The RTN design-feature, the Reconstitutable Bottom Nozzle, and axial blankets
-are described in Reference I which has oeen approved by the NRC.
I so1ve-enon 1
T
-THIMBLE SCREW I4CKWIRE
. /
THIMBLE TUBE
- WEIo i
~
( g NS_\\ p ! , i h _t _ - - 2.
- - - -- u- ,
\ j
\
% %%V/4/ J l
BOTTOM NOZZLE .LOCKWIRE l CONVENTIONAL CONNECTION CRIMPED LOCKING CUP THkMBLESCREW THIMBLE TUBE IDCKING CUP
/
2Y
-S
~
~ Cl=25 +-
-= ==( _ . . _ _
Q') ( ~X AN%Nw/4/ l
.j BOTTOM NOZZLE /
\'
RECONSTITUTABLE CONNECTION f 1 TROJAN UNIT 1 ] FIGURE 2.1 I BOTTOM NOZ:LE TO THIMBLE TUBE CONNECTION 2-4 w---_-__-__
l',^ 1 I e i 3.0 NUCLEAR DESIGN t g .
=
i. i
/
b ___ .
4 9_ 3.0 ' NUCLEAR DESIGN 3.1 tatroduction and Summary The nuclear, design portion of, the Trojan Plant Safety Evaluation'(PSE) has- two' q objectives. First', the' impact on the key safety parameters due to tho' s upgraded fuel product will be evaluated.. .These safety parameters are used as'
> input to'the FSAR Chapter 15 accident analyses. Second, the plant Technical Specifications that. apply to nuclear design must be reviewed to determine ~if they remain' appropriate or must be altered t'o accommodate a core containing
'the upgraded fuel product.-
To satisfy these objectives, a representative core model which contained the upgraded fuel product was coveloped using fuel management typical of previous l Trojan cycles. Key safety parameters were then evaluated to determine the expected ranges of variation of these parameters. The key safety parameters referred to here are those des:ribed in the standard reload design-methodology, Reference 3. The majority of these parameters are-insensitive to fuel type and are primarily loading pattern dependent, e.g., control rod worths and peaking factors. The observed variations in these loading pattern (LP) dependent parameters for the core containing the upgraded fuel product are typical' of the normal cycle to cycle variations for standard fuel reloads. l q A Trojan core containing the upgraded fuel product will necessitate some Technical Specification changes as a result of increased peaking factors, , F g and F . The increased peaking factor limits are needed because of O ' the axially heterogeneous features of the upgraded fuel and will serve to reduce burnable absorber requirements, improve fuel economy, and increase nuclear design flexibility. In summary, the change from the current all standard fuel core to a core containing the upgraded fuel product will not cause changes to the current nuclear design bases given in the Trojan FSAR. The evaluation of the Trojan upgrade demonstrates that the impact of implementing the upgraded fuel product does not cause a significant change to the physics characteristics of the
. Trojan core beyond the normal range of variations seen from cycle to cycle.
i mu-.mn 31 : _ _ _ _ _ _ _ _ _ _ _ J
' 1
)
3.2 Methodology The methods and core models used in the Trojan fuel upgrace analysis are described in References 1, 3. and 4. These licensed methods and models have been used for Trojan and other previous Westinghouse reload designs. Improved nodal analysis methods described in Reference 5 will also be used in future i cycle specific Trojan reload design analysis. No changes to the nuclear j design philosophy, methods, or models are necessary because cf the upgraded fuel product. Increased emphasis will be placed on the use of three- l dimensional nut. lear models because of the axially heterogeneous' nature of the I fuel design when axial blankats are used. The reload design philosophy employed includes the evaluation of the reload core key safety parameters wHch comprise the nuclear design dependent input to the FSAR safety evaluation for each reload cycle. This philosophy is described in References 1 and 3.: These key safety parameters will be evaluated for each Trojan reload cycle. If one er nore of the key parameters fall outside the bounds assumed in the safety analysis, the affected transient will be re-evaluated and the results dccumented in the RSE for that cycle. One objective of the Trojan upgrade analysis is to determine, prior to the cycle specific reload design, if the previously used bounds for these key safety parameters will continue to remain applicable. The results of this upgrade core analysis are described in Section 3.4. 3.3 Design Evaluation - Physics Characteristics and Key Safety Parameters As previously mentioned, a representative model containing the upgraded fuel was generated for nuclear design evaluation. Table 3.1 summarizes the cycle length, the fuel leading, and provides some representative nuclear design data. Table 3.2 compares the safety parameter ranges obtained in this upgrade analysis with those obtained for the Cycle 10 design.
'"*""** 3 - 2 ___ __ _
s? Y 3.4 Design Evaluation - Fewer Distributions and Peaking Factors
. The implementation of axial blankets will have-impacts on core power.
distributions and peaking factors experienced in Trojan. The use of axial-blankets, where the top and bottom of the enriched fuel stack are replaced by: natural uranium pellets and the enrichment of--the remaining fuel is increased' slightly, results in higher axial peaking factors. . The increased radial peaking factor limit allows further implementation of low leakage' fuel management by placing more burned fuel on the periphery of the core. The reduction in power carried.by the peripheral assemblies is offset by increases in power in the remaining assemblies. 'This increased radial peaking is accommodated by increasing the radial peaking factor limits, Fg. For Cycle 11, a representative loading pattern was developed and modeled based
'on the projected energy requirement and other available .information. Results of calculations show an increase in radial peaking from previous. cycles which' is not unexpected. It results from the reduced power carried by.the more highly burned assemblies placed on the core periphery to reduce neutron leakage as well as the advent of blanketed fuel which reduces power at the extreme top and bottom of the fuel, thereby reducing axial leakage.
Figure 3.1 shows the quarter core loading pattern and assembly burnups for Cycle 11 (one region of upgraded fuel.'. The assembly burnups are for the beginning of life (BOL), no xenon. Figures 3.2 through 3.6 give the quarter core power distributions at hot full power (HFP), all rods out (ARO) for beginning-of-life (BOL), middle-of-life (MOL) and end-of-life (EOL), and HFP D-Bank inserted to the insertion limit for BOL and EOL. The' assembly average l powers are typical values and do not necessarily represent bounding values for future designs. These power distributions are for illustrative purposes and do not represent the full scope of the nuclear analysis that was performed. Figure 3.7showsthetotalpeakingfactor,Ff*PRei, versus core height resulting from the Cycle 11 total peaking f actor evaluations. Constant axial offset control (CAOC) operation with a +5, -5% delta I band assuming load follow operation formed the basis for the Fg evaluations. e.-man 3-3 l
? :j .
g. i 3.5 Technical Specification Changes Relative to Nuclear Design l \ l There are two major areas related to nuclear design that will affect Technical
' Specifications: increase in 2 3g limits and. increase in F 0 limits. The limits will be increaseo.for the standard fuel as well as the upgraded!
~
F 3g fuel. The~ limit function would be: N Y AH
.g ,
R 1.56 [1.0 + 0.3 (1-P)J where P is the fraction of full power. WhenFfg is measured, no additional allowances are necessary prior to comparison with the' limits. A measurement error 'of 4% for FN H : has been allowed for .in the determination of the design DNBR value. The F limit will-be increased from 2.32 to 2.50. This will' permit core O designs to be performed with minimal use of burnable absorbers.
.3.6 Nuclear Design Evaluation Conclusion The-key. safety parameters evaluated for the conceptual-design shows that the expected ranges of variation for many of the parameters will lie within the normal cycle-to-cycle variations ob' served for reload designs.
Power distributions and peaking factors show some changes as a result of the incorporation of axial blankets, and increased peaking factor limits, in addition to the normal variations experienced with different loading patterns. The usual methods of loading pattern shuffling and enrichment variation can be employed in future cycles using the upgraded fuel to ensure compliance with the peaking factor Technical Specifications. In summary, Technical Specification changes will be limited to increases in
.F 3g and F g
. Outside of these specific areas there will be no further fuel upgrade related technical specifications required by nuclear design for the Trojan Plant Safety Evaluation.
. smoce-ena: 34
1 Table 3.1 REPRESENTATIVE NUCLEAR DESIGN PARAMETERS AND CYCLING SCHEME Representative' i Nuclear Design Parameters Parameter Value in Each Cycle l 10 11 Cycle Burnup (MWD /MTU) 10100 10000 Core Power (MWt) 3411 3411 Uranium Loading (MTU) 89.10 89.10 Cycle Length Equivalent Full Power 264 261 Days (EFPD) No. of Fresh Burnable Absorbers none none HFP Critical Boron (ppm) BOL, No Xenon 1405 1437 BOL, Equilibrium Xenon 1067 1093
-0.81 -1,68 Moderator Teacerature Coefficient (MTC)
BOL, HZP, No Xenon (pcm/Deg F) Uncoated Fuel Pellet U-235 Region Fuel Diameter Enrichment Number of Assemblies Number Lpa (inches) (w/o) 10 11 3 STD .3225 3.09 8 8 4 STD .3225 3.10 1 9 6 STD .3225 3.20 - 4 7 STD .3225 3.20 8 8 8 STD .3225 3.29 40 - 9 STD .3225 3.31 12 - 10 STD .3225 3.45 44 36 11 STD .3225 3.42 32 32 12 STD .3225 3.40 48 48 13 Upgraded .3225 3.45/.711 - 48 ma+ nm 3-5
l [ Table 3.2' 1
. Range of Key Safety Parameters Trojan Trojan Cycle 10 Cycle 11 Safety Parameter Standard Upgrade -
3411 3411 Reactor Core Power (MWt) Vessel Average Coolant Temp. 586.1 586.1 HFP (Deg F) Coolant System Pressure (psia) 2250- 2250 Core Average Linear Heat Rate 5.44 5.44-(Kw/ft) Most Positive Moderator Temperature 5 5 Coefficient (MTC) (pem/0eg'F) 1 Most Positive Moderator. Density .43 .43 Coefficient (MDC) (Ak/g/cm3 ) Doppler Temperature Coefficient -1.0 to -2.2 -1.0 to -2.2 (pcm/*F) Doppler Only Power Coefficient
- Least Negative (pcm/% Power) -6.68 to -~10.18 -6.68 to -10.18
- Most Negative (pcm/% Power) -12.6 to -19.4 -12.6 to -19.4 Beta-Effective .0044 to .0075 .0044 to .0075 Shutdown Margin (% delta-rho) 1.60 1.60 1.30 (pending) 1.30 (pending) 1.435 1.500 Normal Operation F3g (without Uncertainties) smou-eno" 3-6
j } .'
- l. .)
,ix J
l. l' l
.1 11 12 12 10 13 11 .
~
- 4. 12 -
~
28788,: 12238 22747 13818 7878 29905 0; 28881. 12 12 10 11 13 4 12 11 20757 13072 13321 28032 29182 0 31315 12238 12 11 12 12 11 13 to 11 9900 29888 0 33018' , 22747 13000 22991 13085 I 12 12 12 10 3 13 8 12 13818 13304 13078 12222 24979 30400 0 29878 10 12 10 8 13 - 13~ 12 , 25144 9800 25211 31802 0 0 7978 10 11 11 10 13 13 to 28805 29188 29881 34214 0 0 38217 13 13 13 13 13 10 0 0 0 0 0 38221 AS ASSSNSLY SURNUP 11 4 3 S , I8 3157S 30848 28704 29881 i RSSION 895t!OSISNT TYPS 3 3.088 STD 4 3.100 STD 8 3.108 STD 10 3.452 STD 11 3.410 STD 12 3.400 STD 13 3.400 /.711 UPSRA080 TROJAN UNIT 1 l FIGURE 3.1 l S *NN/Mhbis tYutl$n* 3-7 _ - _ . . - _ - _ _ _ - _ _
4 0.940 1.215 1.154_ 1.285 , 1.257 0.958 .1.303 0.494 - f
- f 1.215 1.151 1.273 1.283 1.038 0.973_ 1.290 0.437
* \
1.154 1.277 1.154 1.270 :1.188 0.888 1.243 0.435 .
.j 1.285' 1.283 1.288 '1.194 0.934 0.853 1.141 0.382 9
1.257 1.035- 1.188 0.937 0.874 1.340 0.979 ! 0.999 0.573 0.964 0.871 -1.345 1.178 0.415 1.303 1.283 1.238 1.141 0.984 '0.417 AP ASSEMBLY POWER 0.494 0.427 0.410 0.340 u
-I i
)
TROJAN UNIT 1 FIGURE 3.2 Cycle 11 Normalized Power Density Distribution ; innm of ynrodded Core, nearHotBYull P8wer,lifE ilibrium Xenon 3-8 _ _ _ _ . _ _ _ _ _ __ __ i
I l I i (
,, 0.893 1.196 1.145 1.200 1.205 0.950 1.317 0.501 1.198 1.141 1,388 1.274 1.022 0.873 1.306 0.443 1.145 1.271 1.157 1.299 1.188 0.371 1.280 0.442 1.288 1.274 1.287 1.192 0.834 0.857 1.157 0.357 1.205 1.021 1.187 0.931 0.850 1.345 0.990 0.950 0.973 0.988 0.875 1.350 1.185 0.419
,WI 1.317 1. 2W 1.253 1.1ST O.985 0.421 AP ASSEMBLY POWER 0.501 0.433 0.418 0.354 TROJAN UNIT 1 i l
FIGURE 3.3 i Cycle 11 Normalized P r ensity Distribution ] near Beginnihg
' Hot ull life. ank ofPower. at insertion Limit quil brium Xenon 3-9
l 6 I l n _ 0.983 1.203 1.138 1.244 1.244 0.979 1.287 0.532 1.203 1.139 1.240 1.249 1.083 0.988 . '1.250 0.475 '; (: mummmmme 1.138 1.243 1.142 1.245 1.199 0.'995 1.217 .0.479 i 1.248 1.249 1.244 -1.198 0.082 0.888 1.129 .0.344 1.248 1.082' 1.190 0.379 0.902 1.285' ~0.982 0.979 0.998 0.983 0.901 1.287 1.128 0.437 1.287 1.248 1.212 1.127 0.984 0.438 AP ASSSMSLY POWER 0.532 0.488 0.481 0.388 l l TROJAN UNIT 1 FIGURE 3.4 Cycle 11 Normalized Power Density Distribution neer Middle of life, Unrodded Core, 1 Not Full Power, Equilibrium Xenon f 3-10
( 1 0.979' 1.199 1.129 . 1.231- 1.240 . 0.992 ~1.232 0.555 1.199 1.138 1.225 1,233 1.001 1.011- 1.227 0.500 1.129 1.229 .1.134' 1.233 1.201 1.010 '1.200 0.901 1 1.231' 1.238 1.231 1.190 1.004 -0.900 1.117. 0.4tV 1.240 1.001 1.200 1.000 0.920 1.254 0.990 l j.- ' O.992 1.011 1.007 0.917- 1.254 1.099 0.450 1.232 1.222 1.195 1.113 0.951 0.451 AP ASSEMBLY POWER 0.sss 0.491 0.477 0.40s J l l i. TROJAN UNIT 1 FIGURE 3.5 Cycle 11 Nor alized Power Density latribution near Unrodded ore. . Hot F indPower, of life!quilibrium
! enon C. .-!.. ..
3-11
0.938 1.184 -1.126 1.218 1.188 0.983 1.248 0.544 1.184 1.130 1.225 1.228 1.050 1.012 1.242 0.508 1.227 1.137 1.235 1.203 1.017 1.216 O.508 1.128 3-1.216 1.228 1.233 1.198 0.998 0.904 1.131 0.418 1.188 1.048 1.201 0.992 0.884 1.249 0.957 k l 0.883 1.012 1.014 0.810 1.248 1.100 0.452 l l 1.24e 1.23s 1.211 1.127 0.ssa 0.453 j AP ASSEMSLY POWER 0.5s4 0.4ee 0.485 0.415 TROJAN UNIT 1 FIGURE 3.6 Cycle 11 Normalized Power Density Distribution near End Hot of life. Full BankEquilibrium Power, _C at insertion XenonLimit 3-12 ;
n . 2.6 , ,,,,l , ,, ;;g; ,3; ,,, ,, , ii i Illi i ii iii ill iii ; i L i' i il I ilL atf' % U l
'<r a I ;i .e :s ti Ill ili F *. 4 P% il 1 24 t~ ag .6' ii l ; I ' la ,i l ,
'-Qi i il
~o 1 il i I i 11 1 1 , a -.I l l-ll ! l
- 6 \ l I I! I i i in 'l\ l I
-li I I i! I I !' Ii d i i!
1 i i I I I i l l V 11 I \ 2.0 , , ,, , , 3,, 11 \l il 1.1 li i *U
$8 I \1 8 a FXY LIWilS i l' E : l\
x I UNRODDED I I IL 31.6 1.90 0 10 4.25 FEET I ia i
- 1.86 4.25 10 12.0 FEET 1 l
1 i l RODDED i l 14 1.98 0 10 12.0 FEET t l 12 1 Il li I I ,ll 1111 I I i 1 I .I l l i 'J l l 1 !l l l 1 1 II
! l l lt i i Itil 1.0 ,i ; ; , , ,,, , , , ,,,,
j l l i'l 1 ;l ti'l 3 l ! l lll l 1 ';ll J
! ! ,li i til i I ; I itil
.80 0 1 2 3 4 5 6 7 8 9 10 11 12 i 1
CORE K IGHT (FEET) TROJAN UNIT 1 FIGURE 3.7 Maximum r of Pan versus Axlat Core Height
, during Cycle 11 Normal Core Operation l 3-13
4.0 THERMAL AND HYDRAULIC DESIGN l l l q l 4 i 5
.K .-.l.. -..__!.. --._.- _ . - . . - _ . _ _ _ _ _ _ . . .
1
~
4.0 THERMAL AND HYDRAULIC DESIGN 4.1 ' Introduction and Summary This section describes the thermal-hydraulic analyses performed.to support the
. . use of thel upgraded' fuel in the Trojan Nuclear Plant. The thermal-hydraulic design for the upgraded fuel product was analyzed for a radial peaking factor limit (FAH) increase from 1.49 to 1.56. The increase in design F AH was accommodated in design margin available in the safety; analysis Departure from Nuclear Boiling Ratio-(DNBR) values such that there was minimal impact on the core limits. Table 4.1. summarizes the thermal-hydraulic design parameters used in these ' analyses. The thermal-hydraulic design criteria and methods
- remain the same as those presented in the Trojan Final Safety Analysis Report (FSAR) with the exceptions noted in the following section. All of the current-thermal-hydraulic design ~ criteria are satisfied.
4.2 Calculational Methods The analysis 'of the upgraded fuel product is based on the Improved Thermal Design Procedure (ITDP), Reference 6, and the Westinghouse Critical Heat Flux (WRB-1) correlation, Reference 7, as described in the Trojan FSAR. With ITDP, the plant parameter uncertainties are used to determine the plant DNBR uncertainty, which when combined with the correlation limit ~ (1.17), establishes a design DNBR value. The design DNBR value must be met by the l plant safety analyses. Since the correlation limit and plant parameter uncertainties are unchanged for the upgraded fuel, the design DNBR values in FSAR Section 4.4.1.1 are applicable, l As described in the FSAR, a specific plant allowance is considered in the DNBR values employed in the safety analyses. The available design DNBR margin incorporated in the safety analysis DNBR limits is defined in the following equation: ! ( Desion Limit DNBR Safety Analysis DNBR = 1.0 - Margin 4-1 j useu -emn 1
Currently, a portion of the design DNBR margin has been allocated to offset DNBR penalties associated with rod bcw and the' increase in best estimate l bypass flow due to upflow conversion (see Table 4.2). A reduction to'the rod- # bow penalty is discussed in Section 4;4. For the upgraded fuel, the effect of the increase in FAH on DNBR was
. accommodated by reducing'the amount of margin between the design'value and the
' safety analysis value as described in Section 4.5. 4.3 Hydraulic Compatibility The upgraded fuel product is hydraulically identical to the standard fuel currently in Trojan. No transition penalty is required.
~
4.4 Effects of Fuel Rod Bow on DNBR
' The phenomenon of fuel rod bowing must be accounted for in the DNBR safety analysis of Condition I and Condition II events. Currently, a rod bow penalty of 10.6% is assessed as described in the Trojan Technical Specifications.
Based on References 8, 9, and 10, the maximum rod bow penalty for 17x17 standard fuel is <1.5% at an assembly average burnup of 14000 MWD /MTU. For burnups greater than 24000 MWD /MTU, credit is taken for the effect of'F3g burndown due to the decrease in fissionable isotopes and the buildup of fission product inventory. Therefore, no additional rod bow penalty is required at burnups greater than 24000 MWD /MTU. , 4.5 DNBR Effect of the Upgraded Fuel The change to the Trojan Technical Specifications, due to the upgraded fuel which impacts DNBR, is the increased design radial peaking factor as defined by the following equations: N F H = 1.56 [ 1.0 + 0.3 ( 1.0 - P) ], and :
.j.
pN sH FR* 1.56 [1.0 + 0.3 (1.0-P)] 4-2 : samm s-ena l e-_ _ __ . __
n_ . - _ _ i.
. where P i fractional core power level at less than 100% rated power' or,
= 1.0 for core power level greater than or equal to 100% rated power.
The radial peaking factor limit increase from 1.49 to 1.56 has a' direct impact on DNBR calculations. The impact on DNBR due to the increased F3g was offset by a reduction in the margin which defines the DNBR value for the safety analyses. The new safety analysis DNBR values were selected such that there were minimal changes in the core limits. In addition to the FAH increase, the new safety limit DNBR values explicitly account for the increased core bypass flow associated with the conversion of the reactor vessel internals from downflow to upflow.. A summary of the DNBR limits, margins, and penalties ~for the upgraded fuel is presented in Table 4.2.' The
^
impact of the minor changes.to the core limits on the non-LOCA analyses is l addressed in Section 5.0. associated with the The axial blankets and the increased allowable F0 upgraded fuel product affect the axial power distribution and therefore, the 1 DNBR analyses. These effects were accounted for in the non-Overtemperature Delta T (non-0 TAT) DNBR analyses by means of a limiting axial power ; I distribution. The impact on OTAT accident analyses is discussed in Section 5.0. l 4.5 Fuel Temperatures for Safety Analysis Thefueltemperatures(asafunctionoflinearheatbgrate)foruseinsafety analysis calculations for the upgraded fuel are the same as those used for the current fuel. The PAD fuel performance code, Reference 11, was used for the calculations. 4-3 5404LG-87102? r _ _ _ _ . _ _ _ _
J' l TABLE 4.1 TROJAN THERMAL AND HYDRAULIC. DESIGN PARAMETERS Current Fuel Upgraded Fuel Thermal and Hydraulic'Desion' Parameters (ITDP Methods) (ITOP Methods)' Reactor core heat ouput,LMWt 3411 3411 Reactor. core heat ou,tput, Stu/hr 11,641.7x10 0 11,641.7x10 6 Heat generat'ed in fuel, % 97.4 97.4 Core pressure, nominal, psia 2280 2280 Core pressure, minimum steady-state, 2250- 2250 psia Radial power distribution .1.49[1+0.3(1-P)] 1.56(l'+0.3(1-P ) ]
- Minimum DNBR at nominal conditions:
Typical flow channel 2.67 2.48 Thimble (cold wall)' flow channel 2.42 2.33 Minimum DNBR for design transients: Typical flow channel gl.73[a] 1.62(b) I Thimble flow channel 31 71 "3 31 59[b] WRB-1 WRB-1 DNB correlation 4-4 ce-enon [( - - f
i 4 hi t ! TABLE 4.1 (Cont.). TROJAN THERMAL AND HYDRAULIC DESIGN PARAMETERS l- [, Current Fuel ' Upgraded Fuel' Thermal'and Hydraulic Design Parameters (ITDP Methods) -(ITDP Methods) Coolant Flow 6 6 Total thermal flow rate,1b/hr 139.0x10 139.0x10 6 6 Effective flow rate for heat 135.2x10 133.1x10 transfer, lb/hr Core bypass flow, percent 2.7 4.28 Effective flow area for heat 51.1 51.1 2 transfer, ft Average velocity along fuel rods, 16.6 16.3 ft/sec 2 6 6 I Core average mass velocity, Ib/hr-ft 2.65x10 2.61x10 Coolant Temperature Nominal inlet, 'F 554.0 554.0 ( Average rise in vessel, 'F 61.5 61 5 l
)
Average rise in' core, *F 63.1 64.2 Average in' core, 'F 585.5 586.1
~ Average in vessel, 'F 584.7 584.7 i
s 4-5 5488LP671022
'k TABLE 4.1'(Cont.)
TROJAN THERNAL AND HYDRAULIC DESIGN PARAMETERS-Current Fuel Upgraded Fuel Thermal and' Hydraulic Desian Parameters (ITDP Wethods) -(ITDP Methods) Heat Transfer- , 2 '59,700 59,700 A:tive heat' transfer surface area, ft 2 189,800 189,800
' Average heat' flux, Stu/hr-ft Maximum heat flux, for normal operation, 440,340(c) 474,500(d)
Stu/hr-ft2 Average thermal output,' Kw/ft' 5.44 5.44' , Maximum thermal output,,for normal 12.6(c) 13.6(d) operation, Kw/ft Fuel Centerline Temperature Peak at peak linear power for prevention 4700 4700 of centerline melt, 'F p.
}
i i o,
- 4-6 )
me..mn / L-________--__.__
31 4
> f
# 4 TABLE 4.1 (Cont.)
TROJAN THERMAL AND HYDRAULIC DESIGN PARAMETERS
- . Upgraded' Current Fuel Fuel Product 1 (ITDP Wethods).
