ML20049H877
| ML20049H877 | |
| Person / Time | |
|---|---|
| Site: | Clinch River |
| Issue date: | 02/28/1982 |
| From: | Coffield R, Lowrie R, Severson W WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML20049H876 | List: |
| References | |
| CRBRP-ARD-0308, CRBRP-ARD-308, NUDOCS 8203040399 | |
| Download: ML20049H877 (97) | |
Text
CRBRP-ARD-0308 Clinch River Breeder Reactor Plant
SUMMARY
REPORT ON THE CURRENT ASSESSMENT OF THE NATURA'L CIRCULATION CAPABILITY WITH THE HETEROGENEOUS CORE February 1982 Prepared for the United States Department of Energy under contracts DE AC15-76CLO2395 and EW 76-C-15-0003.
Any Further Distribution by any Holder of this Document or of the Data Therein to Third Parties Representing Foreign Interest, Foreign Govern-ments, Foreign Companies and Foreign Subsidi-aries or Foreign Divisions of U.S. Companies Should be Coordinated with the Director, Division of Reactor Research and Technology, United States Department of Energy.
@ Westinghouse Electric Corporation ADVANCED REACTORS DIVISION BOX 158 ADI IA 663 O
O 00 37 A
INFORMATION CONCERNING USE OF THIS DOCUMENT PRELIMINARY DOCUMENT This document contains information of a preliminary nature prepared in the course of work for the U.S. Department of Energy. This information is subject to correction or modification upon the collection and evaluation of additional data.
DISCLAIMER This report was prepared as an account of work sponsored by an agency of the United States Government. Neithr the United States Government nor any agency thereof, nor any of their employees, makes any warranty, express or impIled, or assumes any legal liability or responsibility for the accuracy, comp %teness, or usefulness of any informa-tion, apparatus, product, or process disclosed, or represents that its use would not infringe privately owned rights. Reference herein to any specific commercial product, process, or service by trade name, trademark, manufacturer, or otherwise, does not necessarily constitute or imply its endorsement, recommendation, or favoring by the United States Government or any agency thereof.The views and opinions of authors expressed herein do not necessarily state or reflect those of the United States Govern-ment or any agency thereof.
WESTINGHOUSE ELECTRIC CORPORATION ADVANCED REACTORS DIVISION BOX 158 MADISON, PENNSYLVANIA 15663
e CRBRP-ARD-0308
SUMMARY
REPORT ON THE CURRENT ASSESSMENT OF THE NATURAL CIRCULATION CAPABILITY WITH THE HETEROGENEOUS CORE W. J. Severson R. R. Lowrie R. D. Coffield R. A. Markley February,1982 APPROVED BY:
/WW~@
F.
. Fairman, Manager CRBRP Plant ngineering R. M. Viju Manager CRBRP Reactor Analysis Prepared for the U.S. Department of Energy Clinch River Breeder Reactor Project Westinghouse Electric Corporation Advanced Reactors Division P. O. Box 158 Madison, Pennsylvania 15663
P ACKNOWLEDGhENT The authors want especially to recognize the significant contributions of the f ollowing:
J. McCowan J. Lubbe K. Daschke R. Sievers
ABSTRACT The CRBRP has been designed to remove decay heat from the reactor core through the heat transport systems to the steam generator system by natural circulation of the sodium systems and within the steam generator system (the steam drum / evaporator recirculation loop). A preliminary assessment of the natural circulation capability in the plant was made in 1976 which concluded that the CRBRP would provide adequate flows for decay heat removal under natural circulation conditions. Since that time changes have been inade to the core design.
In addition, much attention has been devoted to improving and verifying the natural circulation methodology and input data. As a result of various natural circulation verification tasks, and the availability of component test data, modifications have been made to the plant simulation code, DEMO, and the hot rod analysis code, FORE-2M, which improve the simulation. This report presents a description of the natural circulation event, the analysis methods, input data and results of the current assessment of the CRBRP natural circulation capability with the heterogeneous core.
The current assessment (as well as the 1976 assessment) is based on an
~
analysis of an event involving the loss of all power to the sodium coolant pumps (including pony motors), the steam generator recirculation pumps and the main feed pumps, followed by a reactor scram from full power operation. The results indicate that the peak temperatures (and the time af ter scram that they would occur) seen during the transient in the fuel, inner blanket and racial blanket assemblies would be 1565*F (180 seconds),1544*F (239 seconds) and 1556 F (289 seconds), respectively. With these maximum temperatures, a margin to boiling exceeding 150*F exists. Therefore, it can be concluded from this evaluation, that the Clinch River Breeder Reactor Plant can operate in the natural circulation mode for off normal decay heat removal within acceptable core and blanket temperature limits.