- Thermal and Hydraulic Desion Parameters (ITDP Wethods)' ,
Pressure Drop
'Across core, psi 24.4+2.4I '3 24.4+2.4I '3 Across vessel, including noizle, psi -4.2.6+6.2 42.6+6.2
[a) This value includes margin. above the ' design DNBR values. A' portion of the available' margin accounts for rod bow penalties and increased. core bypass flow resulting_from the reactor internals bypass ^ flow reversal-modification. [b] This value includes margin above the design DNBR values. A portion of the available margin accounts for rod how penalties. [c] This limit is associated with the value of F0 = 2.32 at 100 percent Dower.. [d] This limit is associated with the value of Fg = 2.50 at 100 percent power. ,
-[e] Based on best estimate reactor flow rate as discussed in FSAR Section 5.1.
l l 4-7 m.-.n m _ _ _ _ _ _ _ _ _ d
4
'" TABLE 4.2 DNBR MARGIN
SUMMARY
Upgraded Current Fuel Fuel 1.17 1.17 Correlation Limit Design' Limit Typical Cell 1.38 1.38 Thimble Cell 1.36 1.36 Safety Limit Typical Cell 1.73 1.62 4 Thimble Cell 1.71 1.59 DNBR Margin (Between Design and Safety Limit DNBR's) Typical Cell 20.2% 14.8%- Thimble Cell 20.4% 14.4%
'DNBR Penalties:
Rod Bow 10.6% <1.5% Upflow* 2.5% n/a
- The core bypass flow increased from 2.7% to 4.28% when the reactor internals were converted from downflow to upflow.
1 4-8 SWOLG-671022
._1___ __
yr- --,-_,.__ i
.,~i'
[.
- - c..
l I h 4 5.0 . ACCIDENT ANALYSIS 4 i
h! 5.0 ACCIDENT ANALYSIS 5.0 Introduction This section summarizes the effects of the increases in the F 3H and FO limits for Trojan on the FSAR Chapter 15 non-LOCA and LOCA analyses. 5.1 'Non-LOCA Accidents- , 5.1.1- F 3H Increase An increase in the full power F 3 H limit does not directly affect the-system transient response of the non-LOCA events presented in the Trojan FSAR. Rather.the normal' operation F H limit is used in the determination 3 of the DNBR for those events for which DNB is the safety acceptance ' criterion. The F M3 is not relevant for the non-DNB related non-LOCA events. As noted in section 4.1, revised core limits were generated reflecting the increased F3 H limit. In addition, revised axial offset limits were also generated, as noted in section 4.5. Based upon these new protection limits, the Overtemperature and Overpower Delta-T (OTDT/0PDT) setpoint equation constants have been recalculated, consistent witn the methods outlined in Reference 12. However, the setpoints previously used in the Trojan safety analyses are conservative with respect to the recalculated setpoints. Therefore, the system transient response for events that rely on OTDT/0PDT for protection is not affected. l' ossov;io/1ozos7 5-1 l-
q, .- _. _- - - - - - - _ - _ - _ , _ . - . _ _ - - . . . . _ , . . - _ - . _ . . _ _ - . - - - _ - - . -
. 1
- i. ,
,i l-An< increase.in F H results,in a decrease in the DNBR-value for a given set 3 1 of; thermal-hydraulic conditions. Therefore, the' transient.value of DNBR for ithe affected'DNB events will be lower than previously'shown in'the FSAR.
However, as noted in~section 4.2, the DN8R reduction'is accommodated'by a corresponding' decrease'of the~ safety analysis DNBR llmit value for those DNB- : ; l
' events analyzed using the Improved Thermal' Design Procedure (ITDP).
1 The following section specifically addresses the impact of an F 3H. limit, increase'on the non-LOCA transients' presented in the Trojan FSAR. l Feedwater System Malfunctions that Result.in a Decrease in Feedwater-
. Temperatur,e~(FSAR e Section 15.1.1)
This ANS Condition II event is bounded by "Feedwater System Malfunctions that- 1 Result in an Increase in Feedwater Flow" (15.1'.2); The safety analysis DNB design basis is met and the conclusions of the FSAR remain valid. -l Feedwater System Malfunctions that Result in an Increase in Feedwater Flow-(FSAR Section 15.1.2) For,this ANS Condition II event, cases are analyzed for both full power and zero power conditions. The zero power case, as discussed,in the FSAR, is bounded by " Uncontrolled RCCA Bank Withdrawal from a Suberitical or Lc.w Power- , Startup Condition" (15.4.1). For the full powe'r case, the transient is effectively terminated by a turbine trip and feedwater isolation on high-high steam generator level. The revised core limits incorporating the increased F H are not exceeded at the most limiting point .in the transient. l 3 Therefore, the safety analysis DNBR limit is met and the conclusions of the '
-FSAR remain valid.
Eavioment Malfunction of Operatino Failure that Results in Increasing Steam Flow (FSAR Section 15.1.3) . For this ANS Condition 11 event, cases are analyzed at beginning and and of life conditions both with and without automatic rod control. In all cases, I the transient approaches an equilibrium condition and a reactor trip does not osasv. tono 20:7 5-2 1
result. The revised core limits incorporating the increased F 3H are not exceeded at the most limiting point in the transient. Therefore, the safety analysis DNBR limit is met and the conclusions of the FSAR remain valid. Inadvertent Ooening of a Steam Generator Relief or Safety Valva (FSAR Section 15.1.4) This ANS Condition II event is bounded by " Spectrum of Steam Piping Failures Inside and Outside Containment" (15.1.5). The safety analysis DNB design basis is met and the conclusions of the FSAR remain valid. Soectrum of Steam Pioino Failures Inside and Outside Containment l (FSAR Section 15.1.5) For this ANS Condition IV event, the ANS Condition II criterion of maeting the DNB limit is applied. The analyses are performed at zero power conditions and assume peaking factors consistaat with the most reactive RCCA stuck out of the core. An increase of the full power F3 H limit results in an increase of the stuck RCCA peaking factor. The effects of the increased stuch rod F 3H l value have been evaluated. The safety analysis DNBR limit is met and the conclusions of the FSAR remain valid. Loss of External Electrical Lead (FSAR Section 15.2.2) For this ANS Condition II event, cases are analyzed at beginning and end of life conditions both with and without pressurizer control. For the cases with
;.ressurizar control, the transient is terminated by an Overtemperature Delta-T reactor trip. The OTDT setpoint assumed in the current analysis presented in the FSAR is conservative with respect to the new OTDT setpoint. For the cases without pressurizer control, the transient is terminated by a high pressurizer pressure reactor trip. The revised core limits incorporating the increased F H are not exceeded at the most limiting point in the transient.
3 Therefore, the safety analysis DNBR limit is met and the conclusions of the FSAR remain valid. 0936v:1o/102087 5-3
Turbine Trio (FSAR Section 15.2.3) This ANS Condition II event is bounded by " Loss of External Electrical Load" (15.2.2). The safety analysis DNB design basi:: is met and the conclusions of the FSAR remain valid. Inadvertent Closure of MSIVs (FSAR Section 15.2.4) This ANS Condition II event is bounded by " Loss of External Electrical Lead" (15.2.2). The safety analysis'DNB design basis is met and the conclusions of the FSAR remain valia.
~
Loss of Non-emergency A-C Power to the Station Auxiliaries (FSAR Section 15.2.6) This ANS Condition II event is analyzed to show that the adequate heat removal capability exists to remove core decay heat and stored energy following l reactor trip. This criterion is not affected by the F3 H limit. With respect to the DNB criterion, this event is bounded by " Complete Loss of ForcedReactorCoolantFlow"(15.3.2). The safety analysis DNB design basis is met' and the conclusions of the FSAR remain valid. Loss of Normal Feedwater (FSAR Section 15.2.7) This ANS Condition II event is analyzed to show that adequate heat removal capability exists to remove core decay heat ar.4 stored energy following reactor trip. This criterion is not affected ry the F H3 limit. With respect to the DNB criterion, this event is bonded by " Loss of External Electrical Load" (15.2.2). The safety analysis DNB design basis is met and the conclusions of the FSAR remain valid. l l L l l l 0936v:10/102087 5-4 l 1 _ _ _ _ _ _ - i
L
'Feedwater System Pipe Breaks Inside and Outside Containment (FSAR Section 15.2.8). .This is'a Condition IV event and the criterion of DNB is not' applicable.
Therefore,- the increased F3 rl. limit does not impact the analyses and the conclusions of the FSAR remain valid. Partial loss of Forced Reactor Coolant Flow (FSAR Section 15.3.1) For this ANS Condition II event, the transient is terminated by a low RCS loop flow reactor trip.- The effect of the increase in the F 3H limit on the-transient DNBR is accommodated by the reduction in the DNBR limit. The safety analysis DNBR limit is met and the' conclusions of the FSAR remain valid. Complete Loss of Forced Reactor Coolant Flow (FSAR Section 15.3.2) For this ANS Condition II event, the transient is terminated by an undervoltage or underfrequency reactor trip. The affect of the increase in the F3H limit on the transient DNBR is accommodated by the reduction in the DNBR limit. Therefore, the safety analysis DNBR limit is met and the conclusions of-the FSAR remain valid. Reactor Coolant Pump Shaft Seizure (FSAR Section 15.3.3) This is a Condition IV event and the criterion of DNB is net applicable. Therefore, the increased F H limit does not impact the analyses and the 3 conclusions of the FSAR remain valid. Reactor Coolant Pump Shaft Break (FSAR Section 15.3.4) This ANS Condition IV event is bounded by " Reactor Coolant Pump Shaft Seizure" (15.3.3). Therefore, the conclusions of the FSAR remain valid. 0936v:1D/1020s7 5-5 t _. . _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _
i:
' Uncontrolled RCCA Bank-Withdrawal from a Suberitical or Low ~ Power'Startup
[ Condition (FSAR Section.15.4.1) ,
~
For this ANS Condition II event,Ithe analysis is~ performed at zero power conditions. The current safety analysis. presented in the FSAR did not use the F H in determining that the safety analysis DNB design basis is met. 3 Therefore, the conclusions of the FSAR remain valid. Uncontrolled RCCA Bank Withdrawal at Power (FSAR'Section 15.4.2) ' For'this-ANS Condition II event, v.arious power levels and: reactivity insertion rates for both minimum and maximum reactivity feedback are analyzed. The transients.are terminated by an Overtemperature. Delta-T or High Neutron Flux
' reactor trip. For those cases that trip on OTDT, the current analysis presented in the FSAR'is conservative with respect to the new OTDT setpoint.
The revised core limits incorporating the increased F H3are not~ exceeded et the most-limiting point in the transient. Therefore, the safety analysis-DNBR limit is met and the conclusions of the FSAR remain valid. l RCCA Misalignment (FSAR Section 15.4.3) j For the events presented in this section of the FSAR, the effect of the increase in the F M limit is accommodated by the reduction in the DNBR 3 limit. Therefore, the safety analysis DNBR limit is met and the conclusions of the FSAR remain valid. Startuo of an Inactive Reactor Coolant Pump (RCP) with Low Hot Leo Temperature (FSAR Section 15.4.4) For this ANS Condition II event, the increase in the F 3 H limit is ! accommodated by the reduction in the DNBR limit. Therefore, the safety analysis DNBR limit is met and the conclusions of the FSAR remain valid. 0936v:1o/102087 5-6
n
'CVCS Malfunction that Results in a Decrease in the Boron Concentration in the 3 Reactor Coolant (FSAR Section 15.4.5) i This ANS Condition II-event is analyzed to show that adequate time exists for operator action to terminate a dilution event prior to a loss of shutdown margin. With respect to the DNB criterion, the at power cases, Modes 1 and 2, are bounded by " Uncontrolled RCCA Bank Withdrawal at Power" (15.4.2). For the post reactor trip transient of the Mode 1 and 2 analysis and for events occurring in Modes 3 through 6, the minimum operator action time' ensures that a return to power does not occur. Hence, for the potential return to power the full power F3 H' limit is not applicable. Therefore, the safety- l L
analysis DNB design basis is met for the full power event and the conclusions E of the FSAR remain valid. Spectrum of RCCA E3ection Accidents (FSAR Section 15.4.8) This is a Condition IV event and the criterion of DNB is not applicable. Therefore, the increased F3 H limit does not impact the analyses and the conclusions of the FSAR remain valid. Inadvertent Operation of Emeroency Core Coolina System (ECCS) Durino Power Operation (FSAR Section 15.5.1) For this ANS Condition II event, the transient is initiated by a spurious safety injection signal. The injection of borated water drives nuclear power and RCS temperature down. Hence, the most limiting thermal- hydraulic conditions occur at the initiation of the transient. Therefore, the safety analysis DNBR limit is met and the conclusions of the FSAR remain valid. Accidental Depressurization of the 'leactor Coolant System (FSAR Section 15.6.1) i For this ANS Condition II event, the transient is terminated by an ' Overtemperature Delta-T reactor trip. The revised core limits incorporating q the increased F M are not exceeded at the most limiting point in the l 3 transient. Therefore, the safety analysis DNBR limit is met and the conclusions of the FSAR remain valid.
)
l I 0936v:1o/1o2087 5-7 1
1
- i. z 5.1.2 F gIncrease l
To ensure that cladding integrity and fuel melting at the " hot spot" are-maintained within the applicable safety _' analysis limits, the two impacted I-l- E transients for an. increase in the F glimit were examined.- The Reactor. Coolant Pump Shaft Seizure event presented in FSAR section 15.3.3 assumed an-Fg of.3.0. Therefore, the. conclusions for this event remain bounding for-the increase in F g.. However, the RCCA Ejection Accidents presented in" section 15.4.8 do not bound the increased Fg limit associated with the fuel upgrade. Therefore, the RCCA Ejection accidents were reanalyzed with an assumed initial F value of 3.0. In addition, an increased nominal core-g power level of 3558 Mwt ,was' assumed.to bound Trojan ~for the potential uprating. ' Based upon the reanalysis, all acceptance criteria are met and the 4 conclusions of the FSAR remain valid. On the following page a summary of.the' important input parameters and the results are given in Table 5.1 The non-LOCA FSAR markups, including tables and figures, are also presented. T O osaav:1o/1020s7 5-8
L t. ' Table 5.1 ' Parameters Used in the Analysis of the Red Cluster
. Control Assemoly Ejection Accioent HFP HIP HFP HZP Time In Life BOL BOL EOL EOL Power Level (%) 102 0 -102 0 Initial Tavg ('F) 591.2 557 591.2 557 Ejected Rod Worth 0.18 0.72 0.21 0.90
(% Delta-K/K) DelayedNeutronFraction(%) 0.52 0.52- 0.44 '0.44 Feedback Reactivity Weighting 1.30 2.07 1.60 3.55 Trip Reactivity (% Delta-K/K) 3.2 1.0 3.2 1.2 FgBefore Rod Ejection 3.00 - 3.00 - F After Rod Ejection 6.90 12.0 7.70 21.0 0 Number of Operational Pumps' 4 2 4 2 Max. Fuel Pellet Average Temp 4156 3386 4067 3386 (*F) Max. Fuel Center Temp ('F) 4981 3922- 4908 3820 Peak Clad Temp (*F) 2350 2495 2300 2577 Max. Fuel Stored Energy (cal /gs) 183 143 178 143 Percent Fuel Melt (%) <10 0 <10 0
')
1 1 0936v:10/102087 5-9
_-.qw , yy m.
-,__7,.- . . . . =
1 i 1 fil' .p 3 NON-LOCA' , , FSAR CHANGES m
\
f. s 1 P
}
L-
- c. _. _ - - _ _ _ _ _ -a
i to assure that specified accepta& fuel design limits are not exceeded - during' any condition of normal' operation,' including the effects of
. anticipated operational occurrences. g
, g .{g -
b ~ t va h<c .
. serfy s escontrolandProtection_
The' reactor core with its related coolant, System
.is designed to function throughout its design lifetime without exceeding-Jacceptable fuel damage limits. The reactor trip system is designed'to actuate a reactor trip for any anticipated combination of Plant condi-f tions'when necessary to ensure that the departure from nucleate boiling
~
ratio (DNBR) is not /=&tSTIknd fuel center temperatures are below the
-f melting point ' of U02. . The' core design, together with reliable process and decay heat removal systems, provides for this capability under all expected conditions of normal operation with appropriate margins for.
uncertainties and anticipated transient situations, including ths effects' j of the loss of reactor coolant ' flow,' trip of the turbine generator, loss of' normal feedwater and loss of both normal- and preferred power sources.
' Chapter 4 discusses the design bases and design . evaluation of reactor components. . Chapter 5 discusses the RCs. The details of the reactor
. trip and EST systems design and logic are discussed in Chapter 7. This information supports the accident analyses' presented in Chapter 15.
Criterion 10 is met by Trojan design. Criterion 11 - Rasetor Inherent Protection. The reactor core and associ-aced coolant systems shall be designed so that in the power operating range the net effect of the prompt inherent nuclear feedback character-istics tends to compensate for a rapid increase in reactivity. A negative reactivity coefficient is a basic feature of core nuclear design as discussed in Chapter 4. x Criterion 11 is met by Trojan design.
- l. Criterion 12 - Suppression of Resetor Power Oscillations. The reactor L- core and associated coolant, control and Protection Systems shall be l designed to assure that power oscillations which can result in conditions i- Amendment 3 3,1 10
. (July 1985) gj L _
y
.The Protection' System at Trojan consists of a reactor trip-system and an; ci ESF system. -These systems are described in Sections 7.2 and.7.3.
Operational limits for the reactor trip ' system are defined by analyses of all Plant operating and fault' conditions requiring rapid rod insertion
.to prevent or limit core damage. The' design bases for all anticipated transientsLor' faults are:-
,(1) Minimus DNBR shall noe be lhe &.
daTe . derl
/#. s ,s Ma e fb y sq /s rw o f " Vo/a e .
(2) Clad strain on the fuel element shall not exceed. 1 percent. (3) No ' center melt shall occur in the. fuel elements. A region of permissible core operation is defined ini erms t of power, axial power distribution and coolant flov 'and temperature. The reactor; ; trip system monitors these process variables. If the region limits are approached during operation, protection circuitry will automatically. j actuate' alarms, initiate load runback, pravent control rod withdrawal and/or. trip the. reactor. Operation within the permissible region and complete core protection is assured- by the overtemperature AT and over-power AT reactor. trips over the pressure range defined by the pressurizer high pressure and pressurizer low pressure reactor trips, provided that the transient .is slow with respect to piping delays from the' core to the' temperature sensors. High nuclear flux and low coolant flow reactor-trips provide core protection against transients which are faster than the AT response. Also, thermal transients are anticipated and evoided by reactor trips actuated by turbine trip and primary coolant pump circuit .j breaker position. !
]
The safety function of the ESF system is that of processing signals used I
-for EST actuation and generation of the actuation demand. The conditions leading to ESF actuation are:
q (1) Low pressurizer pressure. h! 3.1-18 j dl-2 1 w_ -
Criterion 24 is met by Trojan design. Criterion 25 - Protection System Requirements for Reactivity Control Malfunctions. The Protection. System shall be designed to assure that specified acceptable fuel design limits are not exceeded for any single malfunction of the Reactivity Control Systems, such as accidental withdrawal (not ejection or dropout) of control rods. Reactor shutdown by full-length rod insertion is completely independent of the normal control function since the trip breakers interrupt power to the rod s chanisms regardless of existing control signals. The Protec-tion System is designed to limit reactivity transients so that DNBR will exceed M[-- for any
% sasingle htymalfunction on ly sisin either
/n .treactor a/u e. control system.
The analysis presented in Chapter 15 shows that for postulated dilution during refueling, startup or manual or automatic operation at power, the operator has maple time to determine the cause of dilution, terminate the source of dilution and initiate reboration before the shutdown margin is lost. The rod control system is discussed in Sections 4.2 and 7.7; the boric acid injection system is discussed in Section 9.3.4. Analyses of the effects of possible malfunctions are discussed in Chapter 15. The analyses show the acceptable fuel damage limits are not exceeded even in the event of a single malfunction of either system. Criterion 25 is met by Trojan design. Criterion 26 - Reactivity Control System Redundancy and Capability. Two independent Reactivity Control Systems of different design principles shall be provided. One of the systems shall use control rods, preferably including a positive means for inserting the rods, and shall be capable of reliably controlling reactivity changes to assure that under condi-tions of norinal operation including anticipated operational occurrences, and with appropriate margin for malfunctions such as stuck rods, speci-fled acceptable fuel design limits are not exceeded. The second Reacti-ity Control System shall be capable of reliably controlling the rate of reactivity changes resulting from planned, normal power changes (including xenon burnout) to assure acceptable fuel d. sign limits are not exceeded. One of the systems shall be capable of holding the reactor core suberitical under cold conditions. 1-Amendment 6 3.1-24 (July 1987) }
l I i j u
, 4 1
^
es ' '
' ' ' ' ' l '
l l I l l l l _
~
OF Comorftens LOCUS OF Comemons wuent Powse e isom op .=
= e*ERE Ossea e1.F8PGA PosTmEAesAL OSSaGesPLew ,
taasAL passen PLOW l Su se07 CMessieEL j 80 - - Phc70As : PEIA
.l m -
75 - - asse pela - N89 7 ^- ,
=. =
g
- \ :
,g ;.; ;_; . AT ..
, 9 .
- TRIP blIIE
. M" - .
g = S {*
~ f.* - .. i im e.
WW M E _ _
! e _ _
}
- er pourts wasme -
STEAas egestAAton SAFETT ~ 1A&&VS$ OPEN m -
; so_ _ -
t - t ._ _ _ t . - . erTttspamAfung AT fnsp UssES - a N"" =~ _ __; Isaassum insimumspr7 Emm0Att At 30sCATSO PRESSWAES
?- m =
j - - i(-
\t i
=
t m - 1 i f I f f f i I f f f 1 I 1 520 Ho 360 500 400 620 eso eeo eso 700 fave (*M fgQ & C,L td / ff/ [02 40cO'N8 SNE
! Figure 15.0-1 Illustration of overtemperature 6 and Overpower AT Protection Amendment 2
__ Ju y 984
.- ?
i ' l ' i ' l l ' i ' l ' Loew r amorfree as0er #WugR = s.sar
. 9 enaneZNAL 999 an!,NZasne LOCm W CwCZT!OGC DOE Dem af LJNtf Put astNtanat
=_
asAAAWD 40
- asammen non orssee nor owest racrats - 3 o
,0 -. .
\ -
\ -
- asso para .,, !
- -l 73- .
. amo es 4 -
- 3900 **E' i
~ -
s m essa 70_ . -
. - 1
- o.a.o .r -
into LZw ., 55-. . - l 1
- - i
- R y ,0 -
~ ~ "
E g
~
1
\}$ -
)
{ j
\
E ss - -
.d c
- Loew a morwis awsur "
straat asunarom s4rrry 60 - - *****N - i
- j 45 - -
q 1 40 ,,,, _ oventrassemafuer er Tara truts _ .. CENcL1,cINs anaxt**t IN5fituteENT 5put0R81 -- Ar recicarum persswers a )
-BS 4
y, .- e e i i e s i i i i sgo 540 SE0 500 000 E20 040 660 #000' ) Tavs (deg-F) Figure 15.0-1 Illustration of overtemperature and overpower AT Protection ,,p 1 C_--_-__---------- I
.i 15.1.1 FIEDWATER SYSTEM MALFUNCTIONS THAT RESULT IN A DECREASE'IN FEEDWATER TEMPERATURE ,
15.1.1.1 Identification of Causes in the Accident Description Baduction in feedvater temperature will cause an increase in core power l by decreasing reactor coolant temperature.. Such transients are atteau-ated by the thermal capacity of the' secondary and of the RCS. The
# overpower-overtemperature protection (neutron overpower, overtemperature - 1 and overpower AT trips) prevents any power increase which could lead to
... ,.... u>,>-- - > r ~~ >, ,, 9
,s._....... a... '
pea a R ta3s %% N .pa/s+y a =*ty = 8* =e.+ 'Ma'- *
- One example of a decrease in feedwater temperature is the accidental opening of a feedwater bypass' valve which diverts flow around a portion l of the feedwater heaters. Without the proper amount of preheating, feed-inster temperature entering the steam generator will be reduced suddenly.
With the Flant at no-load or at power conditions, the addition of cold l l feedwater any cause a decrease in RCS temperature and thus a reactivity J
)
insertion due to the effects.of the negative moderator coefficient of reactivity. This fault is considered a Condition II event. i 15.1.1.2 Analysis of Effects 4 When the Plant vss first licensed, evaluation of an accident resulting in decreasing feedwater temperature was not analyzed per se, nor was .)! l an analysis required to be done. Bowever, one of the accidents discussed in Section 15.1.2, -excessive feedwater flow at zero load, assumes a feedwater inlet temperature of 70*F with a step increase in feedvater flow to one steam generator to 100 percent of nominal. There is no credible circumstance under which I a Feedwater System malfunction could reduce feedwater temperature below 70*F. A complete description of assumptions and methods employed is in Section 15.1.2. 15.1-2 gg
):
K' .,
.A decrease in normal feedwater temperature is classified'as an ANS Condi-tion II event fault of moderate frequency as defined in Section 15.0.1'.2.
~
15.1.1.3 Conclusion r As specified in Section 15.1.2, the DNBRs encountered for ucessive feedwater temperature decrease (and excessive feedwater addition) at sero power are well above the limit % value,o h
.t g 4 , ,
15.1.2- FEEDWATER SYSTEM MALFUNCTIONS THAT RESULT-IN AN INCREASE IN FEEDWATER FLOW 15.1.2.1 Identification of Causes and Accident Description Feedwater system malfunctions'that result in excessive feedwater' flow are means of increasing core power above full power. Such transients are attenuated by the thermal capacity of the secondary plant and of the RCS. The overpower-overtemperature protection (neutron overpower, overtemperature and overpower AT trips) prevents any power increase , E ** M* which could lead to a DNBR #sMe la w ne Oe. O+7 ***Ir S i One 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 error. At power, this' excess flow causes a greater load demand on the RCS due to increased subcooling in the steam generator. With the Plant at no-load conditions the addition of cold feedvater may cause a decrease in RCS temperature and thus a reactivity insertion due to the ef fects of the negative moderator coefficient of reactivity. Continuous addition of excessive feedwater is prevented by the steam
. generator high-high level trip.
This accident is classified as an ANS Condition II event, as defined in Section 15.0.1.2. i b - 15.1-3 Y!"
The full power case.(EOL, with control) gives the largest reactivity feedback and results in the greatest power increase. A turbine trip and reactor trip is actuated when the steam generator level reaches the high-high level setpoint. -The time sequence of events for this accident
. is shown in Table 15.1-1. .
For all excessive feedwater cases continuous addition of cold feedwater-is prevented by closure of all feedwater control valves and a trip of the feedwater pumps on steam generator high-high level. Feedwater pump trip causes closure of the feedwater pump discharge valves. Transient results'(see Figure 15.1-1) show the increases in nuclear power at AT, associated with the increased thermal load on the reactor. Steam generator level rises until the feedwater is terminated as a result of the high-high steam generator level trip. The DNBR does not drop below M99? tk4. 3./a.+7 % > l m .t % h.4.. 15.1.2.3 Conclusion The reactivity insertion rate which occurs at no load following e::ces-sive feedwater addition is less than the maximum value considered in the analysis of the rod withdrawal from a subcritical condition. Also, the ' DNBRs encountered for excessive feedwater addition at power are well above the linie6eg value. ' ' 34A+y 4a.ly $s3 , ( i 15 1.3 EQU1PMENT HALTUNCTION OR OPERATING FAILURE THAT RESULTS IN INCREASING STEAM TLOW f 15.1.3.1 Identification of Causes and Accident Description An increasing steam flow incident is defined as a rapid increase in the steam flow that causes a power mismatch between the reactor core power and the steam generator load demand. The reactor control system is i l designed go acconnodate a 10 percent step load increase or a 5-percent / min ] 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 (RPS). . 15.1-6
~7
l (2) Manually controlled reactor at EOL. (3) Reactor in automatic control at BOL. (4) Reactor in automatic control at EOL. 1 ' The analyses are performed using a detailed digital model of the Plant including the core kinetics, RCS and steam and feedwater system behavior. At BOL the core has the least negative moderator temperature coefficient \ of reactivity and therefore the least inherent transient capability. At EOL 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 being taken for pressurizer heaters. Initial core power, reactor coolant average temperature, and reactor coolant pressure are assumed'to be at their nominal values. Uncertain-ties in initial conditions are included in the limit DNBR as described in WCAP-8568 I) . 15.1.3.2 2 Results Figures 15.1-2 through 15.1-5 illustrate the transient with the reactor in the manual control mode. As expected, for the BOL 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 EOL manually controlled case there is a much larger increase in reactor power due to the moderator feedback. A reduction j in DNBR is experienced, but DNBR remains above W 6 e. s. A h a.** tr o s i,..+ v . h. . I Figures 15.1-6 through 15.1-9 illustrate the transient assuming the reactor is in the automatic control mode. Both the BOL and the EOL cases show that core power increases, thereby reducing the rate of decrease in coolant average temperature and pressurizer pressure. For both the BOL and EOL cases, the minimum DNBR remains above ban. O t. 8 "
- l 15.1-8 S
P The tima sequence of events for these accidents is shown in Table 15.1-1. ! 15.1.3.3 Conclusions It has been demonstrated that for such staan flow increasee, the minimum ! DNBR during the transient will not be below the limit N, va,t wa. ; safe +y My4 5 l 15.1.4' INADVERTENT OPENINC OF A STEAM CENERATOR RELIET OR SAFETY VALVE 15.1.4.1 Identification of Causes and Accident Description , The'most severe core conditions resulting from an accidental depressuriza-
~
tion of the Main Steen System (HSS) are associated with an inadvertent opening of a single steam dump, relief or safety valve. The analyses performed assuming a rupture of a main steam pipe are given in Section 15.1.5. The sceau release as a consequence of this accident results in an initial increase in steam flow which decreases during the accident as the steam pressure falls. The energy removal from the ICS causes a reduction of coolant. temperature and pressure. In the presence of a negative moder-ator temperature coef ficient, the cooldown results in a reduction of core shutdown margin. This, accident is considered an ANS Condition II event as defined in Section 15.0.1.2. The analysis is performed to demonstrate that the following criterion ] is satisfied: Assuming a stuck rod cluster control assembly (RCCA) and I a single failure in the Engineered Safety Features (EST), there will be no consequential damage to the core or RCS after reactor trip for a steam ;Q release equivalent to the spurious opening, with failure to close, of the largest of any single steam dump, relief or safety valve. i ! The following systems provide the necessary protection against an acci-dental depressurization of the MSS: 15.1 9 Amendment 3 (July 1985)
//L-/O
l-t p-15.1.4.4 Subsecuent Analysis In May 1980, to avoid consequential damage of one or more CCys following a secondary system high-energy line rupture. West.b shouse (M) recosamended that the CCp mini-flow isolation valves not be automatically shut upon a SIS . Circuitry changes were unde such that the CCF mini-flow isolation valves no longer close automatically on a sis and plant Emer-
< gency Operating Procedures were modified to instruct the operator to:
(1) Close the CCP sini-flow isolation valves when the actual BCS pressure drops to 1425 psig. 31 (2) Roopen the CCp mini-flow isolation valves should the wide-range BCS pressures subsequently rise to >2000 psig. The effects of maintaining the mini-flow isolation valves in a nomally open position following the accident described in section 15.1.4 were The results of the transient showed the.DNBE limit {i investigate (. criterion continues to be met, w % % @ A .. . 15.1.5 SptCTRUM OF STEAM SY3TDi PIPING FCI,URES INSIDE AND OUTSIDE couTATEMENT i 15.1.5.1 tuoture of Main steen t.ine 15.1.5.1.1 Identification of Causes and Accident Description The steam release arising from a rupture of a main steam pipe would result in an initial increase in steam flow which decreases during the accident as the steen pressure falls. The energy removal from the RCS -I I j causes a reduction of4 coolant temperature and pressure. In the presence ; of a negative moderator temperature coefficient, the cooldown results in a redaction of core shutdown margin. If the most reactive BCCA is assumed stuck in its fully withdrawn position after reactor trip, there l is an increased possibility that the core will become critical and return to power. A return to power following a steam pipe rupture is a l Amendment 4 15.1-14 (July 1986) ML-M !
1,; e Steam flow is measured by monitoring dynamic head in nozzles inside the steam pipes. The nozzles which are of considerably. smaller diameter than the main steam pipe are located inside the Conttisment near the
- steam generators and also serve to limit the maximum steam flow for any.
break further downstrema. 15.1.5.1.2 Analysis of Effects and Consequences Method of Analysis: l 9 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 stem line
. break. The 1,0FTEAN code has been used. 'h (2) The thermal and hydraulic behavior of the core following a steam line break. A detailed thermal and hydraulic .,
digital computer code, THINC, has been used to determine .f if DNB occurs for the core conditions computed in (1) i < above. (3) cause the steam break accident is a relatively a w ev with respect to the flow transient time, sady-state lear and thermal hydraulic analyse are used to l determine minimus DNAR during the tr sient. The TE:3tc system t sient code is used generate the .} 's' core power, inlet parature, pr sure and flow. Typi-cally, the minimum DN occur en the power-to-flow ratio is the largest. Se al state points are chosen at and around where t nazi power-to-flow ratio ! occurs and a prel nary nuclear thermal hydraulic j evaluation is de. Following this e luation, the
. worst sta point with respect to DNBR is hen analysed in de .1. If the minimus DNBR is well above , then it is concluded that the other state points will ea 15.1-17 Amendtnant 3 (July 1985)
/
A1-/A l 1
y i 7-i. s. se case is 1.30 j DNBR . If the DNBR for th
. additional state po analyzed until.it can.be
- ensured that nimum DNBR vi .30 for all .
j t state nes. j l ! The following conditions were assumed to exist at the time of a main steen line break accidents ( y (1) ECL shutdown mergin at no lead, equilibrium sonou condi- 1
'1 tions, and the'most reactive assembly stuck in its fully j uithdrawn ' position: Operation of the control rod banks l d
'during core burnup is restricted in such a way that addi-tion of positive reactivity in a steam line break accident f i
will not lead to a more adverse condition than the case l
)'
analysed' (2) The negative moderator coefficient corresponding to the ) ECL redded core with the most reactive rod in the fully withdrawn position: The variation of the coefficient with tamparature and pressure has been included. The k,gg vs temperature at 1000 poi corresponding to the negative moderator temperature coefficient used is shown in Figure 15.1-10. The effect of power generation in the core on overall reactivity is shown in Figure 15.1-13. The core properties associated with the sector nearest the affected steam generator and those associated with . the remaining sector were conservatively combined to obtain average core properties for reactivity feedback calculations. Further, it was conservatively assumed that the core power distribution was uniform. These two conditions cause underprediction of the reactivity feedback in the high power region near the stuck rod. To verify the conservacias of this method, the reactiv-e ity as well as the power distribution was checked for the state points shown in Table 15.1-2. These core 1 l 15.1-18 g4-/3
V t-I le should be noted that following a steam line break only one steam generator blows down completely. Thus, the remaining steam generators
. are still available for dissipation of decay heat after the initial
.f transient is over. In the case of loss of offsite power this heat is removed to the atmosphere via the stems line safety valves which have been sized to cover this condition.
. .J The seguence of events is shown in Table 15.1-3. -
Margin to Critical Heat Fluz: A DNS.anklysis was performed for the three cases most critical.co DNB.
- - Five points from each case were asamined. It was found that all cases
' had a binimum DNB1 >P"StPT c444 ar % + %. safe 47 = \yses Im.e AW e.. , !
Conclusions:
- The radiological consequences of this accident era presented in section 15.1.5.1.5.
In Section 6.2.1.1.1, the Containment pressure response for the worst , I steam line break results in a peak pressure smaller than that esiculated j for the DBA, which is less than the containment design pressure. 15.1.5.1.4 subsequent Analysis , As previously explained in Section 15.1.4.3, CCP sini-flow isolation valves are no longer being shut automatically on a sIh. 3 e effect of maintaining these mini-flow isolation valves open on a rupture of a main steam line, as discussed in Section 15.1.5.1, was investigated. Analy-sis results showed that, even with reduced safety injection flow into the core, no DNB occurred for any rupture. Amendment 3 15.1-24 (July 1985) kl. '/Y
l i These analyses are 1.11ustrative of a pipe break equivalent'in size to a single valve opening. 15.1.5.3.3 Conclusions The analysis presented in section 15.1.5 demonstrate that the conse-quences of a minor secondary system pipe break are acceptable since o=
-,JG- '
____ for a more critical mejor secondary system pipe g mmre D^l 64 9 g%Yee tan 4k4. se/ 4y My ses ti .f Va.Id l No ndiological effects have been evaluated for this category of accident since other accidents having similar characteristics and comparable or i worse consequences are evaluated (see section 15.1.5.1.5 and 15.2.6.4). l
-)
1 3 i
. J 1
( l Amendmenc 5 15.1-32 (September 1986)
- l. .. _ _ _ _ . _ _ _ _ _ _ _ _ _ _ _
- The' integrity of the core is maintained by operation of the RPS, is, the DNBR will be maintained above the value. Thus, no core safety limit.
will be violated, g yg 3, . f 15.2 2.5 Radiological Evaluation The' radiological consequences are similar to section 15.2,6.4. Radio-activity release through the steam generator relief: valves consists of the concentration present in the secondary due to assumed primary-to- -l I secondary leakage.
.I 15.2.3 TURBINE TRIP ($TCP VAI,VE CLOSURE) 15.2.3.1 Identifiestion of Causes and Accident Description Major load loss can result from a turbine trip event. Offsite power is available for the continued operation of Flant components such as the RCPs. The case of loss of non-emergency a-c power is analysed in @
8ection 15.2.6. For a turbine trip, the reactor would be tripped directly (unless below approximately 10 percent power) from a signal- derived from the electro-hydraulic control trip fluid pressure and turbine stop valves. The automatic steam dump system would accommodate the excess steam genera-tion. Reactor coolant temperatures and pressure do not significantly increase if the steam dump system and pressurizer pressure control system are functioning properly. If the turbine condenser was not available, the excess see.as generation would be dumped to the atmosphere. Additionally, main feedvater flow would be lost if the turbine condenser was not available. For this situation, feedwater flow would be main-
~
tained by the Auxiliary Feedwater System (AFWS). A turbine trip event is classified as an ANS Condition II event (see Section 15.0). Turbine trip initiation signals include: (1) Overspeed trip (sechauical). 15.2-7 Amendment 1 (July 1983) hYb
15.2.3.2 Analysis of Effects and Consequences The method of analysis, including limiting conditions and assumptions, is identical to that for a loss of external electrical load in Section 15.2.2.2. The time' sequence of events is similar to that of loss of arternal electrical load as shown in Table 15 2-1. 15.2.3.3 conclusions l Results of the analyses, including those in Reference 1, show that the Plant design is such that e. consequences of a turbine trip with or without , a direct or immediate reactor trip presents no hazard to the integrity oJ the ICS or the MSS. Frassure-relieving devices incorporated in the two systems are adequate to limit maximum pressures within the design limit.. The integrity of the core is maintained by the operation of the RFS, is, the DNBR will be maintained above the value. Thus, no core safety limit will be violatad. 3Ay am. lyn lia e i" ; 15.2.3.4 Radiological Evaluation The radiological consequences are similar to those analyzed in Section 15.2.6.4. E, radioactivity release through the steam generator relief valves consists of the concentration"present in the secondary due to assumed primary-to-secondary leakage. 15.2.4 INADVERTENT CLOSURE OF M3IVs The analysis of this transient is bounded by the analysis presented in Section 15.2.2. 15.2-9 Amendment 1 (July 1983) k} 'l1
The first few seconds of the transient will closely resemble a simula-tion of the complete loss of flow incident (see Section 15.3). ie, core damage due to rapidly increasing core temperatures is prevented by a After the reactor trip, stored and l promptly tripping the reactor. residual heat must be removed to prevent damage to either the RCS or the l core. The natural circulation flow as a function of resid d l reactor power is presented in Table 15.2-2. l 15.2 6.3 Conclusions Resu11is of the " Decrease in Reactor Coolant Systes Flow Este* analysis (Section 15.3) and the " Loss of 18ormal Feedwater" analysis (section 15.2.7) show that for a loss of all offaite a-c power no adverse conditions The RCS occur in the reactor core. The DNBR is maintained above is not overpressurised and no water relief will occur through the h safc4y exa.t r u ) pressurizer relief or safety valves. Thus, there will be no cladding b=.+ ll damage and no release of fission products to the RCS. vta. 15.2.6.4 Radiological Evaluation A detailed analysic of effects and consequences other than radiological consequences is presented in Section 15.2.2. It is concluded that following a loss of offsite s-c power with subsequent turbine and reactor trip, the steam systes pressure rises, resulting in the auto-matic operation of the relief valves and a steam dump to the atmosphere. Radiologi-Steam dump to the condenser is assumed not to be available. cal consequences of this incident are presented below. 15.2.6.4.1 Fission Froduct Release Assumptions The following conservative assumptions were used to evaluate the activ-ity release from a postulated loss of of fsite power to the station auxiliariass (1) . Loss of of f aite power results in no clad damage (section 15.2.2). Amendment 1 15.2-12 (July 1983) d-/7 ,
-- - - - - - _ - -- ]
l a. L . j 15.2.8.2.3 Conclusions The analysis presented.in Section 15 2.8.1 demonstrates that the.conse-quences of a minor secondary systes pipe break are acceptable'since =
= ' ' ' "" '
- for a more critical asjor secondary system pipe break. -- - \x. W A,m em Od6A C g %4v- Ng . % c- '
y
. sofe+y sa.J s,es I. * . 4- 4 s.\ me.
A discussion of the envireamental consequences of this accident is presented in Section 15.5. 15.2.8.3 Other Considerations Due to Feedwa,ter Systes Fipe Bre..3 15.2.8.3.1 Control Systen Effects Due to Adverse Environmental Conditions Resulting From Righ Energy Line Breaks Based on Westinghouse (V) generic assessment, there are 15 potential . interaction scenarios where the effect of adverse environne,ats on control systems could land te consequences eers limiting than the - results presented in previous sections of this report. V has demon-1 strated that these 15 interactions are bounded by the consideration of the following four limiting scenarios as applied to Trojas. , 1 Induced Malfunction in the Steam Generator Pog7 Control Systas: (1) Susmary of postulated generic scenario Follpwing a feedline rupture outside Containment, the steam generator PORVs are assumed to exhibit a conse-quantial failure due to an adverse environment. Failure of the FORY in the open position results in the depres-surisation of multiple steam generators, which are the source of steam supply for the seses turbine-driven euziliary feedwater pump. Eventually, the turbine-driven suziliary feedwater pump will not be capable of delivering auxiliary feedwater to the intact seems generators. A potential exists that no auxiliary Amendment i 15.2-26 (July 1983)
/ $~lf ;
l
-)
15.3.1.2.2 Inittel Conditions initial core power, reactor coolant sverage temperature,-and reactor
' coolant pressure are assumed to be at their nominal values. Uncertain- .j ties in initial conditions are included in the limit DDR as described in NCAP-8568. With all but one loop operating, the maxisua power level-allowed for that mode of operation is assumed. {
A consev~ntively large absolute vslue of the Doppler-only power coeffi-cient is used (see Table 15.0-2). The total integrated Doppler reac-tivity fros 0- to 100 percent power is assuand to be 0.016 Ak. The lowest absolute anguitude of the moderator temperature coefficient . I (0.0Ak/*F) is assumed since this results in the maximum hot-spot heat flux during the initial part of the transient when the miniana DNBR is reached. 15.3.1.2.3 Results The. calculated sequence of events is shown in Table 15.3-1 for the two cases smalyzed. Figures 15.3-1 through 15.3-6 show the loop coastdowns, the core flow coastdowns, the nuclear power coastdowns and the average , and hot channel heat flux coastdowns for each case. The miniana DNBR for each case is MwaDe gna.ha- % +%e. sdefy m hsu 1.ae.1 Ww 15.3.1.3 Conclusions {j 4 d sby m aly n E The analyses show that the DNBR will not decrease below'the limitdays value e 6 at any time during the transients. Thus no core safety ; limit is violated as a result of a partial loss of forced reactor coolant flow. l l I i. 15.3-3 , t
The following cases have been analyzed: l Loops Initially Operating Punpa Coasting Down 4 4 l 3 3 The method of analysis and the assumptions made regt.rding initial operat-ing conditions and reactivity coefficients are identical to those dis-cussed in Section 15.3.1, except that following the loss of supply to all pumps at power, a reactor trip is actuated by either bus uadervoltage or bus undarfrequency. 15.3.2.2.2 Results The calculated seguance of events is shown in Table 15.3-2 for the two cases analysed. Figures 15.3-7 through 15.3-12 show the loop coast-downs, the reactor vessel coastdowns, the nuclear pow'er coastdowns sad the average and hot channel heat fluz coastdowns for the two cases. The reactor is assumed to trip on the undervoltage signal. The minimus DNBR for each case is McG6e g radar N, 444. 3./.67wyao( .+ h . a 15.3.2.3 Conclusions The analyses performed have demonstrated that for the complete loss of forced reactor coolant flow, the DNBR does not decrease below during the transient and thus, no core safety limit is violated. % 3.Is.+y % ms
- f. . . t
- L ua.
15.3.3 REACTOR COOLANT PUMP SHAFT SEI URE 15.3.3.1 Accident Description The accident postulated is an instantaneous seizure of a RCP rotor. Tiow through the af f acted reactor coolant loop is rapidly reduced. Isading to an initiaf.1on of a reactor trip en a low flow signal. This is considered to be a Condition IV event (limiting fault). 15.3-6 fl*$l
(2) Since the peak clad surface temperature calculated for the hot spot during the worst transient remains con-siderably less than 2700*F and the amount of . zirconium-
-water reaction is small, the' core will remain in place and intact with no consequential loss of core cooling capability. ,
I (3) Typically for a single RCP locked rotor accident, about 3 percent of the core fuel rods have a DNB ratio eP t e.es. 4% A Ag 4ps WPf996 Assuming that the activity from the pellet-h a d d "'-
- cladding gaps for these rods is released to the reactor coolant as a result of the locked rotor sce'ident,.the 1
radiological consequences are much less severe than l those analyzed for the Loss-of-Coolant Accident pre-seated in section 15.6.5. For the Design Basis Accident.- Loss-of-Coolant Accident, it was assumed thet 100 percent of the noble gas and iodina fission products' contained in
. the pelletwladding gaps was released to the Containment. l 1
l I 15.3.4 ' REACTOR C001. ANT PUMP SHAFT BREAK a The accident is postulated as an instantaneous failure of a RCP shaft. i Flow through the affected reactor coolant loop is rapidly' reduced, i though the initial rate of reduction of coolant flow is greater for the RCP shaft seizure event, and all other event characteristics are similar. ] Hence, the consequences of a shaft break event are bounded by those cal-I culated for the shaf t seizure event described in Section 15.3.3. l I i
)
I a { t i-q 1 4 l 15.3-11 fl'h? j
The energy release and the fuel temperature increases are relatively small. The thermal flux response, of interest for departure from nucleate boiling (DNB) considerations, is shown in Figure 15.4-2. The beneficial ef fect of the inherent thermal lag in the fuel is evidenced by a peak heat flux such less than the full power nominal value. There is a large margin to DNB during the transient since the rod surface heat flux remains below the design value, and there is a high degree of subcooling at all times in the core. -Figure 15.4-3 shows the response of the average fuel and cladding temperature. The average fuel tempera-ture increases to a value lower than the nominal full power value. The minimus DNBR st all times remains above the limiting value. The esiculated sequence of events for this accident is shew in Table 15.4-1. With the reactor tripped, the Plant returns to a stable condition. The Flant may subsequently be cooled down further by follow-ing normal Plant shutdown procedures. 15.4.1.3 Conclusions In the event of a RCCA withdrawal accident from the suberitical condition, the core and the RCS are not adversely affected, since the combinatica of thermal power and the coolant temperature result in a DNBR greater J an thodimiteeg value e M . The DNER design basis is described'in M*h Section 4.4; applicable acceptance criteria have been met. L y7g 15.4.2 UNCONTRO1.1.EDRCCABANKWITHDRAWALATPOWbR 15.4.2.1 Identification of Causes and Accident Description Uncontrolled 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 P reaches the relief or safety valve setpoint, there is a net increase in the reactoc coolant temperature. Unless terminated by manual or automatic l action, the power mismatch and resultant coolant temperature rise could eventually result in DNB. Therefore, in order to avert damage to the 15.4-7 J/L 2 3 o - _ __--- _
]
)
I ! I j l i fuel clad, the RPS is designed to terminate any such transient before-k ! the DNBR falls'below be @. Re_ se.fz+y W ywilem.+ 4 .-a, j This incident is classified as an ANS Condition II event (a fault of moderate frequency) as defined in Sectica 15 0.1. The automatic features of the RPS which prevent core damage following j the postulated accident include the followings j ! (1) Power range neutron flux instrumentation actuates a reactor trip if two out of four channels exceed an over-power setpoint. i (2) Reactor trip is actuated if any two out of four AT channels I 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 l exceed an overpower AT setpoint. This setpoint is auto-l matica11y varied with axial power imbalance to ensure that the allowable heat generation rate (kW/f t) is not exceeded. l (4) A high pressurizer pressure reactor trip actuated from any two out of four pressure channels, which is set at a ] I fixed point. This set pressure is less than the set f pressure for the pressurizer safety valves. l [ (5) A high pressurizer water level reactor trip actuated from I any two out of three level channels when the reactor power is above approximately 10 percent (Permissive-7). f In addition to the above listed reactor trips, there are the following RCCA withdrawal blocks: (1) High neutron flux (one out of four power range). j 15.4-8 f) O Y J
(2) overpower AT (two out of four). (3) overcemperature AT (two our of four). The manner in which the combination of overpower.and overtemperature AT trips provida protection over the full range of RCS conditions is described in Chapter 7. Figure 15.0-1 presents allowable reactor coolant loop average tempeuture and AT. for the design power distribution ard flow as a function of primary coolant pressurs. The boundaries of operation defined by the overpower AT trfp and the overtssperature AT trip are represented as " protection lines" on this diagram. The protec.- tion lines are drawn to include all adverse instrumentation and setpoint errors so that under nominal conditions trip would occur well within tha area bounded by these lines. The utility of this diagram is in the fact 1 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 DNBE ==4eM' #3 =+ 'W8-All points below ana to the left of a DMR line for a given pressure have [h,, M*' ' a DNBL1;p The diagram shows that DNR is prevented for all cases if the area snelosed with the maximum protection lines is not traversed by the app 11 cable DNBR line at any point. y emker- % +ke smh47 w\yons ie =se4. M.a a N 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). 15.4.2.2 Analysis of Effects and Consequences Method of Analysis: This transient is analyzed by the LOFTEANI ) code. This code simulates the neutron kinetics, RCS, pressurizer, pressurizer relief and safety valves, pressurizer spray, steen generator and stems generator safety valves. The code computes pertinent Flant variables including tampera- l tunes, pressures and power level. The core limits as illustrated in Figure 15.0*1 are used as input to LOFTRAN to determine the minisua DNBR during the transtant. 15 4-9 fi 'h 5
(5) The maximum positive reactivity insertion rate is greater than that for the' simultaneous withdrawal of the cosoina-tions 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 a decrease in margin to DNB. No singla active failure in any of the systems or equipment available to sitigate the effects of the accident will adverssly offset the con-sequences of the accidens. A discussion of anticipated transients without trip considerations is presented in Reference 4.- Results: Figures 15.4-4'and 15.4-5 show the transient response of neutron' flux, pressure average coolant temperature and DNER to 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 and pressure result and a large margin to DNB is maintained.. The design basis for DNBR is described in Section 4.4. The response of neutron flux, pressure, average coolant temperature and 4 i DNBR for a slow RCCA withdrawal from full power is shown in Figures 15.4-6 l and 15.4-7. Reactor trip on overtemperature AT occurs af ter a longer period, and the rise in temperature and pressure is consequently larger than for rapid RCCA withdrawal. Again, the minimum DNBR is Mda. t~.+ *1 p+aa. n m. s.Je ys s~ty vs . Figure 15.4-8 shows the minimus DNBR as a function of reactivity inser-tion rate from initial full power operation for minimum and nazicus reactivity feedback. It can be seen that two reactor trip channels l' f' provide protection over the whole range of reactivity insertion rates. These are the high neutron flux and overtemperature AT channels. The Im.+ h j minimus DNBR is _ _. u.U. g r**.4 OM N sd+7 e=We3 I 15 4-11 f).
- N L - - - - - - - - ----
Tigures 15.4-9 and 15.4-10 show the minimus DNBR as_a function of reac-tivity 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 i
which the overtemperature AT trip is effective is increased. In neither case does the DNBR fall below A ne. .pn A h m ty u , f .+- Wu . , The shape of the curves of minimus DNER vs reactivity insertion rate in the referenced figures is due both to reactor core and coolant syttes transient response, and to protection sy. stem action in initiating a reactor trip. Referring to Figure 15.4-9, for example, it is noted thats (1) For high reactivity insertion rates (ie, between
~ ~3
-2.5 x 10 ' 6K/see and 1.0 x 10 AK/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 races while core heat flux and coolant system cooperature lag behind due j
to the thermal capacity of the fuel and coolant system
/ fluid. Thus, the reactor is tripped prior to significant
/
increase in heat flux or water temperature with resultant high minimum DNBRs during the transient. As reactivity insertion rate decreases, core heat flux and coolant tem-paratures can remain more nearly in equilibrium with the neutron flux; minimum DNBR during the transient thus decreases with decreasing insertion rate. [ (2) The overtemperature AT reactor trip circuit initiates a reactor trip when the AT power measurement exceeds a setpoint based on measured RCS average temperature and pressure. This trip circuit is described in detail in Chapter 7. 15 4-12 Afl' h ] i J
i 1 v The reactor is-tripped sufficiently fast during the RCCA withdrawal at -) power transient to ensure that the ability of the primary coolant to f remove heat from the fuel' rods is not reduced. Thus, the fuel cladding l i temperature does not rise significantly above its initial value during i the transient. { The calculated sequence of events for this accident is shown in. l Table 15.4-1. With the reactor tripped, the Flant eventually returns to a stable condition. The Flant may subsequently be cooled down further ! by following normal Flant shutdown procedures. 15.4.2.3 Conclusions The high neutron flux and overtemperature AT trip chanaals provide adequate protection over the entire range of possible reactivity inser-tion rates; le, the minimum value of DNER is 2-- ,_ 'l O. grede +w.% % amis +y andyses (es.+ w.1%. ( 15.4.3 2CCA MISALIGNMENT 15.4.3.1 Identification of Causes and Accident Description RCCA misalignment accidents includes (1) A dropped full-length assembly. (2) A dropped full-length assembly bank. (3) Statically misaligned full-length assembly (see Table 15.4-2). (4) Withdrawal of a single full-length assembly. , Each RCCA has a position indicator channel which displays position of j 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. Group demand position is also indicated. 15.4-14
/Yd~ $11k l
L (b) Dropped RCCA group A dropped RCCA group typically results in a reactivity insertion of -1200 pcm which will be detected by che power range negative neutron flux rate trip circuitry. The reactor is tripped within approximately 2.5 see following the drop of a RCCA. The core is not adversely
~
affected during this period, since power is decreasing rapidly. Following reactor trip, normal shutdown procedures may subsequently be followed to further cool down the Plant. (a,) Statically utsaligned RCCA: The most severe misalignment situations with respect to DK51 at significant power levels arise from cases in which bank D is fully , inserted with one RCCA fully withdrawn; this ] represents a 12-ft misalignment error. ] hitiple independent alarms, including a bank insertion limit alarm, alert the operator well I before the postulated conditions are approached. The bank can be inserted to its insertion limit with any one assembly fully withdrawn without j
.he DNBR falling below h +>v= % e. d '+7
- ly b * *Ia** l 1
1 The insertion limits in the Technical Specifi-cations may vary from time to time depending on a number of limiting criteria. It is pref-erable, therefore, to analyse the miss11gned i RCCA case at full power for a position of the j control bank as deeply inserted as the criteria on minimum DNBR and power peaking factor (see l Section 4.4) will allow. The full power inser- i
. . l tion limits on control bank D are then choseu 15.4-21 Yb" h
i i Results: 1 For the single rod withdrawal event, two cases have been considered as follows: (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 j
' channel factor in the area of the withdrawing I J
RCCA. In terms of the over all system response, ] this case is similar to those presented in .) 1 section'15.4.2; however, the increased local power j
.)
peaking in the area of the withdrawn RCCA results in lower minimus DNBRs than for the withdrawn i hauk cases. Depending on initial bank insertion J and location of 'the withdrawn ECCA, eutomatic reactor trip may not occur sufficiently fast to 0 j prevent the miniaua core DNBR from falling below +k e. s*A4y I c.Ay s.s lia.+ V4.k4eM . Evaluation of this case at the power and o coolant conditions at which the overtemperature ) i AT trip would be expected to trip the Plant sbows j that an upper limit for the number of rods'with a DNBR is 5 percent. '] tess +% %e. .wfs+y == ty us I m +* *b*- 1 (b) If the reactor is in the automatic control mode, the multiple failures that resuit in the with- f i 1 drawal of a single RCCA vill result in the { immobility of the other RCCAs in:the control- ! ling bank. The transient will then proceed in the same manner as (a) described above. l l for the above cases a reactor trip will result, although not sufficiently fast in all instances j to prevent a miniana DNBR in the core of x.. L Id % O* t NE To11owing reactor trip, normal operating procedures T "' I T' 8 L ts.t vat u . I may be followed to further cool down the Pl.ent. { I 15.4-23 ; C_ _ _ _
15 4.3.3 Conclusions yrsadar- -l%.n +% d=+y sAst so$ y b-..& vst.,e It has been sho' that in'all cases'of dropped single RCCA, the DNBR remains at power and, consequently, drupped single assemblies do. not cause core safety limits to be violated. For all cases of dropped banks, the reactor is tripped by the power 6 range negative.asutron flux rate trip and consequently the DNER design criterion as described in Section 4.4 is set. For all cases of any bank inserted to its rod insertion limits with any single 1CCA in that bank fully withdrawn (static misalignment), ths DEBR
- remaina PrefM g ca ,4er
%., M a. 1>.f.hr my,., f.%.+ w.w.
i 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 powsr with bank D at the insertion limit, an upper _ bound of the number of fuel rods experiencing DNB6 is 5 percent of the total fuel rods in the core. - k 15.4.4 STARTUP OF AN INACTIVE REACTOR COOLANT FUMP (RCF) WITH , LOW HOT LEC TEK1ERATURE 1 15.4.4.1 Identification of Causes and Accident Description If the Plant is aparating 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 in. active loop is identical to the cold-leg temperature of the active loops (t!.e reactor core inlet temperature). If the reactor is operated at power. and assuming the secondary side of the ste:w generator in the inactive loop is not iso- l j lated, 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. 1 15 4-24 l ge si l 1 -
l i I difference between the core inlet cesperature and the inactive loop hot-leg temperature. Uncertainties in' initial conditions ) i are included in the limit DNBR as described is WCAP-8568. (2) Following the startup of the idle pump, the inactive loop flow reverses and accelerates to its nominal full-flow value. (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 appro-priate values for one pump out of service. In the analysis, reactor trip is conservatively. assumed to be actuated by the high neutron flux reactor trip. The trip setpoint' was assumed to be 118' percent of nominal " full power. In practice, however, reactor trip would be expected to occur on power range neutron fluz exceeding the P-8 setpoint (set at -75 percent of nominal full power). The F-8 setpoint will remain active until flow in the inactive loop reaches 90 percent of its nominal value. I No single active failure in any of the systems or equipment which are svailable to mitigate the effects of the accident will adversely affect the consequences of the accident. Results: The results following the startup of an idle pump with the above listed assumptions are shown in Figure 15.4-12. As shown in this curve, during the first part of the transient, the increase in core flow with l cooler water results in an increase in nuclear power and a decrease in core average temperature. The minimum DNBR during the transient is 1 l considerably See section 4.4 for a description of the DIGR design basis. m4er %w h M4 7 4 3,3 l ,+ 4 , , 15.4-26 1' N
- - ---- -- 9
.l q
Emactivity addition'for the inactive loop startup accident is due to the 1 decrease in core water temperature. During the transient, this decrease is due both to the increase in reactor coolant' flow and, as the inastive loop flow reverses, to the colder water entering the core from the hot-les side (colder temperature side prior to the start of the transient) l of the steen generator in the inactive loop. ;Thus, the reactivity insertien rate for this transient changes with time. The resultant core. I nuclear power transient, cesputed with seasideration of'both moderator j and Doppler reactivity feedback effects, is shown in rigure 15.4-12. y The calculated sequence of events for this e.cident is shown in Table 33.4-1. The transient results illustrated in Figure 15.4-12 indicate that a stabilised plant condition, with the reac' tor tripped, is approached rapidly. plant cooldown any subsequently be achieved by following normal shutdown procedures. 15.4.4.3 Cone 1usions i
)
The transient results show that the core is not adversely affected. I { There is considerable margin to the l' ' _ ~- 'r. F DMd4 fi # Va.lu. 1 15.4.5 A MALFUNCTION OR FAILURE OF THE FLOW CONTROLLER IN A BWR Loop THAT RESULTS IN AN INC****ED t h" TOR C00fAMT FLOW RATE This section is not applicable to the Tro.ian Nuclear plant. I 15.4.6 CVCS MALFUNCTION THAT RESULTS IN A DgCREASE IN THE BORON 66mCr.mIRATION IN THE em" TOR COOLANT 15.4.6.1 Identification of Causes s.nd Accident Description One of the two principal means of positive reactivity insertion to the core is the addition of unborated, primary grade water from the primary g aske ap water system into the BCS through the reactor askeup portion of the CYCS. Boron dilution with these systems is a manually initiated I l~ operation under strict administrative controls requiring close operator 4 surveillance with procedures limiting the dilution. A boric acid bland l Amendment !. 15.4-U (July 1986) fS 0
1
. system overpressure Analysis:
Because safety limits for. fuel damage specified earlier are not exceedad, there is little likelihood of fuel dispersal into the coolant. The pressure surge may therefore be calculated on the basis of conventional 1 beat transfer from the fuel and p'rompt heat generation in the coolant. The pressure surge is calculated by first performing the fuel heat trans-fer calculation to determine the average and hot spot heat fluz vs-time. ] 1 Using this heat flux data, a THINC (Section 4.4) calculation is conducted to determine the volume surge. Finally, the volume surge is simuisted in a Flant-transient computer code. This code calculates the pressure, anJ l
. transient taking into account fluid transport in the system,Vheat trans-fer to the steam generators, : ' " :--f- rf t'^ ;rrrrr-fr- r; ry xf ; :::ur: :1fr' - '--- No credit is taken for the possible pres- 3 sure reduction caused by the assumed failure of the control rod pressure housing.
1 15.4.8.2.2 Calculation of Basic Parameters
- i Input parameters for the analysis are conservatively selected on the }
hasis of values calculated for this type of core. The more important parameters are discussed below. Table 15 4-3 presents the parameters used in this analysis. j { Ejected Rod Vorths and Hot Chsanel Factors: The values for ejected rod worths and hot channel factors are calculated using either three-dimensional static methods or by a synthesis method l naploying one-dimensional and two-dimensional calculations. Stacdard nuclear design codes are used in the analysis. No credit is taken for the flux flattening ef fects 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 zenon distributions are considered in the calculation. 15.4-46
& 3y
.i )
l coefficient curvaa which are conservative' compared to actual design conditions for the Plant. As discussed above, no weighting factor is ] applied to these results. ! l The Doppler reactivity defect is determined as a function of power level using a one-dimensional steady-state computer code with a Doppler weight-1 ing factor of 1.0. The Doppler defect used is given in Table 15.4-3. j The Doppler weighting factor will increase under accident conditions, 5 as discussed above. Delayed Neutron Fraction, 8:
)
I Calculations of the effective delayed neutron fraction ( 8,ff) typically yield values of 0.70 percent at BOL and 0.50 percent at EOL for the first cycle. The accident is sensitive to 8 if the ejected rod worth is 18 as in zero power transients. In order to allow for future cycles, possi- I mistic estimates of 8 of 0.52 percent at beginning of cycle and 0.44 per- f cent' at and of cycle were used in the analysis. Trip Reactivity Insertion: 3.1. The trip reactivity insertion assumed to be + percen6 from hot full power f.L l and 4 percent from hot sero power 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 see after the high neutron flux trip point was reached. This delay is assumed to consist of 0.2 see for the instrument channel to produce a signal. 0.15 see for the trip breaker to open and 0.15 see for the coil to release the rods. The analyses presented are applicable for a rod insertion time of 2.2 sec l 1 from coil release to entrance to the dashpot, although measurements indicate that this value should be closer to 1.8 sec. The choice of such a conservative insertion rate means that there is over 1 see af ter the l trip point is reached before significant shutdown reactivity is inserted . into the core. This is particularly important conservatism for hot full 4 power accidents. { 15.4-48 i l~ \ l t
1 Results: The values of the parameters used in the analysis, as well as the results 1 of the analysis, are presented in Table 15.4-3 and discussed below. Berinning of Cycle. Full Fover
.1 Centrol beak D uns assumed to be inserted to its insertion limit. The worst ejected rod worth and het channel factor were 0.16 percent AK and 6.34 respectively; however, the analysis was done with more conservative <
.I o 6.*f o and an FO of + rep. The peak hot spot i
values, namely evet percent AKS3SO I n !
- clad average temperature was 3GPF. The peak hot spot fuel center M ac temperature exceeded the BOL life seit temperature of 6#00*F. Bowever, selting was restricted to (10 percent of the pellet. i l
Berinnine of evele. Zero Power: For this condition, control bank D was assumed to be fully insertad and C was at its insertion limit.- The worst ejected rod is located in con-trol bank D and has a worth of 0.520 percent AK and a hot channel factor of 11.7. Eowever, the analysia was done with more conservative values,
- o. 7s 2.o l anmely erP9S percent AK and an Fq of lev 6. The peak hot spot clad aqa i temperature reached only# F. -
End of Cycle. Full Powert Control bank D was assumed to be inserted to its insertion limit. The 7.70 ejected rod worth and hot channal factorsuwere e-0.21 percent AK and het, uo respectively. This resulted in a peak clad *'t erature of G M F F. The peak hot spot fuel temperature ascoeded the EOL melt temperature of 4400*F. Bowever, salting was restricted to <10 percent of the pellet. The variation in melt temperature with burasp is discussed in Section 4.4.1. 4 e Amendment 1 15 .4 -50 (July 1983) th - 3!,
l 1 End of Cycle. Zero Power: The ejei:ted 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 p
- c. 90 z1. o limit. The results were 4 46 percent AK and tert respectively. stLo The f ts r1 peak clad and fuel center temperatures were 3M&*F and 49H*F.
{ adegc. A summary of the cases presented above is given tu Table 15.4-3. The nuclear power and het spot fuel and clad temperature transients for the worst esses (50L full power and EOL sero power) are presented in Figures 15.4-21 through 15.4-24. 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 4.8.2.2, the reactor will remain suberitical following reactor trip. The ejection of a RCCA constitutes a break in the RCS, located in the reactor pressure vessel head. The effects and consequences of Loss-of-Coolant Accidents (LOCAs) are discussed in Section 15.6 5. Fo7. lowing the RCCA ejection, the operator would follow the same emergency instructions as for any other LOCA 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, <10 percent of the rods entered DNB based on a detailed three-dimensional THINC analysis ( ) . Although limited fuel melting at the hot spot was predicted for the full power cases, in practice melting is not expected since the analysis conservatively sasumed that the hot spots before and af ter ejection were coincident. Pressure Surte: A detailed calculation of the pressure surge for an ejection worth of one dollar at BCL, hot full power, indicates that the peak pressure does not exceed that which would cause stress to exceed the faulted condition 15.4-51
/L-37
.i 15.6.1'.2.2 Results Figure 15.6-1 illustrates the flux transient following the accident..
Reactor trip on overtemperature AT occurs as shown in Fizure 15.6-1. The pressure d4esy transient following the accident is given in Figure 15.6-2. The resulting DNBR never goes below as shown in. Figure 15.6-3. g ,, a i 15.6.1.3 Conclusions i The pressurizer low pressure and the overtemperature AT RPS signals provide adequate protection egainst this accident, and the minimum DNRR. remains t _ ____ .: *..- . 9 % fw- N tw dy wtyw le = f" Maae.. i 15.6.2 BREAK IN INSTRUMENT LINE OR OTHER LINES FROM REACTOR
' 1 COOLANT PRESSURE ROUNDARY THAT PENETRATE CONTAINMENT 1 Analysis of this event was not required as part of the original licens-ing basis for the Trojan Nuclear Plant. ;
l 15.6.3 STEAM CENERATOR TURE RUPTURE 15.6.3.1 Accident Description The accident examined is the complete severance of a single steam generator tube. The accident is assumed to take place at power with l l the reactor coolant contaminated with fission products corresponding to j continuous operation with a limited amount of defective fuel rods. .Ihe accident leads to an increase in contamination of the secondary system due to leakage of radioactive coolant from the RCS. In the event of a coincident loss of of fsite power, or failure of the condenser dump system, dischstge of activity to the atmosphere takes place via the steam generator safety valves and/or power-operated relief valves (PORVs). l In view of the fact that the steam generator tube sacerial is Inconel 600 j and is a highly ductile material, it is considered that the assumption i 15.6-3
/4-37 i
. _ _ _ _ ________-____a
l C - I 0 0 9 8 0 O. 1 f 0
% 9 4 5 . -
W48 41511 83-9 d n 0 9, 4 5 9 3 _ 3 4 9s%v4 9 14e4 E 3 2 - ? ! 2 3S39 z 0. T s3 -- 7, 7 s N % - 7 E D d n 2 1 0 2 4 6 4 2 e 5 e 4 9 4e o& c&F +- 0+t I C C E 1 1 W2 . h4 44 s37e t284 A F-7 7 N g n 0 0 o. O z r I n . f s T C i n 5
- 6 t CLJa3-M, 2 5 94: W P 3 +et4 S E n % 91 3G9iH149 J
F i g 0 9 2 - 9 2 343S* ism e 9 A B C
. C F 0 R 8 3 1 0 1 .
g 0 3 E H n l 3 3 4 a
- T i % ' e 4 n n
2 0 2 & 0 2 5
- r 23 07 S 9s2 9 s5 MP53- 6 trse F t s9?1 r C 7 5
1 O i g 1
+ 9 2 J 4 y4y UPeb S e E I B L S B Y A L T A F N F
- A * =
. g E g . e /
H i e r l T % i t r ut a N t u c
, h t a _
_ I n g a a r r e y D K o i P e E e A it e n C g e p g f w K o n R a p m r S r m e en U i 1 c a y A it oi g e e t L , r t 1 c t n v t e S e c i a e R n h f i v , j e t r g d E i t T r n i y e j a t F e a e e o o t t e r e* t r r E e n e o - M m w rt c i d p l l . e v t A i a v o d R T l d u e i rt o o ree c a pr s A e o e r v r n c e f u l d l __ F k a r r o l t e a eu e ea u l l d d c e o e e ey ba r _ f t r ur f c f r t e f e f e e c a d p b a b p . . . m nm m m x w e i _ - o ej le e r 9 u a e a a e F E D F T F F N Ht h H H _ N3g.3 - ss d% CE<
. "' y
_ ' !llL
a
;t f.
l l: t
- 4 3.
! A 2.5000 - - i, 2x 2.0000 - -
= -
$I ..
I ! 38 g,$oco . .
$5 W: .
EW -- l ! $ 1.0000'- - ! 4
.50000 - -
\' '
8 O.0 - 8 e a 1
* *
- a a i a.
5
~
~ e ~
a TIME (SEC) l l Q lLt.,k fdi'f}l [0A AM/N$ lA$E \ Figure 15.4-21 Nuclear Power Transient, BOL EFP, l Bod Ejection Accident
.- Amendment 1
- (July 1983)
$' D A-__-_________ __
j
5 e ,
.g 3.0 2.5 ,
7 E
, , .: '" 2.0 o
- E e 1.5
$.2 '
8t
- 2 1.0 6
. 0. 5 .-
i 0 i l I i i i l I l- 1 l 2 3 4 5 6 7 8 9 10 0 1
. TIME (Seconds)
' i Figure 15.4-21 Nuclear Power Transient, B01., HFP, Rod Ejection Accident ;
t
- - - - a
k' 1 t 6000.0 : : : LTING N " 5000.0 3 5080' F ; n FUEL CENTER E TEMPERATURE 4000.0 4 - h g FUEL AVERAGE w ~TEMPERATUtY . - g 3000.0 - -
- i' CLAD 2000.0 .
OUTER TEMPERA E 1000.00 - l 0.0 i TIME (SEC) 6,LQ,,, id 10/ [0A A 0OlA l $ fAdf Figure 15.4-22 Hot Spot Fuel and Clad Temperature vs Time, BOL HTP, Rod Ejection Accident Amendment 1 (July 1983) :
r g h 3 a 4 6000 i 5000 ---- , l _ MEL.T.I.NG 4900.'.T. ..
- - FUEL CENTER T W ERATURE .
$ 4000- -
w. E FUEL AVERACE , T M ERATURE ! 2 3000 l 5 g 1 w
" 2000_ CLAD OUTER TEMPERATURE {
1000 1 i i 4 l l l l 1 i l i 0 1 2 3 4 5 6 7 8 9 10 TIME (Seconds) l
^
Figure 15.4-22 Hot Spot Fuel and Clad Temperature vs ;
- Time, BOL HFP, Rod Ejection Accident l i
l l F
b 1 1 6937-66 ) I 10-2 gou j l i 10-8 - - 103 1 1 i 10-" - - 102 1E as
- % 10 - 10 8
5 f' c W f er 10-8 : - 100 I 10 - - 10-8 I 10 - g
- 10-2 io-e i I I i I in-
.5 1.0 1.5 2.0 2.5 10.0 TI E (SECONDS)
Figure 15.4-23 ' Nuclear Power Transient E01, HZP Rod Ejection Accident
- m. \ 6,A R. (A) /T/f [0U bN/N6 fA$E llL ' YY
_______._m-_.____ _ . . _ - . _ _ _ _ . _ _ _ _
i l 102 j I id1-100 _ 10-L 1 g 10-L 3 I10-L o e 10 4 5 510-L s g 10- ' 8 10 10-8. 10-' g g g g g i j g l 0 1 2 3 4 5 6 7 8 9 10 TIME (Seconds) Figure 15.4-23 Nuclear Power Transient EOL HIP Rod Ejection Accident
'l 6917 1 d
.6000 m
w0o - o ^
-- -- Z.
4000 - f L CENTER TEMPERAT E q
- I
( FUEL AVG TFMP TURE l CLAD TEMPERATWR 2000 . - 4 L 3 I 1000 i - l j e 1 0 i i i ! 0 / l.0 2.0 3.0 4.0 10.0 TIME (SECONDS) l Figtire 15.4-24 Bot Spot Fuel and Clad Temperature versus Time, EOL RZP Rod Ejection Accidest fg p A, L L t<> l1~h/ [04/bNINS PAdf WL%
[' I li 6000 5000 -___ _ _ _ _ _ _ _ _ _ _ _ 9 TF L'5 1*L _ _ _ l m cm mmm E4000-TUEL AVG h
- 3000 _ T M ERATURE E
w CLAD TEMPERATURE h 2000 _. I 1000 _ I O l l l l l 1 i i 1 0 1 2 3 4 5 6 7 8 9 10 !
.i l
TIME (Seconds) 1 I
/
1 Figure 15.4-24 Hot Spot Fuel and Clad Temperature versus Time, EOL HZP Rod Ejection Accident I + ) I 1
\
$L-Y7
_ _ _ _ _ - _ _ _ _ _ . _ . _ __.____________________m__.m___._______
i < ;;r) . {_ 1 l
< 'i G
1 1
. ,1 . 4
( k 9 SECTION 5.2. LOSS OF COOLANT ACCIDENTS 1 i I 1 + I
)
!;'f .. 4 l L 5.2 Loss-Of-Coolant Accidents (LOCA)
.c .
5.2.1 .Large Break LOCA
.5.2.1.1- Introduction-This section reports the results of an. analysis which was performed to demonstrate that the Trojan Nuclear. Plant cenforms to.the' requirements of
-Title 10 of the Code'of Federal Regulations, Part 50, Section 46 (10CFR50.46, Reference 13) in accordance.with Appendix K (10CFR50, Appendix X) for Large Break Loss-of-Coolant-Accidents (LOCA) for a nuclear enthalpy rise hot channel
-factor.(F3g) of 1.62.
m 5.2.1.2 Background and Assumptions Pursuant to an increase in F3g from.1.55 to 1.62, a Large Sreak LOCA analysis was performed for the Trojan Nuclear Plant to demonstrate the continued conformance of this unit to the Acceptance Criteria of 10CFR50.46. In addition to the increased FAH, the analysis also assumed a core thermal
' power level of 102% of 3558 Megawatts thermal-(MWt) and 11.5% uniform Steam Generator Tube Plugging (SGTP). The core was assumed to consist >of 193 assemblies of Westinghouse 17 x 17 STANDARD (STD) fuel.
5.2.1.3. Method of Analysis The analysis was performed using the Westinghouse BASH Evaluation Model, (Reference 14), for a full spectrum of break discharge coefficients (CD* 0.4,.0.6 and 0.8). The Westinghouse BASH Emergency Core Cooling System (ECCS) Large Break Evaluation Model was developed to determine the Reactor Coolant System (RCS) response to design basis large break LOCAs and consists of the SATAN-VI, WREFLOOD, COCO, BASH and LOCBART computer codes. The SATAN-VI code. (Reference 15), was used to generate the blowdown portion of the transier;t. The WREFLOOD code, (Reference 15), was used to calculate the refill portion of the transient. The COCO code, (Reference 17), operates interactively 5-10 wn e-sma L .-
, , n c-~
I + l!, p , with the WREFLOOOl code to evaluate the containment pressure response. The f BASH code, (Reference 14), .was used to calculate the system hydraulics during o ,1 the reflood phase and cladding ;the-mal analyses were performed with the ' :LOCB4RT. code. The LOCBART code, (Reference 14), .is a synthesis of the LOCTA-IV-(Reference to) and SART, (Reference'19), codes and uses'the RCS
- pressure, fuel rod power history.. steam flow past the uncovered part of-the core, and mixture height history from the SATAN-IV, WREFLOOD and BASH codes as input. The hydraulic.' analyses and core thermal transient analyses assumed 102. ~
, percent of core thermal power. The pumped ECCS injection was assumed to be delivering to the RCS 27 seconds after the generation of a safety injection signal. The 27-second delay includes time required for diesel start up and
. loading.of the ECCS pumps onto the emergency buses. Minimum safeguards
. Emergency Core Cooling System capability and' operability has been assumed in the full' spectrum analysis. .However, since the assumption of maximum safeguards.has proven ~ to be more limiting for some Westinghouse four-loop,-
non-upper-head injection, non-burst node limited plants, a maximum safeguards case for the most limiting discharge coefficient was also performed. 5.2.1.4 Results The transient was considered to be terminated when-the hot rod clad average
' temperature " turned around" (i.e., hot rod clad average temperature began to decline) indicating that the peak clad temperature had been reached. The analysis determined a peak clad temperature of 1983*F for the limiting discharge coefficient (CD = 0.6.) assuting minimum safeguards at a total peaking factor (Fg ) of 2.50. Tae calculated peak clad temperatures for the The CD = 0.4 and CD = 0.8 cases were 1892*F and 1880*F, respectively.
maximum safeguards case for the most limiting discharge coefficient (CD' O.6) resulted in a peak clad temperature of 1916*F, demonstrating the minimLm safeguards case to be most limiting. The final peak clad temperatures for ali the cases were below the 2200*F Acceptance Criteria limit established by Appendix K of 10CFR50.46. 5-11 00137:6-671022
v g s 5.2.1.5 Conclusions for. breaks up to and inclucing the double-ended severance' of 'a reactor coolant
' pips, the emergency cere cociing system will meet the acceptance criteria as presented in 10CFR50.46. That is: .
- 1. The calculated peak fuel element clad temperature is below the-requirement of 2200*F.
~
2.- The amount of fuel element cladding that reacts chemically with water or steam does not exceed one percent of the total amount of zircaloy. in the reactor. q
< , j
- 3. The localized cladding oxidation limit of 17 percent is not exceeded
~during or afte'r 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. This is required to remove the heat from the long-lived reactivity in the core.
The Large Break LOCA analysis for the Trojan Plant, utilizing the BASH model, resulted in a peak clad temperature of 1983*F for the limiting break case at a total' peaking factor of 2.50. The maximum local metal-water reaction was 4.46 percent, and the total metal-water reaction was less than 0.3 percent for all cases analyzed. The clad temperature turned around at a time when the core geometry is still amenable to cooling. Criteria 5 is addressed separately in a specific evaluation for each reload cycle. . i A more complete description of the Large Break LOCA analysis and results are contained in Section 5.2.3 which contains the modifications to the Trojan FSAR Chapter 15.6.5 for the fuel upgrade. l l 5-12 aave-mm i i h_ _ __ _ _ - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ __
7 5.2.2. Small Break LOCA-5.2.2.1 Introduction . This:section reports the results of an analysis.which was performed to demonstratEth'attheTrojanNuclearPlantconformstotherequirements'~of Title 10 of the Code of Federal Regulations, Part 50,.Section 46 (10CFR30.46, < Reference 13) in accordance with Appendix' K (10CFR50, Appendix K) and Item
'II.K.3.30'of NUREG-0737 for Small Break Loss-of-Coolant-Accidents (LOCA) for an F AH of 1.62.
5.2.2.2 Background and Assumptions Pursuant to an increase in F AH fr a 1.55 to 1.62, a Small Break LOCA analysis was perfonned for the Trojan Nuclear Plant to demonstrate the continued conformance of this unit to the Acceptance Criteria of 10CFR50.46. In Addition to the increased FAH, the analysis also assumed a core thermal power level of 102% of 3558 Megawatts thermal (MWt) and 11.5% uniform Steam Generator Tube Plugging (SGTP). The core was assumed to consist of 193 assemblies of Westinghouse 17x17 STANDARD (STD) fuel. The power shape used for the Small Break LOCA analysis was chosen to allow an increase in the peaking factor at the top of the core from 1.0 to 1.5. 5.2.2.3 Method of Analysis The analysis was perfonned using the Westinghouse NOTRUMP Small Break Evaluation Model, (Reference 20), for a full spectrum of break sizes (2 in., 3 in, and 4 in.). The Westinghouse NOTRUMP Emergency Core Cooling System (ECCS) Small Break Evaluation Model was developed to determine the Reactor Coolant System (RCS) response to design basis Small Break LOCAs and consists of the NOTRUMP and LOCTA-IV computer codes, References 21 and 18 respectively. 5-13 som -onen
$ 1
! k'
(
.The NOTRUMP code was used te calculate the~ system hydraul.ics throughout the:
transient. -The LOCTA-IV-coce which calculated the' cladding thermal response uses the RCS pressure, fusi od power history, steam flow past' the uncovered
-part of the core, and mixture height history from the NOTRUMP code to
' determine the peak clad . temperature during the Small Break' LOCA. The hydraulic analyses'and core thermal transient analyses assumed 102. percent of:
- core termal. power. The pumoed ECCS' injection was. assumed to be delivering to the RCS 25 seconds after the generation of a safety' injection signal. - The 25-second dela) :ncludes time required for diesel' start-up and loading of the ECCS pumps onto the emergency l buses. Minimum safeguards Emergency Core Cooling System capability and operability has been assumed in the analysis.
5.2.2.4' Results- , ' The transient was considered to be terminated when the hot rod clad average temperature " turned around"-(i.e. - hot rod clad average temperature began'to-decline) indicating that the peak clad. temperature had been reached. The analysis determined a peak clad temperature of 1925'F for the_ limiting break
'(3 in.) consistant with an Fg snvelope based on a total peaking factor of 2.50. -The calculated peak clad temperatures for.the 2 in. and 4 in. breaks' were 1060*F, and 1402*F, respectively. .All peak clad temperatures for the Small Break analyses were below the values determined by the large Break analyses and were below the 2200*F Acceptance Criteria limit established by 10CFR50.46.
5.2.2.5 Conclusions In the event of a Small Break LOCA the emergency core cooling system provides 1 adequate core protection by satisfying the acceptance criteria of 10CFR50.46. That'is: j i
- 1. The calculated peak fuel element clad temperature is below the -
requirement of 2200*F. I 5-14 oom mn i _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . _ _
4-I
~,
6
- 2. The amount of fuel element cladding .that reacts chemically with water or steam does not exceed one percent of the total. amount of zircaloy in the' reactor.
3.- The : localized cladding oxidation limit ef 17 percent is.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. .This is required to~ remove the heat from the long-lived reactivity in the core.
The Small Break LOCA analysis for the Trojan Plant,' utilizing the NOTRUMP
^
model, resulted in a peak clad temperature of 1925'F for the. limiting break case for an Fg envelope based on a total peaking factor of 2.50. The maximum local metal-water reaction was 4.59 percent, and the total metal water L reaction wcs less than 0.3 percent for all cases analyzed. The clad temperature turned around at a time when the core geometry is still amenable to cooling. Criteria 5 is addressed separately in a specific evaluation for each reload cycle. A more complete description of the Small Break LOCA analysis and results is contained in Section 5.2.3 which contains the modifications to the Trojan FSAR Chapter 15.6.5 for the fuel upgrade. 5.2.3 LOCA Chapter 15.6.5 FSAR Modifications 5-15 ann e-mm
't 3 9
., e o
. .s .
, _ u. f-:
- f. ~
I1 l- ., 6 . r .. .o n :. i
#.' i
' r LARGE At0 SMALL BREAK'LOCA-FSAR CHANGES ll
.1 e
. L 'i
- 11. .
( I 'I k l 4 1 1 l l
(. 15.6.5- LOSS'OF REACTOR COOLANT ACCIDENTS RESULTING FROM A SPECTRUM OF POSTULATED PIPING BREAKS WITHIN THE REACTOR COOLANT PRESSURE EuUnBARY is.s.s.1 marident namerintions A Ioss-of-Coolant Accident (IDCA) is the result of a pipe rupture of the Reactor Coolant System (RCS) pressure boundary. For the analyses
-report ~ed here,.a ma j or ip pe break (large 'oreak) is defined as a rupture with a total cross-sectional area equal to or greater than 1.0 square 2
foot (ft ).
;This event is considered an American Nuclear Society (ANS) condition 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 design basis. See Section 15.0.1.4 for a discussion of l
. Condition IV events.
l A minor pipe break (small break),,as considered in this section, is defined as a rupture of the reactbr coolant pressure boundary with a 2 total cross-sectional area less than 1.0 ft in which the normally operating charging system flow is not sufficient to sustain pressurizer level and pressure. This is considered a Condition III event, in that it is an infrequent fault which may occur during the life of the plant. See section 15.0.1.3 for a discussion of condition III events. The Acceptance Criteria for the LOCA are described in Title 10 of the Code of Federal Regulations Part 50 Section 46 (10CFR50.46; Reference
- 4) as follows: .
- 1. The calculated peak fuel element clad temperature is below the requirement of 2200*F.
- 2. The amount of fuel element cladding that reacts chemically with water or steam does not exceed 1 percent of the total amount of zircaloy in the reactor.
P/
g i
- 3. The clad temperature transient is terminated at a time when'the 4 core geometry is still amenable to cooling. The localized cladding-oxidation limit of 17 percent'is 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. This is required to remove the~ heat from
; the lorg lived radioactivity remaining in the core. !
1I
- These criteria were, established to provide significant margin in
. Emergency Core Cooling System (ECCS) . performance following a IDCA. .
Reference 3 includes a study of the probability of the occurrence of RCS pipe failures.
~In all cases, small breaks (less than 1.0 ft2 ) yield results with more margin to the Acceptance criteria limits than large breaks.
Should a major break occur, depressurization of the RCu results in a pressure decrease in the pressurizer. The reactor trip: signal
- subsequently occurs when the pressurizer low pressure trip setpoint is reached. A safety injection signal is generated whhn the appropriate setpoint.is reached. These countermeasures will limit the consequences of the accident in two ways:
(1) Reactor trip and borated water injection complement void formation in the core and cause a rapid reduction of nuclear power to a residual level corresponding to the delayed fission and fission product decay heat. However, no credit is taken during the IACA blowdown for negative reactivity due to boren content of the injection water. In addition, the insertion of control rods to shut down the reactor is neglected in the large break analysis. 1-2 1
T i J (2) Injection of borated water provides the fluid medium for heat transfer from the core and prevents excessive clad temperatures. 15.6.5.1.1 Description'of-Lnrge-Break IOCA Transient The time sequence of events following a large break Inca is. presented
~ ~ in Table 15.6-5. .
Before the break occurs, th4 unit is in an equilibrium condition, i.e., a the heat generated in the core is being removed via the secondary system. 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 nuclects boiling. Thereafter, the core heat transfer is based on local conditions with transition boiling, film boiling, and forced convection to steam as the major heat transfer mechanisms. The heat transfer between the RCS and the secondary systen may be in-either direction depending on the relative temperatures. In the case of continued heat addition to the secondary, secondary system pressure increases and the atmospheric relief and/or main steam safety valves may actuate to limit the pressure. Makeup water to the secondary side is automatically provided by the Auxiliary Feedwater System. The safety injection signal actuates a feedwater isolation signal which
' isolates normal feedvater flow by closing the main feedwater isolation
. valves and also initiates auxiliary feedwater flow by start'ing the
-auxiliary feedwater pumps. The secondary flow aids in the reduction of RCS pressure. l I
L-3 \ 1
F p , 4 1
' When theLRCS.depressurizes to approximately 600 pounds per square inch q absolute (psia), the cold, leg accumulators begin to inject borated water'into the reactor coolant loops.
1 s since the loss of offsite power-is assumed, the reactor coolant pumps are assumed to trip at the time of reactor trip during the accident. l _ The effects of pump coastdown are included in the blowdown analyses.
~
)
The blowdown phase of the transient' ends after the RCS pressure (initially assumed at a. nominal 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 j
- emergency core cooling water injected into tho'RCS are calculated not to be' effective. At this time (called end-of-bypass) refill of the -
~
reactor vessel lower planum 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 temperature r.ise has been terminated. . Frou. the later stage of blowdown and then the beginning-of-reflood, the safety injection accumulator tanks rapidly discharge borated cooling water into the RCS, contributing to the filling .cf the reactor vessel downconer. The downconer water elevation head provides the driving force required for the reflooding of the reactor core. The centrifugal charging, safety injection, and RHR pumps aid in the filling of the downconer and subsequently supply water to maintain a full downconer and complete the reflooding process. L - ?'
~
y i '.
-Continued operation of the ECCS pumps supplies water during long term-cooling.. Core temperatures have been reduced.to long-term steady state levels l associated with dissipation of residual heat generation. After-the water level' of the refueling water storage tank: reaches a minimum l
' allowable value, coolant for long-term cooling of the core'is obtained by. switching to the cold leg recirculation phase of operation in which spilled borated water is' drawn from the containment sump by the.
residual heat removal pumps and returned to the RCS cold legs. The containment Spray System continues to operate to further reduce-containment pressure. Approximately 13 hours after initiation of the IOCA, the ECCS is realigned to supply water to the RCS hot legs in 1 order to control the boric acid. concentration in the reactor vessel. 15.6.5.1.2 Description of Small Break ICCA Transient The time sequence of events following a small break LCCA is presented in Table'15.6-6. Rupturesofsmallcrosssectionwillcauseb.eakageofthecoolantata rate which can be accommodated by the charging pumps. These pumps 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 for the power and fuel history prior to the event. 1 The maximum break size for which the normal makeup system 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 Gystem pressure, i.e., 2250 psia. A makeup flow rate from one charging pump is adequate L to sustain pressurizer level at 2250 psia for a break through a 0.37:5
' inch diameter hole. This break results in a loss of approximately 17.25' pounds mass per second (1b/sec).
L-5 __--___ _ ~
[. . j '
.n Yi,
- D ' '
L i Should a la'rger break occur, depressurization of the Reactor Coolant . System causes-fluid to flow into the loops from the pressurizer
!resulting in a pressure and level decrease in the pressurizer. Reactor l- trip occurs when the low pressurizer pressure trip setpoint is
< . reached.- During the earlier part cf the small break transient, the effect of the break flow is not strong enough to overcome the flow maintained by the' reactor coolant pumps through the core'as-they are coasting down following reactor trip. Therefore, upward flow through the core is maintained. The ECCS'is actuated when the appropriate setpoint is reached.
Before the break' occurs the plant is in an equilibrium condition, i.e.,
.the heat generated in the core is being removed via the secondary systaa. . During blowdown, heat from fission product decay, hot internals, and the vesd1 continues to be transferred to the Reactor.
Coolant System. Tha heat transfer between the Reactor Coolant Syster and the secondary system may be in either direction depending on the relative temperatures. In the case of continued heat addition to the-
' secondary, secondary system pras.sure increases and stama relief via the atmospheric relief and/or safety valves may occur. Makeup to the secondary side is automatically provided by the auxiliary feedwater pumps. The safety injection signal isolates normal feedwater flow by closing the main feedwater isolation valves and initiates auxiliary
- feedwater flow by starting the auxiliary feedwater pumps. The secondary flow aids in the reduction of Reactor Coolant System pressure.
When the RCS depressurizas to approximately 600 psia, the cold leg accumulators begin to inject borated water into the reactor coolant
. loops. The vessel mixture level starts to move up to cover the fuel before the accumulator injection for most breaks. For all breaks, the 1
f
l 4 accumulator injection provides.enough water supply to bring the mixture level up' to the upper plenum region where it is maintained. Due to the loss of offsite power assumption, the reactor coolant pumps.'are assumed to be tripped at.the time of reactor trip during the accident and the effects of pump coastdown are included in the blowdown analyses. l 1 15.6.5,2 Analvais The requirements.of an acceptable ECCS evaluation model are presented in Appendix K of 10CFR50 (Reference 4). The requirements of Appendix K
, regarding specific model features were met by selecting models which .
'I provide a significant overall-conservatism in the analysis.- The i
-astfuaptions made pertain to the conditions'of the reactor.and associated safety systen equipment at the time that the IDCA occurs 'and include.such items as the core peaking factors, the containment pressure, and the performance of the ECCS. Decay heat generated-
~
throughout the transient is also conservatively calculated as required by Appendix K'of 10 CFR 50. The NRC approved Westinghouse ECCS Evaluation Model assumes single failure of RCCS equipment as required by 10 CFR 50, Appendix K, Section + I.D.1. The ICCS systen design and the Westinghouse system design criteria satisfy the General Design criteria, Appendix A, of 10.CFR
- 50. Single active failures, as defined by ANSI N658, have been considered in the ECCS analyses. For the Large Break analysis, the loss of a low-head ECCS pump is the assumed worst single failure. For the Small Break IDCA analyses, the most limiting ringle failure is the assumed loss of one train of pumped ECCS'due to the loss of an emergency diesel generator. The analyses assume the loss of offsite power coincident with the IDCA.
L-7
p 115.6.5.2.1- Large. Break-IOCA Evaluation Model The analysis of a large break IDCA transient is divided into three phases: 1) blowdown, 2) refill, and 3) reflood. There are three l L distinct transients analysed in each phase: 1) the thermal-hydraulic transient in the RCS, 2) the pressure and temperature transient within the containment, and 3) 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 IDCA.
The description of the various aspects of the IOCA analysis methodology is'given in References 5, 10, 11, 16, 17, 19, and 20. These documents describe the major phonemana modeled, the interfaces among'the computer codes, and the features of the codes which ensure compliance with the Acceptance Criteria. The SATAN-VI, WREFIDOD, COCO, and ICCTA-IV codas which are used in the LOCA analysi,s, are described in detail in References 6 through 9. Modifications to these codes are specified in References 10 through 12. The BART code is described in References 16 and 17. Modifications to the BART code and BART methodology are l described in References 21 and 22. The BASH and LOCBART codes are described in References 19 and 20. These codes are used to assess the I core heat transfer geometry and to determine if the core remains amenable to cooling throughout the blowdown, refill, and reflood phases l of the IDCA. The SATAN-VI computer code analyzes the t'hermal-hydraulic transient in the RCS during blowdown and the WREFIDOD and BASH computer codes are used to calculate this transient during the refill and reflood phases of the accident. The BART computer code is used to
-calculate the fluid and heat transfer conditions in the core during '
reflood. The COCO computer code is used to calculate the containment pressure transient during all three phases of the LOCA analysis. similarly, the IDCBART computer code is used to compute the thermal l* transient of the hottest fuel rod during the three phases. Z-?
The large ~ break arialysis was performed with the approved December,1981
. version of the Evaluation Ngd g (Reference 11), with the approved version of ' BASH (ReferenceI 20) .
SATAN-VI is used to calculate the RCS pressure, enthalpy, density, and 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 IDCA. 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 l phases, these data are' transferred to.the WREFLOOD code. Also at the
.and-of-blowdown, the mass and energy release rates during. blowdown'are
. transferred to the COCO code for use in the determination of the containment pressure response during these phases of the IDCA.
Additional SATAN-VI output data from the end-of-blowdown, including the core pressure, and the core power. decay transient, are input to the IDCSART code. With input from the SATAN-VI code, WREFIDOD uses a system thermal-hydraulic model to determine the core flooding rate (i.e.-, the rate at which coolant enters.the bottom of the core), the coolant pressure and temperature, and the quench front height during the refill phase' of the IDCA. WREFIDOD 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 WREFIDOD and COCO codes are interactively linked. BASH is an integral part of the ECCS evaluation model which provides a meme realistic thermal-hydraulic simulation of the reactor core and RCS during the reflood phase of a LOCA. Instantaneous values of I accumulator conditions and safety injection flow at the time of ; e l A-7 \ l
<- )
y 1 ", 4 I completion of lower plenum refill are provided' to BASH by WREFI40D. BASH has been substituted for WRETIAOD in calculating transient values . L of core inlet'. flow and enthalpy for the detailed fuel rod model, 1 l
-IDCBART. A more detailed description of the BASH code is available in Reference 19. The BASE code provides a e sophisticated treatment of. steam / water flow phenomena in the reactor coolant system a during core reflood. A more dynamic-interaction between the core thermal-hydraulics and system behavior is expected,..and recent experiments have borne-this out. In the BASH code reflood model,'BART
.j provides the entrainment rate for a given flooding rate, and.then a i system model determines loop ~ flows and pressure drops in response.to
'the calculated core exit flow.- An updated inlet flow is'used to
! calculate a new entrainment rate. This system produces e,very dynamic flooding transient, which' reflects the close coupling between core l thermabhydraulics and loop behavior. f The Coco code is a mathematical model of the containment. Coco is run .)
,using mass and energy releases to'the containment provided by SATAN and WREFIDOD. - COCO is described in detail in Reference 8. Calculated-
'l pressures from Coco are presented in Figures 15.6-48 through 15.6-50 for each of the analyzed discharge coefficients.
The IACBART code is a coupling of IDCTA-IV and BART. The I4CTA-IV code is a computer program that evaluates fuel, cladding and coolant , temperatures during a LOCA. A more complete description than is presented here can be found in Reference 9. In the I4CTA detailed fuel rod model, for the calculation of local heat transfer coefficients, the empirical'FLECHT correlation is replaced by the'BART code. BART employs rigorous mechanistic models to generate heat transfer: coefficients. appropriate to the actual flow and heat transfer regimes experienced by. the IDCTA fusi rods. This is considered a more dynamic realistic approach than relying on a static empirical correlation. l 1 L -/o ,
3' ,. 4 -15.6.5.2.2 sma11' Break IDCA Evaluation Model The small-break analysis was performed with the approved . Westinghouse l ECCS Small Break. Evaluation Model (References 9, 14, and 18).. 3 1 l d es are t er co used in the analysis of, "The NOTRUMP and IDCTA-IV compu loss-of-coolant accidents due to small breaks in the Reactor Coolant System. The NOTRUMP computer code is a state-of-the-art. one-dimensional general network code consisting of a number of advanced q features. Among these features are the calculation of thermal l non-equilibrium in all fluid volumes, flow regine-dependent drift flux. calculations with counter-current flow limitations, mixture level f
.j tracking logic in multiple-stacked. fluid nodes, and regine-dependent.
]
heat transfar correlations. The' NOTRUMP small break IDCA' energency'~ core cooling system (ECCS).svaluation model was developed to determine the RCS response to design basis small break IDCAs and to Eddress the - Nuclear Regulatory Commission (NRC) concerns axpressed in NUREG-0611,
~
" Generic Evaluation of Feedwater Transients and Small Break i Ioss-of-coolant Accidents in Westinghouse Designed Operating Plants."- l In NOTRUMP, 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 throughout the system. A detailed description of NOTRUMP is given in References 14 and 18.-
The use of NOTRUMP in the analysis involves, among other things, the representation of the reactor core as heated control volumes with an 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. ; l
- j l
4-//
b ' cladding thermal. analyses are performed with the IACTA-IV (Reference'9) -l
; code which uses the RCS pressure,' fuel rod power. history, steam flow past'the-uncovered part of the core, and mixture height history from j the NOTRUMP hydraulic calculations, as input. l 15.6.5.1 Beaulta 15.6.5.3.1 Iarge Break Results-Based on the results of the I4CA sensitivity studies (Reference 15),
the. limiting large break was found to be the double ended cold leg guillotine (DECIA) . . Therafore, only the DECIA break is considered . in-the large break ECCS performance analysis. Calculations were' performed' for a range of Woody break discharge coefficients under min N w . safeguards conditions.- Minimum safeguards conditions' assume the worst single failure of ECCS ogaipment, which for. the Large Break I4CA is the . loss of one low-head ECCS pump. However, for some four-loop, a i non-upper-head injection, non-barst node limited plants with a Eastinghous Hsss design, the current nature of the Appendix K ECCS Evaluation Models is such that it may be more limiting to assume maximum possible ECCS delivery resulting from no single failure of the ECCS. Thus, after a range of break discharge coefficients under minimum safeguards conditions were analyzed, the limiting discharge coefficient was, analyzed with maximum safeguard conditions. Table j 15.6-4 lists important input parameters and initial conditions used in f the large break analyses. The results of these calculations are i summarized in Tables 1S.6-5 and 15.6-7. The het spot is defined as the location of maximum peak clad temperature. The location is specified in Table 15.6.7 for each break analyzed. The location is indicated in feet which represents elevation l above the bottom of the active fuel stack. 1 a_-______:_________-
li i 1 Table 15.6-8 presentsfa summary of the various Containment systems-parameters and structural. parameters which were used as input'to the COCO computer code used for this analysis. Tables 15.6-9'and 15.6-10 present reflood mass'and energy releases to the Containment and the broken loop accumulator mass and' energy release to the Containment. Figures 15.6-9 through 15.6-55 present the parameters of principal' interest from the large break IcCA analyses. Transients of the following parameters are presented for each. discharge coefficient , analyzed, and where appropriate for the limiting discharge coefficient l (CD = 0.6 DEC14) ..with maximum safeguards assumed. ! Figures 15.6 15.6-17: . Quality, mass velocity and clad heat- q transfer coefficient for theihot spot and s clad burst locations. Tne heat transfer coefficient shown is calculated by the ICCBART code. i Figures 15.6 15.6-26: Core pressure, break flow and core pressure .q drop. The break flow is the sum of the flow rates from both ends of the guillotine i break. The core pressure drop is taken as the pressure just before the core inlet to the pressure just after the core outlet. l j j
~
Figures 15.6 15.6~35: Clad temperature, fluid temperature and core flow. The clad and fluid temperatures are given for the hot spot and burst j locations. Figures 15.6 '36 - 15.6-41: Downconer and core water level during reflood, and flooding rate. 4
)
I
.i 1
f .
} - /}
L m
'(
^
+ ,
71gures 15.6 15.6-47:- ECCS flow rates, for both accumulator and 1 t pumped safety injection. l s I j Figures 15.6 15.5-53: Containment pressure and core power
.. )
transients. , a Figures 15.6-54, 55: . Break energy release during blowdown and: the Containment wall condensing heat transfer coefficient for the worst break. ] l q The' maximum clad temperature calculated for a'large break.isE1983'F,-
- which is less than the Acceptance Criteria limit of 2200*F of .i n
V fl0CFR50.46. The marimum local metal-water reaction is 4.46 percent, { which.is well balow the embrittlement limit of 17 percent as' required by 20CFR50.46. The total core metal-water reaction is less than 0.3 percent for all breaks, . as compared with the 1 percent critai-ion of 10CFR50.46, and the clad temperature transiant is terminated at a time- ' when the core geometry is still. amenable to cooling. .As a result, the 1 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 j provided. 15.6.5.3.2 Small Break Results As'noted previously, the calculated peak clad temperature resulting
~
l from a small break IDCA is less than that calculated for a large braak. A range of small break analyses is presented which establishes
'the limiting break size. Table 15.6-4 lists important input parameters and initial conditions used in the small break analyses. .
0
-The core decay power and axial power distribution assumed for the small break analyses are shown in Figures 15.6-68A and 15.6-685. Note that the power. shape shown in Figure 15.6-688 is the most limiting for small 1 -/y 4
]
1 1
. break transients ratherLthan a chopped cosine power shape'with a peak- q at the center. This. power shape maximizes local power in the upper'
]
regions of the reactor core, with the peak occurring at the 10.0 foot l core elevation. This is limiting for the small break analysis because-- of the core uncovery proces.s for small breaks., .As the core uncovers,- I 3
#' the cladding in the upper elevations of the core heats up and'is sensitive to the local' power at that elevation. The cladding
~
a
)
temperatures in thetlower elevation of the core, below the'two-phase
~
mixture height, remain low. The peak clad' temperature occurs above 10 feet'for all small breaks. { d ECCS flow rate to the Reactor Coolant System as a function of the n4 . system pressure is used as part of the input. The ECCS was assumed to , j be delivering to the RCS 25 seconds after the generation of-a safety injection signal. This' includes the assumption of a lo second de3ay in
-the startup of.the diesel generator and a 2 second sensor response time" delay. _.
For these analyses, the ECCS delivery considers pumped injection flow i which is depicted in Figure 15.6-8 as a function'of RCS pressure. This figure represents injection flow from the centrifugal charging pumps (CCP) and safety injection '(SI) pumps. The 25 second delay includes time required for sensor response, diesel startup, and loading of the CCP and SI, pumps onto the emergency buses. The effect of flow from the RER, pumps is not considered here since their shutoff head is lower than RCS pressure during the portion of the transient considered-here. Also, minimum safeguards Emergency Core Cooling System capability and operability have been assumed in this analysis. 1 The hydraulic analyses.are performed with'the NOTRUMP code using 102%
.of the reactor. core design thermal power. The core thermal transient analyses are performed with the IOCTA-IV code using 102% of reactor core design thermal power. The results of these analyses are summarized in Tables 15.6-6 and 15.6-11.
l l' / - /[ j I i
1 i i
' l Figures 15.6-56 through 15.6-68B present the principal parameters of 1nterest for the small break ECCS analyses. These include:. .
.' J
-(1) The RCS pressure (Figures 15.6-56 through 15.6-58) .
(2) Core mixture level time' histories (Figures 15.6-59 through f j 15.6-61). i a 1 j
. r (3) Bot spot clad temperature histdry (Figures 15.6-62 through 15.6-64). ,
i
.For the limiting break analyzed (3-inch), the'following additional transient parameters are presented:
(1) Core steam flow rate (Figure 15.6-65) (2) Core heat' transfer coefficient (Figure 15.G-66) T (3) Hot spot fluid temperature (Figure 15.6-67) l The core power after reactor trip is given in Figure 15.6-68A, and the limiting small Break axial power distribution is given in Figure 15.6-688.- The maximum calculated peak' cladding temperature for the small breaks analyzed is 1925*F. The maximum local metal-water reaction is 4.59%. Further, the. total core metal-water reaction is less than 0.3% for all cases analyzed and the clad temperature transients turn around 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 l decay heat generated in the fuel for an extended period of time will be l Provided. These results are well below all Acceptance Criteria limits , of 10 CFR 50.46 and no case is limiting when compared to the results presented for large breaks. l I
~YO
/
/
is.s.s.4 conclusions - Thermal Annivsis L For breaks up to and including the double-ended severance of a. reactor ~ coolant pipe,- the F.CCS will meet the Acceptance criteria of 10 CFR 50.46.-~That is:- (1) The calculated fuel' element clad temperature provides a margin 8 ' to the requirement of 2200 F. (2) The. amount of' fuel element cladding that reacts chemisally with water or steam does.not exceed 1 percent of the total' amount of 31rcaloy in the reactor. (3) The clad temperature transient is terminated at a time when the core geometry is still amenable to cooling. The cladding oxidation limits of 17 percent are not exceeded during or after quenching. (4) 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. L /7 .
Rafarances for Section 15.6
- 1. Burnett, T. . W. T. , et al. ," I4FT".AN Code Description,"
.WCAP-7907-P.-A (freprietary), WCAP-7907-A-(Non-Proprietary), April 1984..
- 2. W. K. . Brunot, em - A Frwen for the calculation of Activity ggleases and Potantini Dosas from a Pressurized _, Water Reactor g nt, Pacific Cas and Electric Cannany (Octoner 1971 6
- 3. " Reactor Safety Study - An Assessment of Accident-Risk in U.S.
Commercial Nuclear Power Plants," WASH-1400, NUREG-75/014, October 1975.
- 4. " Acceptance Criteria fer Emergency Core Cooling System for Light Water Cooled Nuclear Power Reactors", 10 CFR 50.46 and Appendix K of 10 CFR 50. Federal Register, Volume 39, Number 3, January 4, 1974.
- 5. Bordelon, F. M., Massie, H. W. and Borden, T. A., " Westinghouse ICCS Evaluation Model-Summary", WCAP-8339, (Non-Proprietary), July.
1974.
- 6. Bordelon, F. M., et., al., "EATAN-VI Program: Comprehensive Space l Time Dependent Analysis of. Loss of Coolant", WCAP-8302,.
(Proprietary) June 1974, and WCAP-8306, (Non-Proprietary), June l 1974.
- 7. Kelly, R. D., et. al., "Calcu ated Model for Core Reflooding After l a Loss of Coolant Accident (WRETI4OD Code)", WCAP-8170 (Proprietary) and WCAP-8171 (Non-Proprietary), June 1974
- 8. Bordelon, F. M. and Murphy, E. T., " Containment Pressure Analysis Code (COCO)," WCAP-8327 (Proprietary) and WCAP-8326 (Non-Proprietary), June 1974.
- 9. Bordelon, F. M. , et. al. , "IACTA-IV Program: Loss of Coolant Transient Analysis", WCAP-8301, (Proprietary) and WCAP-8305, (Non-Proprietary), June 1974.
- 10. Bordelon, F. M., et.al., " Westinghouse ECCS Evaluation Model -
Supplementary Information," WCAP-8471-P-A, April, 1975 ; i (Proprietary) and WCAP-8472-A, April, 1975 (Non-Proprietary).
- 11. " Westinghouse ECCS Evaluation Model, 1981 Version," WCAP-9220-P-A, Rev. 1 (Proprietary) , WCAP-9221-A, Rev. 1 (Non-Proprietary) ,
February, 1982.
- 12. Letter from C. Eicheldinger.(Westinghouse) to D. B. Vassallo (NRC),
Imtter Number NS-CE-924 dated January 23, 1976.
- 13. Letter ULNRC-1207, dated November 15, 1985. .
L -f f
. i e
- 14. Lee, N.,.Rupprecht, S. D., Schwarz, W. R.,-Tauche, W. D.,
" Westinghouse'Small Break ECCS Evaluation Model Using the'NOTRUMP Code," WCAP-10054-P-A (Proprietary) and WCAP-10081-A (Non-Proprietary) August 1985. -15.'Salvatori, R., *W' westinghouse Emergency Core Cooling System - Plant Sensitivity Studies", WCAP-8340, (Proprietary) July 1974.
- 16. Young, M., et.:al., *BART-1A: A Computer Code for the Best Estimate Analyzed Reflood Transients", WCAP-9561-P-A, 1984 (Westinghouse Proprietary).
- 17. Chiou, J. S., et.al.,. "Models for PWR Reflood Calculations Using the BART code," WCAP-10062.
- 18. Meyer, P. E., "NOTRUMP, A Nodal Transient Small Break-and General Network Code", WCAP-10079-P-A (Proprietary) and WCAP-10080-A (Non-Proprietary), August 1985.
- 19. Kabadi, J. N. , et. al., "The 1981 version of the Westinghouse ECCS Evaluation Model Using the BASH Code," WCAP-10266, Revision 2, August 1986, (Westinghouse Proprietary).
- 20. Letter C. E. Rossi (NRC) to E. P. Rahe, Jr , (1), " Acceptance for Referencing of Licensing Topical Report WCAP-10266," November 1986.c
- 21. McIntyre, B. A.," ADDENDUM To BART-A1: A COMPUTER CODE FOR THE BEST ESTIMATE ANALYSIS OF REFICOD TRANSIENTS (SPECIAL REPORT:
MS-NRC-85-3025-A)," WCAP-9561-NP-A, Addendum 2 (Non-proprietary), Westinghouse Electric Corporation,' July 1985.
- 22. Young, N. Y., " ADDENDUM To: BART-A1: A COMPUTER CODE FOR THE BEST ESTIMATE ANALYSIS OF REFLOOD TRANSIENTS (SPECIAL REPORT: THIMBLE MODELING IN WESTINGHOUSE ECCS EVALUATION MODEL),: WCAP-9561-P-A, Addendum 3, Revision 1 (Proprietary) and WCAP-9561-NP-A, Addendum 3, Revision 1 (Non-proprietary), Westinghouse Electric Corporation, July 1986.
- 23. Information on the long-term build-up.of boric acid in the core .
region following a postulated I4CA transmitted from C. L. Caso, Westinghouse NES, to T. M. Novak, NRC, as enclosures to letter CLC-NS-309 (April 1, 1975).
- 24. _ Supplemental information on the long-term build-up of boric acid in the core region following a postulated IOCA transmitted from J. O.
Cermak, Westinghouse NES, to T. M. Novak, NRC, as letter JOC-NS-364 (July 23, 1975).
.25. Letter from D. B. Vassallo, NRC, to C. Eicheldinger, Westinghouse NES (May 30, 1975). .
I j
- 26. As umstions Used for Evaluatina the Petantial Radioloalcal '
consacruances of a Lsss of Coolant Accident for Pressurized Water j Reactors,- Regulatory Guide 1.4, Directorate of Regulatory + j Standards, U. S. Atomic Energy commission (June 1973) .
- 27. N. A. Styrikovich and O. I. Martynova, " Transfer of Iodine from- ,
Aqueous Solutions to saturated Vapor *, A*n=niva-Enaraiva, n , 4, (July 1964) pp. .'45-49. ].
,1
- 28. L. F. Parsly, Damian considerations of Rameter Containment serav Syst=== = Part TV. Calculation of Todina - Water Partition ,
Coefficients, ORNL TM-2412, January 1970. 29..D. E. Slade, Mateoreleev and Atomic Enarav, TID-24190, U. 5. Atomic Energy Commission (1972). ,
- 30. D. S. Duncan and A. 5. Speir, CRACE-TT - A PrGuisa and for Connutina Scharical ca=== = Rav Attenuation and Maatina in cylindrical Gammatrica,-Atomics International (1959).
'31. D. S. Duncan and A. B. Speir, GRACE-T - A Fiu-r-a Danioned for
'cmm,atina an=== = Rav Attenuation and Heatina in Reacter Shields, Atomics International (1959).
Analysis
- 32. ggNRC Standard review Plan for the Review of Safety Reports, NUREG-0800, Section 6.4, July 1981.
- 33. K. G. Murphy and K. M. Cnape, " Nuclear Power Plant control Room Ventilation System Design for Meeting General Design Criteria 19",
13th AEC Air Cleaning conference, August 1974.
- 34. W. K. Brunot, zmexam - A Fremrem for the Calculation of Activity Raiansas and Potential Desas from a Pressurized Water Raaeter Plant, Pacific Gas and Electric connany (October 1971).
I
I
.)
TABLE 15.6-4 ; Input Parameters Used in the LOCA Analyses Iaram Break Small Break ~ Parameter ! 3558 3558-Reactor Core Design Thermal Power * (Mwt) 13.744 14.482 Peak Linear Power (kw/ft) at'6.0 ft at 10.0 ft 2.50 2.50 Total Peaking Factor (Fg) Chopped See Figure Power Shape cosine, 15.6-68B F, = 1.543 17 X 17 17 X 17 Fuel Assembly Array
~J 900 900 lator Nominal Cold Water Leg Ay/
Volume (ft accumulator) lator 1350 1350 'l l Nominal TankCold Leg Acy/ accumulator) Volume.(ft 615 615 Minimum cold Leg Accumulator Gas Pressure (psia) ., i See Figures See Figure ) Pumped safety Injection Flow 15.6-45 15.6-8 thru 47 842 842 steam Generator Initial Pressure (psia) 3 1 11.5 11.5 Steam Generator Tube ! Plugging Level (4) Containment Parameters (See Table 15.6-8) 9237 9237 Initial Flow In Each Loop (1b/sec) 550.6 550.6 Vessel Inlet Temperature (*F) 618.S 618.8 Vessel Outlet Temperature ('F) Reactor Coolant Pressure (psia) 2280.00 2280.00 1
- Two percent is added to this power to account for calorimetric error.
Reactor coolant pump heat is not modeled in the IDCA analyses. l 1 N/ !
TABLI 15.6-5 (Sheet li Large Break LOCA Time Sequence of Events Minimum Safeguards C 0.4 C 0.6 C 0.8 p D=ECIA p D=ECIA p D=ECLG _ (sec) (see) -(seel 2133 0.0 0.0 0.0 Break Reactor Trip Signal 0.518 0.547 0.500 0.830 0. 6' O 0.590 SI-Signal. 14.20 11.60 Intact Loop Accumulator Injection 20.00 27.83 27.67 27.59 Pump Injection 36.48 28.047 24.042 End of Bypass 36.48 28.047 24.042 End of Blowdown 50.722 40.663 36.249 BOC Time Intact Loop Acc Empty 61.394 54.992 51.640 48.80 38.80 41.20 Hot Rod Burst Time Peak Clad Tamparature (PCT) Time 162.05 168.25 159.15 1
-1 TABLE 15.6-5 (Sheet 21 Iarge Break IDCA-Time Sequence ofiEvents ,
I Maximum Safeguards (
, ,Cp = 0.6 l DECLG (see) p: t 0.0
- Break 0.507
- Reactor Trip signal 0.670 8-I~ Signal 14.20
. Intact Loop Accumulator Injection' ,
27.67 Pump Injection End of B'ypass 28.047 28.047
- End of Blowdown 40.339 BOC Time Intact Loop Accumulator Empty 56.313 i 38.80 Hot Rod Durst Time Peak Clad Temperature (PCT) Time 163.61 9
0 4 L-23 ,
- - - - - -- _- _------ - _ _ - _ J
TABL2 15.6-6
- Sma11' Break LOCA Time Sequence of Events 3 in- ' 4 in 2 in fSac) M yl- #
(Sec) .,
\
~0.00 - 0. 00 - 0.00 Break-i 18.10 8.07 4.73 Reactor Trip Signal 28.08 15.85 12.11-Safety Injection Signal 53.36 41.43 37.73 Safety Injection Begins Loop Seal venting Begins 1259. 575. 339.
2185. 972. 643. .f Top of Core Uncovered ! (First Uncovery) N/A 895. Cold' Leg Accumulator Injection N/A N/A* 1705. 972. Peak Clad Temperature occurs (First Uncovery) 3682. 2627. 1436.. Top of core covered (First Uncovery) . 5088. N/A N/A Top of Core Uncovered (Second Uncovery) Peak Clad Temperature occurs 5477. N/A N/A
'(Second Uncovery)
None** N/A N/A Top of Core Covered (Second Uncovery) l
- The Peak. Clad Temperature for the 2-inch break occured during the ;
second uncovery. The highest clad average temperature during the I first uncovery was 8648F at 2587 s.
' ** j The 2-inch transient was terminated at 7703 s. with the core mixture level 0.7 ft. below the top of the core and j increasing. Two phase break flow bed stabalized with not RCS inventory increasing gradually, ensuring immenent core recovery. j Prior to this time, clad average temperatures for all elevations ha' i
peaked and were declining. I f l L-py
I ..u TABLE 15.6-7 (Sheet 11
.'Large Break LOCA Results Fuel Cladding Data Minimum Safeguards -
C 0.4 C 0.6 C 0.8 1 DD=E N p D=EdM DD =ECM RESULTS 1891.98- 1982.615 1879.213 Peak Clad Temperature (*F) 8.5 8.5 8.5 Peak Clad Temperature Incation (ft) 3.3211 4.4546 2.9933 Local Er/H O 2 Reaction (maximum %) 8.5 8.5 8.5 Incal.Er/N 20 Reaction Location for maximum reaction (ft)
<0.30 <0.30 <0.30 Total.Er/H 2O Reaction, (%)
6.00 6.00 6.00 l Hot Rod Burst Location, (ft) 3 I a b'hS
TABLE 15.6-7 (Sheet 21 j Large Break LOCA Results Fuel Cladding Data Maximum Safeguards C 0.6
'DD =ECIA i JtESULTS Peak Clad Temperature ('F) 1915.017 8.5
- Peak Clad Temperature Location (ft) 3.6693 ,
Local Zr/E 02 Reaction (maximum 4) 8.5 Incal Zr/H O2 Reaction, Location for maximum reaction (ft)
<0.30 Total Zr/H 2O Reaction, ($)
6.00 Hot Rod Burst Location, (ft) 9 i i l 1
)
i Yb _ _ - - _ - _ _ _ ~
I t a taste 15.bs Sheet 1 of 2 LARGE BREAK CONTA110ENT DATA
.I Est fres volume 2.145,000 it Initial senditions:
Pressure. 14.7 pais i Temperature 90*F - , EWST temperature 37'F , ~ Service water temperature 40*F Outs".4e tempr.rature 1*F ; Spray systems manber of posps operating 2 kament flow rate 4,000 sys actuation time 40 see Safe mards fan emelers , Eumber of faa seelers operating 8 I i Fastest post-sesident initiaties of fas eselers 29 see
. STRUCTURAL EAT SIl0C DATA Thickassa Area, (is.) Material se it 42.0 Cesarete 50,000 '
0.25 Steel
. 30.0 Concrete 27,000 0.25 Steel ,
See Table Eine Castias 6,000 } 6.2-11 42 0 concrete 5,000 . 0.25 Steel
. 15 0 Coacrate 81,000 24.0 . Concrete 17,000 0.25 Steel 72.0 Caecrete 14,000 42.0 Concrete 7,000 0.68 Steel
.me p 9 L -? 7 8 e
1 l 1 k TABLE 15 6-8 Sheet 2 of 2
. 1
~
d STRUCTURAL EAT SINK DATA Arsa. Thickasse Material se ft (in.) 1 i concrete 3.500 ! 36 96 S Steel
- 0.094 " Steel 46.000 0 055 Steel 48.000 0.077 .
Steel 10.000 0.054 Steel 20.000 0.145 Steel 12.000 0.5 Steel 14.000 20 Steel 3.500 1.0 Steel 6.500 Steel 29,000 0.33 , 0.312 Steel 20.000
# 4.500 0.75 Steel FAINTED STRUCWRAL MEAT SINK DATA Structural Seat Sink Structural Saat Siak .
Paint Thickness Surface Area Thickness (Mils) (is.) se it 6 42 25 50.000 30 25 6 27.700
- 15 6 6.000 42 25 6 5.000 42.68 6 7.000 44.96 6 3,500 0.5 6 14.000 20 6 3.500 1.0 6 ~
6.500 ' 6 29,000 O.25 0.75 6 4.500 I l 1 . . L-2 8
[' TABLE 15.6-9 REFIDOD MASS AND ENERGY RELEASES 1. Enerav Zina Maam 0.0 0.0
' 40.463 0.0245 31.597 41.238 0.3451' 447.757 41.938 52.3501 67944.45' 48.120 58.067 237.08 231857.22 71.361 274.53 235253.77 86.767 292.81 231452.69 103.467 305.33 225868.55 121.267 316.57 219522.91-165.567 401.22 223144.99 e
4
?
i
TABLE 15.6-10 BROKEN LOOP ACCUMULATOR MASS AND ENERGY RELEASES MA31 Enerav ZiER
'O.000 3293.645 .196498.848 1 010 3030.298 180787.581 2 010 2822.960 168417.816 3.010 2652.611 158254.777.
'4.010 2510.123- .149753.941 5 010 2388.073- 142472.411 6.010 2281.532 136116.184 7.010 2186.618 130453.620 8.010 2101.278 125362.236.
9.010 2024.062 110755.514 10.010 1953.892 116569.212 11.010 1889.841 117.747.940 12.010 1830.986 109236.647 13.010 1776.689 105997.244 14.010 1726.296 102990.841 15.010 1679.393 100192.578 16.010 1635.481 97572.784 17.010 1594.264 95113.809 18.010 1555.523 92802.509 19.010 1519.178 90634.164 20.010 1485.677 88635.479 21.010 .. 1454.784 86792.433 22.010 1425.936 85071.370 23.010 1398'.617 83441.510 24.010 1372.898 81907.101 25.010 1348.824 80470.864 26.010 1326.277 79125.704 27.010 1304.837 77846.584
)
)
i i D L
2 k TABLE 15.6-11 Small Break IDCA Results Fuel Cladding Data d 2 in i_in RESULTS- ,
~ Peak Clad Temperature (*F) 1059.201 1924,109 1401.156 12.0 12.0 Peak Clad Temperature Location (ft) 12.0 0.0460 4.5886 0.2150 Incal: Er/H O2 Reaction (maximum %)~
12.0 12.0 Incal Er/H 2O Reaction, Incation for 12.0 maximum reaction (ft)
<0.3 <0.3 . Total Er/H2O Reaction, (4) <0.3 Mot Rod Burst Time,.(sec) N/A N/A N/A Hot Rod Burst Location, (ft) N/A N/A 'N/A s
e 2 - 3/ L - - - - - - - - - - - - - - - - _ - - _ - - _
L IAEuU2a.I".ll coeause er m:1mai ammisen reasses most enansersas Trojas Teles fee teameh Toshateal Trojan 4alw for tecs aut.sts presawe aulvete - terrant ee , teatrien ess 61 ealw s.m.,
- Saaawac , 1.as stas for enemma 130 98 lastdo semperesere. *F sperassan M.? M.? a Genessammes reemeure. pose 1Amasass for amend eposertatu 99 Bessans humidtsy . 33 talasese W 48tre I 8 90 3m6$ad tuse&de temperasere. *F h d'r Ass 8 3. MS e 3e8 8 homemma ens inom 1.906187 e 38 emesasaamme telems is Sanseammmes amas Ammunna speemusa - 2 A11 speems 1 ames Sumere 30se F3aw smeas spetemedb sames ammemas fles 3r Os manaam. es amt sem souse. at $2
==
maarstes same. sus 8 4 all ensse M ase endare &$$ hanemma W emmLams omsar sampusesare. Y.* ==- 38 8 39 enarsans sema. ame
=== 3 e 30 t , = 6.5 e 10'I*I S.141 e 10 T ,= S.126 s as' 8
assa summent sene. ase/tr/ameh amis . panesse sees saaks: es be61s Samaansmums esmo 36.300 27.000 ases. f,a 38.3M '38.338 1h6ahmens, h as belas GS.Me emmaeammmes untle 38.300 Emed amen, f 8 48. 0 9 A1.4M , 8mereen he h es basis Other esent base asake 317.388 1M .617 tend ases, re 8 3 S.,3600 9 6007 emerege th6ehames. $s. ( as be&1s Insesed enumress sesessures 117.000 as6.6M Susan meus, ts8 M.131 34.339 emerees ehishamas en. Thesempaym4ena proporsaass S.92 1.8 As tasas t . See/trafs 8 *f ' as 31
. asett: 8.*F As tesis Ese ,)
M.S M es besla tg . Som/heafs2.*F 42 as husts M 9sette s y Epe ,) ,,w Ames tramofer eseffsssensas 8 8 8
- Sassial ammusensaas seassa.
asethr=fe4.*F 5 s aerejes Tesans e s Assojos 1 meant endefied Sm6ama had of 41sadman en8mm 9sep es Bahade 3mpamensial desef trat.h L- -*d damer
* ?remetsian se 1 mag =eems emmesses of 0.05 pass.61andsma pensa sash suussans of S.833 1 8 s Tasame Sundde lame =taan pess=61mmenn 3.3 s tuhade ,
tal 7 ,se she cantanamnus es,se seseerasure se *F.
- I O
/ e P 3,2
- t_
m j i
. i
_8* l
,g
.S ,j
.1 <
#1
== .{
l n in \ U
} $1 v
.u t; "
"U ew x1
, 8' % B ,
O b 5. 8 N lE*
'?
f2
- F.
E a a i i a i ,O i a e a a N N . e4 c4 e4 4 - o o o o o : (spuoenot.q) (YlSd) eanssoJd SCW L-33 4
i f TROJAN 11.5 PER,'i, IT TUBE PLUGGING ANALYSIS DOUBLE ENDED GUA h 'INE - COLD LEG - CD"88
- foe 2 ) Mst.82 i
1.6 l l E 1 .4 PEAK ELEVATION (X) - d or BURST ELEVATION (-) M_ 1.2 1 , . . , , E , 38 6 - 2e i 5 .6 r y M2 l wb [ ,1
- t. ~, . a l % 4 ' '
~
t g
.2 E , i a ;
'8 B. 50. 100. 150. 200. 250. 500.
TIME (SEC.) FIG.15.8 9 FLUlO QUAUTY - DECLG (CD = 0.8) i t
i J.
- t TROJAN 11.5 PERCENT TUBE PLUGGING ANALYSf3
' DOUBLE ENDED GUILLOTINE-COLD LEG-CD= 0.6 70 2.5 FDuet.82 1.8 g g C .
PEAK ELEVATION iX) " 3 E 1.4 BURST ELEVATION (*) M
!.2 %
1. e -I' g.. ,/ -.- g """
* ~'
*6 Hm==
pgy ' q 7 n
.P., - i i ,I .
,e
*, v q.
s$ L
.2 E
de z. se. iee. Ise.. aee. ase. see. TIME ISEC.) i FIG.15.s.10 FLUID QUALITY - DECLG (CD=0.8) P 35 _=
i I TROJAN -11.5 PERCENT TUBE PLUGGING ANALTS!$ DOUBLE ENDED GUILLOTINE - COLD LEG - CD = 0.6 70 2.5 Fewes.52 l.6 1 PEAK ELEVATION (X) 6 1.4 N BURST ELEVATION (*) W l.2 euw tasutA.evt 1 ,~.,. 3 .e 9
.6 4*f: 7
$g d p p guy x< y W ~,
- x-- .
/1 HL b# *
.2 E
a
' B 50. 100. 150. 200. 250. 500.
B. TIME (SEC.) FIG.15.610A FLUlO QUAUTY - DECLG (Co = 0.8) (MAXIMUM SAFETY INJECTION)
+
V
r 1 TROJ AN -11.5 PERCENT TUSE PLUGGING ANALY$15 DOUBLE ENDED GUILLOTINE - COLD LEG - CD = 0.4 [ 70*2.5 FONsi.82 I'.6
~
Z l +d PEAK ELEVATION (X) e as BURST ELEVATION (-) i 1.2 . i I. - - C.
.s
,f,,- l 5 i pyww E '6 . q I
$, , 7.hl y ,e
--1 u-y k
p y'u su ;3nrn
=
g - g
. .2 -
- s e
a I
' B 1. 50. 100. 150. 200. 250. - 500.
TIME ISEC.) l l I. FIG.15.6 11 FLUID QUALITY - DECLG (CD= 0.4) t b37 4
-- _ _. \
l .. 4 TROJAN.11.5 PEACENT TUSE PLUGGING ANALY$ts DOUBLE ENDED GUILLOTINE-COLO LEG-C D* 8 8 fos 2.5 FDMe1.52 '
- 58. l l u
- 40. PEAK ELEVATION (X)
(-) ,,,,, U g.,,. sonsT ELEVATION m d ae. 1** W yp.- , p
- e. .
-18. p w
*20. ,
u 8 a .se. _. W mm
*de. 2,
-50 15 . 50. 100. 150, 200. 250. 500.
TIME (SEC.) I l I { I FIG.1s.s.12 M ASS VELOCITY - DECt.G (CD=0.8) l z-37 6
le: 4 4 TROJAN -11.5 PERCENT TURE PLUGGING ANALYSl3 3 DOUBLE ENDED GUILLOTINE - COLD LEG - CD = 0.8 l ro.a.s ron.s.s2 I i. 1 t G l l D 40. . PEAK ELEVATION (X) 4 U suasi ELEVATION (-) , i g 50. e d 20. A 8 .. fc=<g . ay73 - c-g
-10. -
w 0 -20. b - g 1 a -se. 1 W u ' m m
-40.
2 50 1. 50. 100. 150. 200. 250. 500. ., TIME (SEC.) '! I i a FIG.15.s.13 MASS VELOCITY - DECLG (CD = 0.8)
/-39 ,
_ - - _ - _ _a
I d TROJAN -11.5 PEMCENT TUBE PLUGGING ANALYS15 DOUSLE ENDED GUILLOTINE-COLD LEG - CD= 0.8 70s2.5 FDNet.52
- se. ., ;
h 40. PEAK ELEVATION (X) U BURST ELEVATION (-) w s se. - - e d ae. l ' 19'
. e. .
%)Q La ~,_,.
-10. ,
U t1
-20. r y
8 - g sa. ,
- u i m -40.
m 15 . 50. 100, 150. 200. 250. 500. TIME (SEC.)
-i rio.1s.s.13 A Mass VELOCITY - DECLG (CD= 0.8)
(MAXIMUM SAFETY INJECTION)
i
.1 l
I TROJAN -11.5 PERCENT TUBE PLUGGING ANALY$t$ DOWSLE ENDED GUILLOTINE -COLD LEG-CD = 0.4 f j F082.5 FMest.62 l l
- 50. -
l l a o PEAK ELEVATION (X) h 48. , sunsT ELEVATION (-) I O w -se. H ' I s e - -l d 20. l
- {
3 Angg-._;+, ;
-i..
y . l 0 -20. " 1
.W l
> -50. )
$ 1 m
m
-48. -
1 E -se i, ,g, 3,,, 33,, ,,,, ,,,, 3,e, TIMC (SEC.) FIG.15.s.14 M ASS VELOCITY - DECLG (CD = 0.4) l l' 1 l 1
)
l [-
L E .
, l.]
1 g 'T w a , TROJAN 11.5 PERCENT TU8E PLUGGING ANALYST $ 1 DOWSLE ENDED GUILLOTINE-COLD LEG- CD
- 8 8 -
Feet.5 FDwet.62
.I l
w 55 3: E 45. U PEAK ELEVATION (X,1 .! - 1 g 4s. 3 sun 5T ELEVATION ( = l' [ f b
,se. /1 ;
/[
.)
.-- j s es. ! .
/ / -
. l 1
6 28 h'15 gg-eLI ka v _ 5-g, n - s ,,
!f g ~
aee. ase. . see. is. se. see. ise. TIME ISEC.) FlCL 15.618 HEAT TRANSFER C0 EFFICIENT- DECLG (CD= 0.8) 1 i f. c !
~Yk
.+
4 t e j TROJAM-11.5PERCENTTURE PLUGGING ANALYSIS cousLE ENDED GUILLOTINE - COLD LEG - CDe 0.6 po.a.s rom s.sa. . k t i f' pse. , g i 44s. tv
- 4g, _ , .' PEAR ELEVATION- (X) l.
)
- ss. ,
BURST ELEVATION (-) f , E / > ._ g ug*
- 3 r 3 as. ,
/
w 2,. u / A W Is is 1 gg-p ygy- . g s. / We s. 'se .- iee. ise. see. .ase. see. 1 l TIME tsEC.).- i FIG.15.614 MEAT TRANSFER COEFFICIENT- DECLG (Cg = 0.8) 1 L-Y3
I
. 4:
TROJAN -11.5 PERCENT TUSE PLUGGING AN ALYSIS DOUSLE ENDED GUILLOTINE - COLO LEG - C D* 8 8 res2.5 FDM*1.82 . g se. I as. hgg, PEAK ELEVATION (X) g sussi ELEVATION (-) ss. ,
) .
se.
/
e n. , .
// ,,
g gg n F , /
~ "
k ,s,. a N(,$m-.A g,~ L.x
,7-W p
- Lu $j
; s. =
Ee v. 'se. ies. ise. aee. ase. see. TIME ISEC.) l 1 J FIG.15.s.16A HEAT TRANSFER COEFFICIENT- DECLG (C D
- 8 8I (MAXIMUM SAFETY INJECTION) 2-Y7
- _ _ _ _ _ _ _ _ _ _ . _ _ . _ _ _ _ . _ _ _ _ . _ _ _ _ _ _ _ w
t
?
t TNOJ AN -11.5 PERCENT TUBE PLUCGfMG ANALY$l5 DOUBLE ENDED GUILLOTINE - COLD LEG - CD'88 F0a2.5 FDM*1.82 y 50. I ' 545. 7 U PEAK ELEVATION (X) suRST ELEVATION (-) j 55 n A f E as. a s e
'c h
- Y . f. , .. <
m 8 ,,' JL b9 . I #j@* u 15 -
^ * ' ' "# ' "
e 10-gq - 7 a s- w Ee 1s. Te. see. ise. see. ase. see. TIME ISEC.) O g e O I_ _ _ _ _ _ . _ _ _ _ _ _ _ _ _ _
TROJAN-11.5 PERCENT TUSE PLUGGING ANALYSIS DOWSLE ENDED GUILLOTINE - COLD LEG - CD = 0.8 - 70 2.5 FDuet.62 2.N ages i 1 WOW s * :
\
T
\ ,k . .
12 18 18 18 N 32 pa 26 9 3 8 0 0 18 TIGAE (SECOteOS) MGURE 15.618 CORE PRESSURE- DECLG (CD = 0.8)
~ !'1?o
TROJAN 11.5 PERCENT TUBE PLUGGING ANALYSIS DOUBLE ENDED GUILLOTINE - COLD LEG - CD = 0.8 F0e2.5 FONet.62 NOS 3005 I.i-
\
N
- l w
e 5 F5 le II 5 15 17 5 at 33 5 N. 37 5 58 8 25 l TItst (SECtHe00) 9 I MGURE 15.619 CORE PRESSURE - DECLG (C, = 0.8) 1 I o -- ._ - - - - - - - - - - - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . _ _ _ _ _
a k 1 E i
. TROJAN -11.5 PERCENT TUBE PLUGGING ANALYSIS DOUBLE ENDED GUILLOTINE - COLD LEG - CD s 0.4 l 70e2.5 Penet.82
'j I
1 1 3585 _ NB5 ISN ISOS N
\ w w
% as 8 5 18 15 N N SS 55 TMSE(SECOIIO4 l
l l t FIGURE 15.& 20 CORE PRESSURE - DECLG (C D =0.4) 2-Y7
l d
'\
f . TROJAN -11.5 PERCENT TUSE PLUGGING ANALYS15 DOUBLE ENDED GUILLOTINE = COLD LEG - CD a 0.0 feet.S FDuet.52 i 9055C E C.. X s 00gst 6
\ w k ,_.. k 400K
- 3
# \
g 5000C+6 SO90C+5 IIggeC+5
. N '
\[ \
8 3 4 8 9 18 13 14 le le 29 32 28 . 36 TIIAE (SEC0000SI FIG.1s.6 21 BRF.AK FLOW RATE - DECLG (CD = 0.8) A~Y?
i TROJAN -11.5PEMCENT TUBE PLUCGING ANALYS1$ DOUBLE ENDED GUILLOTINE - COLD LEG - CD= 0.8 Feet.5 Fenst.82 { i 7seec*l J esost.s l i 1 g .. N - m X 3 esse.s
\
I N x i .. N
. J D -
= Igest.t 75 18 la 5 15 17 5 as 33 5 36 37 8 Se 8 35 5 TIABE 18ECONOS) l 1
1 FIG.15.6 22 BREAK FLOW RATE- DECLG (CD= 0.8) l l l A-50
.J
-1 i
i TROJAN -11.5 PERCENT TUSE PLUGGING ANALYSIS DOWSLE ENDED GUILLOTINE - COLD LEG - CD = 0.4 l 7082.$ 1*L"st.82
,1 I
e
, 9899C*9 i gg ,,g.g g.c.. ;
'geset 6 -
I f 2005C+5 3 I .. x N # A I o
. isost s, , ,, ,, , ,, ,, ,, ,,
TIIAE (SSN I FIG.15.6 23 BREAK FLOW RATE - DECLG (CD=0.4) I' I d -5/
,t i
TROJAN-11.5 PERCENT TUBE PLucctNG ANALYSIS DOUBLE ENDED GUILLOTINE - COLD LEG - CD = 0.8 - g F0e2.5 FDMe1.82 t H g .. N y
- l. x n.. ;
I... 90
.esi , , , is le 18 8' "
' , is is Tiens(secossost FIGURE 15.5 24 CORE PRESSURE DROP - DECLGD(C *'
( I y-f2
c , L 1 l ; k a TROJAN-t1.5PERCENTTUSE PLUGGING ANALYSIS-DOUSLE ENDED GUILLOTINE - COLD LEG - CD s 0.8 . Post.5 PONat.82 f 1 l i
, 'i!
99 a l N ..- de N t _ : w 1 ' E.N 4,
*N 8 35 5 FS 18 13 5 15 IF 5 N N5 N 27 8 N l
liest (SECONOSI - ) 4 MOURE 15.6 25 CORE PRESSURE DROP - DECLG (C, = 0.8) j i
. l 1
2-53 i
l 1 L . .) 1
- TROJAN 11.5 PERCENT TUBE PLUGGING ANALYSIS COUSLE ENDED GUILLOTINE - COLD LEG - CD e 0.4 74*2.5 FDNst.82
)
N '
- r. .
N l-N I, y. ~ E=N, I . M
*N N N N 35 de 8 5 IS 15 TItst (SECONOS)
FIGURE 15.6 28 CORE PRESSURE DROP - DECLG (C D
- L-S't
l TMOJ Atl .11.5 PERCENT TUBE PLUGGING ANALYSIS= DOUBLI ENDED GUILLOTINE - COLD LEG - CD . 0.8 7 M ,5 FDH.1.62 i
- .25C*e4 , ,
6 PEAK ELEVATION N, m .225E eg (-) SUR5T ELEVATION b .2C*e4 g - & '
~N-
~ .175E ed
.isc.e, S/W M N B.i,st.e, 1// %
E .It.e4 f[ -
\ '
' M. >
kw 7se.
\ .
g see. e g 2se. . d a see. ase. see.
- 2. se. see. ise.
TINC ISEC.) FIG.15.6 27 PEAK CLAD TEMPERATURE- DECLG (CD= 0.8)
\
l i l 1 P55
- - .c
)
l TROJ AN -11.5 PERCENT TUgE PLUGGING ANALYSIS DOUBLE ENDED GUILLOTINE-COLD LEG-- CD= 0.6 F0e2.5 IDnet.82 i f
- .25C*04 , s PEAM ELEVATION (X)' -
sn ,gggg.ge SUR$T ELEVATION (-) W'.
. y
.2C*04 -,,
iv5c...
^
s"
.15C 04 125E 64
~\vf ,
-n 3 e .ne.es
\
N-w 750.
\ '
~ :
g See. 250. 8 150. 200. 250. 500. 15 50. 100. TIME.tSEC.) 1 i t l Fla.15.6 28 PEAK CLAD TEMPERATURE - DECLG (C D
- 8*8I l
l A*STc
s y 5-t TROJAN-11.5 PERCENT TUSE PLUGGING ANALYSf3 DOUBLE ENDED GUILLOTINE-COLO LEG-CD s 0.4 F0e2.5 FDNet.82 1. i
- .2SC*04 I i w
PEAR ELEVATION (X) _
$ .225E*Od BURST ELEVATION ( .- )
y ' B .2C 04 y _ .se- -ww _n.,- . _ _
,37,g,,,
c.., b ZK N 3.32st.., W/
.i
\ %
Y L E. .lE*04 N w 750.
\ .
9' see.
\ -
-g e
.g 250.
d e ise. see. ase. see.
- e. se. ies.
TIME (SEC'.) Me.15.s.2s A PEAK c(.Ao igurenArung - DecLa (cg = o.s) (MAXIMUM 5AFETY INJECTION) L -5 7
?
TROJ AN -t 1.5 PERCENT TUSE PLUCCING ANALY$15 DOUBLE ENDED GUILLOTINE- COLD LEG = CD = 0.4 Fest.S e t.82 i s I
- . 25E 94 i 1 w PEAR ELEVATION (X) ,
0 225E 94 BURST ELEVATION (-) y D .2E*84 '
.175C*84 8
15E 84 1W/ X N
,,,E.,,
vy -
. q '
E- .lE 04 . 750. N' - gli See.
- I g 25"8.
d e
- 50. 100 150. 200. 250. 500 D.
TIME ISEC.) FIG.15.6-29 PEAK CLAD TEMPERATURE - DECLG (CD = 0.4) A~58
1 i TROJAN .11.5 PERCENT TUSE PLUGGING ANALYSIS (' DouSLE ENDED GUILLOTINE-COLO LEG-CD
- 8 8 l
' Fes2.5 FDNet.82 r
.25E*84 u
$ .225E*84 ,
g PEAK ELEVATION (X) _
.2C, sunst ELEVATION' (-)
d.
.175E*84
% /l f,,& L t..
w 15E*Bd .
..%.,.-g .
.125E*84
> .lE*04 ,
m I 750
/ -
. ; y \. .-
; 50.. s y .
n .. . 8 150. 200. 250. 500. D. 50. 100 TIME ISEC.3 FIG.15.8 30 FtulD TEMPERATURE - DECLG (CD=0.8) L-57
'w j ,
%I k!: TROJAN 11.5PERCENTTUSE PLUGGING ANALYSIS
- i- DOUSLE ENDED GUILLOTINE - COLD LEG - CD e 0.8 it poet.5 paw 1.52
.25C 04 6
0 225E 84 PEAR ELEVATION (X) g .2C 84 sURST ELEVATION (-) g "
.175E 84 ily .15t es f t . 9Y a r.
.125E 84 ,
> .st e4 W w k 750. ^
g
~ 500
}' '
}A{
\ -
b _
\ _
x- ' E e,, ,,, 3 , , ,, 33,, 2,,, 2se. see. j TIME (SEC.) 1 Fla.1s.s. 31 Ftuto TEMPERATunt- oteto (ca. o.s) D 4 -6o
l 4 1 TROJ AN -11.5 PERCENT TUBE PLUGGING ANALYST $
. DOUBLE ENDED GUILLOTINE - COLD LEG - CD*88 i res2.5 FDMet.52 i
- .25E*04 6
$ .225E 04 '
PEAK ELEVAT!DN (X) -' gg ggg gy
.2E*04
.175E 04 '
iw .15E 04 Ae z-llAbit u
~
~
f u\ A' B .125E 04 . k, s l - N- , it.e4 k 750'
'vy
- See.
s N .
+
b -
\
250. ; x B. 50. 100. 150. 200. 250. 500. TIME ISEC.) 1 FIG.15.6 31A FLUID TEMPERATURE - DECt.G (CD=0.8) l
.. (MAXIMUM. SAFETY INdECTION) i 1
l 1 . 1-6/ \ __ ____-_ - _ A
l
-i l
TROJAN -11.5 PERCENT TURE PLUGGING ANALY$t$ DOWSLE ENDED GUILLOTINE - COLD LEG - CD = 0.4 ro.a.s row s.se l I
- :i
-, .25E*e4 c
- y.225C04 PE#X ELEVATION (X) _
$ .2C e4 sun 57 ELEVATION (-)
~ .175E 04 O " '" r" ,-A-b_ .!5E e4 *>
- $ .125E e4
.lg.e4
$ 7se. w m
- See.
@yr' N ,
" \ '
b 25e.
~
x E e see. ise. 2ee. ase. see.
- 2. se.
TIME ISEC.) t . j FLUID TEMPERATURE - DECLG (Cg = 0.4) FIG.15.6 32 l L -d2
i
'i.' TROJAN 11,.5 PERCENT TU8E PLUGGING ANALYSIS
- DOUBLE ENDED GUILLOTINE- COLD LEG - Cg = 0.8 F0e2.5 FDHet.e2 1
f- l l I I sees. 7 sees CORE TOP (* l' CORE BOTT001 -(-) j L - - - g ng g . b /> # 3 ,,,,, , T F' ee 4000 Qf
*M 8000 e a a e e is la 14 is se as . as as as fient(sacessee l
l I m FIG.15.8- 33 CORE FLOW -TOP AND SOTTOM - DECLG (CD= 0.8) Q - __ -_ __
r i e i TROJAN -11.5 PERCENT TURE PLUGGING ANALYSIS DOUBLE. ENDED GUILLOTINE- COLD LEG = CD s 0.8 F0 2.5 FDHs1.82 5809 , I f '" o CORE CORE BCTToll TOP (*I (-) A W my a l*088 oe
+4005 9008 8M , es s 7s is la s is ir s as as s as af 5 ss tiens (seconos)
MG.15.4 34 CCRE FLOW-TOP AND BOTTOM -DECLG Cg 4 = 0.8) t -- 1 -d Y J
f f IIIL 4 l TROJAN 11.5 PERCENT TUBE PLUGGING ANALYSIS ' DOUBLE ENDED GUILLOTINE - COLD LEG - CD = 0.4 F0:2.5 rowet.s2 1 i 9008 l
, If !
esse t CORE TOP (e) CORE 80TTOne (-) g ^ 4 g 3000 l w ~. p_- ': i
-( -
g 3e00 si
+4006 5098 35 SS 35 48 8 5 IG 15 39 TION (SECO8808)
FIG.15.6 35 CORE FLOW - TOP AND 80TTOM - DECLG (CD = 0.4)
p TROJAN'-11.5 PERCENT TUSE PLUGGING ANALYS15 DOUBLE ENDED GUILLOTINE- COLD LEG - CD= 0.8 F0 2.5 FDMet.32 I 22.
\A -
DOWNCOMER LEVEL 2.0 . 3 ' 18.
~
- 16. .
t auEac" 12 j 10. R , Y _ ....myj;gugpym,ymyff ff;4 j
% f, COLLAPSED LIQUID LEVEL 8*
6.
.,g .
7 100. 150. 200. 250. 500. D. 50. TIME (SEC) FIG.15.6 36 REFLOOD TRAhSIENT- DECLG (CO= 0.8) DOWNCOMER and CORE WATER LEVELS
~bb
.tf.
l l
. F:
l l TMOJAN-11.5 PERCENT TUBE PLUGGING ANALYSIS DOUBLE ENDED GUILLOTINE - COLD LEG - CO= 0.8 foe 2.5 FDMet.82 ! 22.
,,- w% --
DOWNCONER LEVEL y 1s.
- 16.
s-J
'14 m .
- 12. g gg V8 @ tEyst
, a e*r l s.
)fMNHa.%16:^coLLapsto :4, ,,'":;;., ,
Liouro LEVEL
- 6. 7 4
200. 250. 500. B. 50. 100. -150. > TIME ISEC) i . I I M G.15.6 37 REPLOOD TRANSIENT - DECLG (CD = 0.8) DOWNCOMER and CORE WATER LEVELS k A -4 7
q,- l
- T l
i { TROJAN-11.5 PERCENT TUBE PLUGGING ANALYS!$ DOUBLE ENDED GUILLOTINE - COLD LEG - CD = 0.8 foe 2.5 posest.62 22. Ww-
- 20. DowncouEn LEVEL 18.
16. U
- 14.
- 22. f*#
s ,,- . . i #gtuc8
. ...~i;,v;;;;;;;;;;,v"nwgvinti;&y.Lailk 9.
,f "f <
COLLAP5ED LIQUID LEVEL i i S- r
%.'; se. ies.- ise. 2ee. 2se. see.
TIME (SEC) e MO.15.6-37 A REFLOOO TRANS!ENT - DECLG (CO = 0.9) DOWNCOMER and CORE WATER LEVELS
'(MAXIMUM SAFETY INJECTION)
,{
W
- p. .. ,
.n h: .,
4 TROJ AN -11.5 PERCENT TU8E PLUGGING ANALYSIS ' DOWSLE ENDED GUILLOTINE - COLO LEG - CD'0d' po.a.s row.t.sz -l t 22.
- 20. 7 DowncouER LEVEL l
E!!,. 18.
. . ... \
!6. ,
5 1' 2# tt"5' A y 12. W Maut # ,, ,5 y-yy . J 10- . s ..
- 8. COLLAPSED LIQUID. LEVEL ,
1 \
, 6. J.
g 50$ 100. 150. 200. 250. 500.- 13 . TIME ISEC) i e FIG.15.6 38 REFLOOD TRANSIENT- DECLG (CD=0.4) DOWNCOMEM and CORE WATER LEVELS 2-GP
l l n t TROJAN .11.5 PERCENT TURE PLUGGING ANALYSIS DOUBLE ENDED GUILLOTINE - COLD LEG - CO*88 F0 2.5 F0Het.82 l l l
- 10. 1 i
1 9. S. ' i G 7. 1 y . i s
- g. .z 6. -
>< m.o .
~
- 5. .
z.,_ 4 m.- 4.
,z 5. -
2. 1. 0 50. 75. 100. 125. 150. 175. 200. 225. 250. D. 25. . TIME (SEC) 1 l
- 1 1
1 FIGURE 15.6 39 REFLOOD TRANSIENT- DECLG (C U =0.8) CORE INLET VELOCITY
& 70
~ . .h (q
...h f ;
jv:
. 9 TROJAM-11' .5PERCENTTUEZPLUGGING ANALYSIS DOUBLE ENDED GUILLOTINE - COLD LEG - CD = 0.6 >
f eet.S FDne t.s2 ' 4
- 10.
4
- 9. -
- 8.
U 7.
\% -
s E b. z
- 5. .
i
2 4. . 3
. 5.
2. 1. 15 . 25. 50. 75. 100. 125. 150. 175. 200. 225. 250. TIME (SEC) MGURE 15.3 40 REFLOOD TRANSIENT- DECLG (C D =0.8) CORE INLET VELOCTfY 4 l 4 7/
/
1l - -.
, i) _' . . . )^
, j;
; i-6 l
TROJ AN -11.5 PERCENT TU8E PLUGGING AN ALYSIS pousLE ENDED GUILLOTlHE - COLD LEG - CD = 0.6 , 1 4 no.a.s ro i.s 1
- .a .
10. 4 cg , s 9. u 7. d
':?c 6.
z 5. _
.2 4.
to 5. 2. 1. E. 25. 50. 75. 100. 125. 150. 175. 200. 225. 250. TIME (SEC) s FIGURE 15.6 40A REFLOOD TRANSIENT - DECLG (C D *e*8I CORE INLET VELOCITY (MAXIMUM SAFETY INdECTION)
& 72 1
_ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ __-___-_________D
i v.g . L /> TROJAN -11.5 PEMCENT TUSE PLUGGING ANALYS15 ' DOUBLE ENDED GUILLOTINE - COLD LEG - CD = 0.4 Feet.S FImes.s2
- 18. ,
8I . 9. G 7. W j s.
. 55-Z 4. - a !
i l* 5. : l' e
- 2. -
- i
- 1. .
)
O D. 25. 50. 75. 100. 12.5. 150. 175. 200. 225. 250. TTHE ISEC) i i 1
)
1 REFLOOD TRANSIENT- DECLG (C D ** FIGURE 15.6 41 CORE INLET VELOctTY ,
.n W
TM0JAN.11.5 PERCENT TUSE PLUGGING ANALYSl3 DOUBLE ENDED GUILLOTINE - COLD LEG - CD e 0.8 F0e2.5 F0eest.52~ 7958
\
# /
/
I8055 E e l
. E "" . j
. /
y N 0 2 8 0 8 19 12 18 16 18 N 22 24 Tleft ISECOtsOS) 4 l 1 FIG.15.6 42 ACCUMULATOR FLOW (SLOWDOWN) - DECLG (CD = 0.8) l l 4-7y . u-- i
1 1 l vnoJAN 11.5 PERCENT TU8E PLUGGING ANALY$l$ . pousLE ENDED GUILLOTINE - COLO LEG - CD = 0.6 ( ro.a.s r m.i.5 l l l t rose h
- / .,
Isees E
, esse g '
t ( E= less . e e 25 5 75 le l2 5 15 IF 5 as 22 5 25 27 5 Se me essconosi me.1s.s.43 ACCUMULATOR FLOW (SLOWDOWN)-DECLG (CD = 0.8) L -75
i.
- +
l-l TMOJAN -11.5 PERCENT TURE PLUGGING ANALYSIS ; DOUBLE ENDED GUILLOTINE - COLD LEG - Cg = 0.4 Fes2.5 FONet.82
, \
W j
' s SSSS E4908 I
s e-8 IPSOS1000 S 5 5 If 15 N N N 55 de TM85 (SSC0060s) i l FIG.15.s-44 ACCUMULATOR FLOW (SLOWDOWN)- DECLG (CD=0.4)
TROJAN -11.5 PERCENT TUSE PLUGGING ANALYS15 DOUBLE ENDED GUILLOTINE-COLO LIG-CD= 0.8 - F0*2.5 FDNe t.52 1 ( 10." M 2 S
- s. -
l I D S s. 8 c' ' g i 4. U ' S 2 - O .O . 50. 100. 150. 200. 250. TIME (SECONDS) i Mount 15.s-45 PUMPED ECCS FLOW (REFLOOD)- DECLG (C D =0.8) , j l
/-??
( ( l. TROJAN -11.5 PERCENT TUBE PLUGGING ANALYSf3 DOUBLE ENDED GUILLOTINE - COLD LEG - Cg
- 0.6 90 2.5 FDNet.62 16.'
~
u M D s.
.S 6.
E d a s. . u W . c w 2. 8 .' ' 50' 100- 150. 200, ggg, TIME iSECONDS) 1 I FIGURE 15.6 44 PUMPED ECCS FLOW (REFLOOD)- DECLG (CD"'I i l L - 79' ;
1 i f TROJAN t1.5PERCENTTUSE plt [GGING ANALYST $ DOUBLE ENDED GUILLOTINE - COLD LEG - CD= 0.8 Feat.5 FDust.62 20.
~
u ' r M 17.5-D
': 15.
8
~ 12.5; 8 10.. -
g , 5 7.5 d : C 5.- . g 2.5- .
- 68. 50. 100. 150. 200. 250.
TIME (SECON05)
. FIGURE 15.8 46 A PUMPED ECCS FLOW (REFLOOD)- DECLG (CO
(MAXIMUM SAFETY INJECTION) f l l 2-77
I' l . 2 TROJ AN -11.5 PERCENT TUBE PLUGGING ANALYS!$ DOWSLE ENDED GUILLOTINE - COLD LEG - CD e 0.4 F0e2.5 FDuet.82 i le.-
~
o M 2 s. o S 6.
-a d
$4
.y o
2 8 .e . se. ies. ise. see. 2se. q TIME (SECONDS) ( i MGURE 15.6 47 PUMPED ECCS FLOW (REFLOOD)- DECLG D (C **' 4 { l j
i TM0JAN 11.5 PERCENTTUBE PLUGGING ANALYSIS l DOUBLE ENDEt, GUILLOTINF. - COLD LEG - CO = 0.6 F0=2.5 FDNet.52 58." _G 49.< g... t i E2.. c s
$_i.. :
5 E 150. 200. 250
e '8. 58. 100.
TIME ISECONOS) FIG.15.6 44 CONTAINMENT PRESSURE- DECLG (CD = 0.8) i 2 -S /
t 7 TROJAN-11.5PERCENTTUBE PLUGGING ANALYSIS DOUBLE ENDED GUILLOTINE - COLD LEG - Cg = 0.5 FO*2.5 FDMal.82
- 58. '
G 48. E b g...
- a. 2e. '
E .
- g19. .
5 5 200. 250.
" e '8. 58, 100. 150 TIME (SECONDS) l 6
FIG.15.5 49 CONTAINMENT PRESSURE - DECLG (C D = 0.8) L-$2 L_- __ __ _ __ _ _ _ __ _
TROJAN -11.5 PERCENT TUBE PLUGGING ANALYSl$^ -{, DOUBLE ENDED GUILLOTINE - COLO LEG - CD = 0.5 j F0 2.5 FONet.82 i i 58.< G
~
- 48. f e
W 58. R m Y28. . 5 .
! 18. . .
s E o 8.B . 58. 100. 150. 200. .250.
' TIME ISECONDS)'
FIG.15.4 49A CONTAINMENT PRESSURE - DECLG (CD*08) (MAXIMUM SAFETY INdECTION) I 4-13 ________-__-_-a
i 'yy
. TROJAN -11.5 PERCENT TU8E PLUGGING ANALYS!$ '
DOWSLE ENDED GUILLOTINE - COLD LEG- CD'84 79 2.5 FDNet.82 E { 58.- G 48. E_ 58. m W a. . .;.
~
z ( 18.-
- 1 5
j E. 3 200. 258.
" 8 8.
- 58. 100. 158. t TINE ISECONOS) 4 I
4
, FIG.15.8 50 CONTAINMENT PRESSURE - DECLG (CD=0.4) l I
L - sy
-------------a
i;' - d I a TROJ AN -11.5 PERCENT TURE PLUGGtNG ANALYSIS DOUBLE ENDED GUILLOTINE-COLD LEG- CD = 0.8 4 FO=2.5 FDNe t.52 ] y
., I t .,
1.E l
- 1. l r ,,
.e r
.4
- )
..t I
s a 4 e s le 12 14 tw4tsacowos) le ie se 22 as as
,x 4 FIG.1s.s.si core POWER TRANSIENT (CD=0.8) { l 1
< 2-%s
f t TROJ AN -11.5 PERCENT TUSE PLUGGING ANALYSIS DOUBLE ENDED GulLLOTINE - COLD LEG - CD s 0.5 Feat.5 FDNet.62 l 1.3 1 8 s I 2 .. B 25 5 75 le 12 5 Il 17 5 29 22 5 25 27 5 58 TIME (SEcotsOS) 1 l I I l FIG.15.5-52 CORE POWER TRANSIENT (CD = 0.8)
!'$b l
( .k tm
' TROJ AN -11.5 PERCENT Tutt PLUGGING ANALY$l$
' DOWSLE ENDED GUILLOTINE - COLO LEG - CD ". 0.4 - 70e2.5 FONef,62 a i F I.E I. l .E 8
.4
.3
. N de 8 5 IS . 15 N -N 50
- Tibet (BECONosp MG.15.8 53 CORE PCWER TRANSIENT (CD = 0.4) f p.-
Z47
i I 4 e 5 TROJAN -11.5 PERCENT TUBE PLUGGING AN ALYS!$ ' DOUBLE ENDED GUILLOTINE - COLD LEG - CD = 0.8 F0 2.5 FDNet.s2 i asset e
,. EsSet*O t..
\
G 5 asset e k
. .. N ,
y teset e E teest*9
\
3 N
'a m seest.7 V%
*"*# a? 5 se g as s 7s le la s is 17 s as 22 s as Time (seconos)
FIG.15.8- 54 BREAK ENERGY RELEASED TO CONTAINMENT l 1 _ - - - _ - -- ---- I
% 'a
(
,r TROJAN-11.5PERCENTTUSE PLdGGING ANALYSIS DOUBLE ENDED GulLLOTINE == COLD LIG - CD e 0.6 F0*2.5 FDNet.82 l
Y .32E*04-U . u h .lE*04- .i 5 E see. b W u 600. E l tee. E g 200. W
- 50. 100. 150. 200. 250.
e .9 . TINE ISECONOS) i I l FIG.15.6 55 CONTAINMENT WALL CONDENSING HEAT TRANSFER COEFFICIENT )
~
r
t TROJAN - 11.5 PERCENT TUBE PLUGGING ANALYSIS COLD LEG SMALL BREAX loCA (2 IN.) Fg=2.50 FDH=1.62 2488. , i 2299." l 2999.<
~
e em i E sees.. ! M - 1898.< l u . i M
~~
gi4.... .N : e E k_ _- , 1298. " 1998. "
- ~-
I
*g . less. 20s0. 50ss. dese. 50se. Sess. 7000. 8000.
TIE (SEC) Figure 15.6-56 RCS Depressurization Transient (2 in ) l
~
b
45 : TROJAN - 11.5 PERCENT TUBE PLUGGING ANALYSIS COLO 1.EG SMALL BREAK IDCA (3 IN . ) .' Fg=2.50 FDH=1.62 .] i 0 2488. 1
.)
2288." 2888. < - g 1888.< _t g 1888. - g . , m h 1488. " : !
&_1288. ' - at !
3 m b1888. " 088. " i 888. " , 5888. 2588. 488 8. 588. 1888. 1588. 2888. TIME (SEC) Figure 15.6-57 RCS Depressurization Transient (3 in.) i j -7/
'd.
TROJAN - 11.5 PERCENT TUBE PLUGGING ANALYSIS COLD LEG SNALL BREAK IOCA (4 IN.) Fg=2.50 FDH=1.62 2488. 2288.' 2988." 1998.< E E1888.<
~ .
~
.. 1498."
r ~ g 1288. " l 1888." 888.. . 888. " m 488. " ESS.8. 488. 888. 1288. 1688. 2888. TIE (SEC) .
. Figure 15.6-58 RCS Depressurization Transient (4 in.)
1 - 7.2 i
'V i
1 i
- TROJAN - 11.s PERCENT TUBE PWGGING ANALYSIS-COLD LEG SMALL BREAK MCA (2 IN.)
re-2.so rou-1. 2
- 54. -
E'a . ' ss." , , C
- as.- -
Y : u k\ w as.~ , i g l x E y 24.< ! B TOP OF CORE w# WW
.. N '
sees, 1s.s. sees. ases. sees. 4ees. sees. sees. 7ees q
,' TInc (sces n , 1s..-s. cor. x.txem. n.ies c2 u.3 j .
/ 13
is t TROJAN - 11.5 PERCENT TUBE PLUGGING ANALYSIS ~
. cold LEG sMALL BREAK MCA (3 IN. )
, , Fg=2.50 FDH=1;42 ;
n ; J a
, j 54. q 52.'
58." {28." 28.' :- 24.'
.u
= -
TOP OF CORE } g
- 22. " l
- 28. "
- 18. '
1908.. 1589. 2988. 2500. 5000. 16 3. '580. TIME (SEC) Figure 15.6-40 Core Mixture Height (3 in.) A ~ ff.
i 4; I i i j TROJAN - 11.5 PERCENT TUBE PIEGGING ANALYSIS ,j
. COLD LEG SMALL BREAK IDCA- (4 IN.)
rg=2.so 70s-1.62 - 1 54 1 ! 52." 58." {28." h .
]
y 26." 24." _M
> W TOP OF CORE ,
'*- - - {
$22.-
FM l
- 28. "
- 18. "
m 18 8. 200. 40'O. 600. 800. 1000.1280.14s3.16se.1880. 2000. TIME' (SEC) Figure 15.6-61 Core Mixture Height (4 in.) d *ff _ . _ - - _ _-___________-__a
'S
. TROJAN SF1ALL BREAK FDH=1.62-INCH DI At1ETER COLD LEG BREAK I' PEAK.12.00 FT
' CLAD AVG.TEf1P. HOT ROD ,
P seas. C ,
. asas. l lW 1
- asse.
I i s. ! _R g ses.
~
vw, _
/- 1 6.............TIM IEC8 t
Figure 15.6-62 Clad Temporgture Transient (2 in.) h L----- ___
i I i 1 l 4 t
]
TROJAN SMALL BREAK FDH=1.62 - 3-INCH DIA!1ETER COLD LEG BREAK
' PEAK.12.00 FT ,
CLAD AVG. TEMP.HDT RDD . 1 N sees. I i
~~
a.
, 2530.
= *
~ 3050. I
/ ,
1985. 1 i .
/
n W ING.
/
R-
= ;
) -
}see.
l sees. 2005. 230s. 3400. 3600. 2000. ecc. sees. 130s. 34es. 1600. Titt teKC3 , l rigure 15.6-s3, clad Temperature Transient (3 in.) F A-97
_$l' t-t< :... t t t , , v E t. l- . TROJAN SMALL BREAK FDH=1'.62 - y F 4-INCH DIAMETER COLD LEG BREAK . -
- PEAK.12.00 FT i CLAD ~ AVG. TEMP. HOT ROD' o
8000. . 4 l~ C
,3EOS.
s ,. ), ,
- asas.
I 1588. h r" j. . 7
/ \ ~
I
}
i580. 8M . 488. 588. 088. 788. 888. w Iges. Itse. 1200. 1988. 140s. Itse. 1888. l. ' ' Tite: IECI . J.' Figure 15.5-64 Clad Temperature Transient (4 in.) 6
r- ,e i 4 Tao:rAN - 11.5 PERCENT TUBE PIRGGING ANALYSIS ! coLo LEG SMALL BREAK IDCA (3 IN.) Fg=2.50 FDH=1.62
~488.
558." 588. ' s .1 1-Ta 258." i
=
1 288." r , b158. w M i ' V' ; 188. " . p . a .... _
- 58. " ,
l l t 58'8. 18se. 1588. 2988. 2588. 5808.
'8 .
TIME -(SEC) Figure 15.6-65 Core steam Flow Rate (3 in.) ,
'h
g'_ l- . q...,.- . , a .
)
TROJAN SMALL BREAK FDH=1.62 .
- I
~
S-INCH DIAMETER COLD LEG BREAK PEAK.12.00 FT HEAT TRANS.COEF. is8 - m .- a ._. e
.isa l
- ~
g- s e N -- 3.i J j see. ses.. asse. mes. m. aass. asse. Ib. asse. sass. sees. 73m istcs t Figure l5.6-66 Core Heat Transfer Coefficient
- (Rod Film Coefficient) (3 in.)
E
e.
}
~
t
~
j 1 TROJAN SMALL' BREAK' FDH=1.62 3-INCH DIAMETER CDLD LEG BREAK PEAK.12.00 FT FLUID TEMPERATURE E800. C 2588. 2008.
/-% \
7 SEE. i g i=0.
' (
\
E ,
'N % -
p M. 1000. 2000. 22e8. 2400. 2688. 2000. 500. 1888. 120s. 1400. 1884. TIPC IEEC3
. Figure 15.6-67 Bot Spot Fluid Temperature (3 in.) . = ,
L -/0 /
i i o a
~ .
m
~
N _= . y _
- p u
- o
$ i E - g :
. g
~ u M E k "o @ 3
$ =_ -
m w 4
" N i 4 - e :
_ ~ 0 h
= -
W. . .-
- . g g .
- oc e- " 4
- C e
w
-- o :
C Y
- p -
- e r w 1 g '
- t
_ *o-IIIll I I I I Ill I I I lIIII I -I I I To
- Ta 7o o,- .
*a - -
H3M0d 1Y 11VM/S11YM L -/02 !
)
l i e e r e N e l 9 e 4 e '
- . u M
W i
- G k
e Pm
'.I s ..
i
.a i
on 4 B
=
O,
" O M g
\\n
-- lif .
S E. Y., s$n . .
-n 5 S m% . i.
8 - u
- n 8, b
1
.D W i m
- g 8 8 l g 3
~g a 8 6 4 4 l
W N 9 9 $ H
- d d
^
a a . - - - (z)ad A -/0 3 .
.- v-w.-- -,q Ya_ _
k* $ $
~ -
g4
'! ) ?
- l. s; e
, 4 2
2 + e e T 6 . 0. REFERENCES 4 8 i > \ Sa s s f, i i '+
...j A.
q k g . o t 4 s J f i i i 4 m__-. .._
a 4 1
6.0 REFERENCES
- 1. Davidson, S. L. and Kramer, W.R.; (Ed.) " Reference Core Report VANTAGE 5 Fuel Assembly," WCAP-10444-P-A,' September 1985.. .
- 2. Davidson, S. L. (Ed.) et al., " Extended Burnup Evaluation of Westinghouse Fuel," WCAP-10125-P-A, December 1985.
'3. Davidson, S. L. (Ed.), et. al., " Westinghouse Reload Safety. Evaluation-
. Methodology," WCAP-9272-P-A, July 1985.
- 4. Camden, T. M.,- et al., "PALADON-Westinghouse Nodal Computer Code,"
WCAP-9485-P-A, December 1979 and Supplement 1, September 1981.
- 5. Davidson, S. L. (Ed.), et al., "ANC: Westinghouse Advanced Nodal Computer Code," WCAP-10965-P-A, September 1986.
- 6. Chelemer, H. , Boman, L. H. , Sharp, D. R. , " Improved Thermal- Design Procedure," WCAP-8567, July 1975.
- 7. Motley, F. E., et al., "New Westinghouse Octrelation WRB-1 for Predicting Critical Heat Flux in Rod Bundles'with Mixing Vane Grids," WCAP-8762-P-A and WCAP-8763-A, July 1984.
- 8. Skaritka, J., (Ed.), " Fuel Rod Bow Evaluation," WCAP-8691, Revision 1 (Proprietary), July 1979.
- 9. " Partial Reston;* to Request Number 1 for Additonal Infoi ation on WCAP-8691, Revision l" letter, E. P. Rahe, Jr. (Westinghouse) to J. R.
Mi'ler (NRC). NJ-EPR-251:. dated October 9, 1981; " Remaining Response to Request Number 1 T..- Additional Information on WCAP-8691, Revision 1" 1stter, E. P. Rahe, Jr. ' Westinghouse) to J. R. Miller (NRC), NS-EPR-2572, dated March 16, 1982. l l 6-1 MesL6-S?1023
- y. ,- - -
I 9
'1 l
10.- Letter C. Berlinger (NRC) to E. P. Rahe, Jr. (Westinghouse) " Request for Reduction in Fuel Assembly Burnup Limit.for Calculation of Maximum Rod Bow I
. Penalty," June 18, 1986.
- 11. . Leech, W. J., et.. al., " Revised PAD Code Thermal Safety .Model," WCAP-8720, Addenda 2 (Proprietary), October, .1982. :
- 12. S. L. Ellenberger, et. al., " Design Bases for the Thermal Overpower Delta j T and Thermal Overtemperature Delta T. Trip Functions," WCAP-8745-P-A f (Proprietary) arid WCAP-8746-A. (Non-Proprietary), September 1986.
- 13. " Acceptance Criteria for. Emergency Core Cooling System for Light Water Cooled Nuclear Power Reactors," 10CFR50.46 and Appendix K of 10CFR50.. .
' " Federal Register, Volume 39, Number 3, January 4, 1974.
- 14. Kabadi, J. N. et. al., "The 1981 Version of the Westinghouse ECCS Evaluation Model Using the BASH Code," WCAP-10266-P-A, Revision 2, March 1987, (Westinghouse Proprietary).
- 15. Bordelon, F. M., et. al., " SATAN-VI Program: Comprehensive Space Time Dependent Analysis ~of Loss of Coolant," WCAP-8302, (Proprietary) June 1974, and WCAP-8306, (Non-Proprietary), June 1974.
- 16. Kelly, R. D., et. al., " Calculated Model for Core Reflooding After a Loss of Coolant Accident (WREFLOOD Code)," WCAP-8170 (Proprietary) and i WCAP-8171 (Non-Proprietary), June 1974.
l
- 17. Bordelon, F. M. and Murphy, E. T., " Containment Pressure Analysis Code (C0CO),' WCAP-8327 (Proprietary) and WCAP-8326 (Non-Proprietary, June 1974.
- 18. Bordelon, F. M., et. al., "LOCTA-IV Program: Loss of Coolant Transient Analysis," WCAP-8301, (Proprietary) and WCAP-8305, (Non-Proprietary), "
l- June 1974. l 6-2 l ) i mm mu
Y f
- 19. Young,- M. , et. al . , "BART-1A: A Computer Code for the Best Estimate
' Analysis'of Reflood Transients," WCAP-9561-P-A, 1984 (Westinghouse Proprietary). .
- 20.- Lee, N. , Rupprecht, S. D., Schwarz, W. R., Tauche, W. D. , " Westinghouse Small Break ECCS Evaluation Model Using the NOTRUMP Code," WCAP-10054-P-A (Proprietary) and WCAP-10081-A (Non-Proprietary) August 1985.
2'.1 Meyer, P. E.,
"NOTRUMP, A Nodal Tran2 .ent Small Break and General Network Code," WCAP-10079-P-A (Proprietary) and WCAP-10080-A (Non-Proprietary),
August 1985 ________m._ _ _ 1
APPENDIX 4 RADIOLOGICAL CONSEQUENCES I: - -- - - - - - . - . - _
L l APPENDIX ] i IMPACT OF EXTENDED FUEL BURNUP ON RADIOLOGICAL CONSE0VENCES OF ACCIDENTS The Trojan upgrade will be capable ~of achieving a lead rod average burnups up to the maximum lead rod burnup permitted by the approved Westinghouse Topical, WCAP-10125-P-A.
'As presently discussed in Section 15.7.4.2.1.1 of the Trojan FSAR, an evaluation was performed to determine the effects on the radiological consequences of the' fuel' handling accident of increasing the discharge assembly average burnup from 25,000 MWD /MTU to 44,000 MWD /MTU. The only change identified was that due to the differences in the fuel rod radionuclides
. inventories and the calculated. increases in the radiological consequences were approximately three percent for both the whole body dose and the thyroid dose. As the FSAR states, this increase is considered to be negligible, given the overall accuracy of the calculation.
While fission product inventories a're roughly proportional to operating power level, the level of fuel burnup has little impact except for isotopes with long half-lives. This is supported by the Westinghouse topical report, WCAP-10125-P-A (Proprietary), titled " Extended Burnup Evaluation of Westinghouse Fuel," which demonstrates that extension of fuel burnup from 33,000 MWD /MTV to much higher discharge region average burnups evaluated in the topical report would have ;nly a small impact on the core fission product inventories. This would not significantly affect the radiological consequences of postulated accidents. For the fuel handling accident, extending the discharge regiors average burnup from 33,000 MWD /MTV to the maximum value evaluated in WCAP-10125-P-A would result in an increase of approximately four percent in the thyroid dose. This increase is based on continued use of the fuel handling accident analysis assumption, defined in Regulatory Guide 1.25, that ten percent of the core inventory of short lived isotopes and thirty percent of the core 1 s033H A ssC22 i _ _ _ _ _ _ _ _ _ _ I
1 i l inventory of long lived isotopes are in the fuel red gap. The short-lived isotopes are.of greatest concern in regard to radiological consequences of the accident, and analysis shows that the fraction of short-lived isotopes in the fuel rod gap, when at its maximum, would be bout one percent or less of the core inventory. Extending fuel burnup.would actually result in a reduction in the gap inventories of short-lived nuclides due to operation at lower fuel temperatures and tnus a lower rate of diffusion of fission products through the fuel pellets into the gap. The radiological consequences of accidents other than the fuel handling i accident are also impacted to a slight degree. As discussed in WCAP-10125-P-A, the impact of extended fuel burnup on the radiological
]
consequences of a rod ejection accident would be to slightly increase the. ] thyroid dese (about two cercent) and to decrease the whole body dose. .This same affect on radiological consequences would also be seen in other accidents involving release of reactor coolant activity whether or not there is any fuel damage as a result of the accident (e.g., Steam Generator Tube Rupture,
' Loss-of-Coolant Accident, etc.). These increases in doses are insignificant, being within the uncertainty of the calculations; thus, there is no need to recalculate the radiological ennsequences of any of the accidents due to extending the fuel burnup within the limits of the study performed and S4 reported in WCAP-10125-P-A. In addition, it is noted that the radiological consequences of accidents reported in the FSAR are well within the limits of 10 CFR 100; thus, if increases such as those discussed above were applied to the FSAR doses, thJre would be no impact on their acceptability.
l i nasy s-sness _ _ _ _ _ _ - _ .}}