1 i
i 1
TABLE OF CONTENTS Page
1.0 INTRODUCTION
1 2.0 ACCEPTANCE CRITERIA 3
3.0 PLANT ANALYSIS 4
3.1 DESCRIPTION
OF NATURAL CIRCULATION EVENT 4
3.2 DESCRIPTION
OF WHOLE-PLANT ANALYSIS METHOD 5
3.2.1 MODELS 6
3.2.1.1 PUMP C0ASTDOWN 6
3.2.1.2 IN-REACTOR MODELS 9
3.2.1.3 PRIMARY HEAT TRANSPORT SYSTEM MODELS 11 3.2.1.4 INTERMEDIATE HEAT TRANSPORT SYSTEM 11 AND STEAM GENERATOR SYSTEM MODELS 3.2.2 DATA 12 3.2.2.1 DECAY POWER 12 3.2.2.2 R0D WORTHS 13 3.2.2.3 INITIAL CONDITIONS 13 3.2.2.'4 PRESSURE DROP CORRELATIONS 2l 3.3 RESULT S 27 3.3.1 PRIMARY FLOW 30 3.3.2 FACTORS INFLUENCING REACTOR FLOWS 30 3.3.3 REACTOR INLET TEMPERATURE.
34 3.3.4 INTERMEDIATE HEAT TRANSPORT SYSTEM EFFECTS ON 39 PRIMARY FLOW 4.0 CORE ANALYSIS 42 4.1 DESCRIPTION OF CORE RELATED NATURAL CIRCULATION PHENOMENA 42 4.2 UESCRIPTION OF DETAILED CORE ANALYSIS METHOD 43 4.2.1 MODELS USED 43 4.2.1.1 SENSITIVITY TO PARAMETERS 43 4.2.1.2 CONSERVATISMS OF ANALYSIS 47 4.2.1.3 LOCAL HOT R00 MODELING 51 4.2.2 DATA FOR HOT ROD ANALYSIS 56 4.2.2.1 HOT CORE INITIAL CONDITIONS 56 4.2.2.2 PLANT INPUT DATA 56 4.2.2.3 DECAY HEAT 58 11
i TABLE OF CONTENTS (Continued) fage 4.3 RESULTS 59 4.3.1 FUEL, INNER BLANKET AND RADIAL BLANKET HOT R00 59 TEMPERATURES 4.3.2 COMPARISUN OF HOT R00 TEMPERATURES TO SATURATION 63 TEMPERATURE 65 5.0 '
SUMMARY
AND CONCLUSIONS 66
6.0 REFERENCES
1 l
l I
4 l
1 a
t 4
i 111
- - _ ~
LIST OF APPENDICES Appendix Title Page A
SELECTED PLANT TRANSIENT DATA A.1 B
COMPARIS0N OF RESULTS FOR HOMOGENE0US CORE B.1 B.1 DATA COMPARISON B.1 B.2 TRANSIENT COMPARIS0NS B.6 B.3 REFERENCES B.6 C
FACTORS AFFECTING THE NATURAL CIRCULATION RESULTS C.1 C.1 PRINCIPAL DIFFERENCES BETWEEN THE PRELIMINARY C.1 EVALUATION AND THE CURRENT ASSESSMENT C.2 INTRA-AND INTER-ASSEMBLY FLOW AND HEAT C.3 REDISTRIBUTION EFFECT C.3 REFERENCES C.6 l
l 1
iv
1.0 INTRODUCTION
The CRBRP has been designed to provide for decay heat removal through the steam generator system with forced circulation of sodiun in both the primary and intermediate systems in at least one of the three heat transport loops.
Forced circulation during decay heat removal through the sodium loops is provided by the main coolant pumps driven at approximately 10% speed by small pony motors. The pony motors in each loop are connected to one of the three divisions of Class lE AC power. Each Class lE AC power distribution load division has the capability of receiving power from the following four sources:
o Plant Power Supply o
Preferred Power Supply o
Reserve Power Supply Onsite Standby AC Power Supply (Class lE Diesel Generator Sets) o Basically, the pony motor power can be provided with off-site power, or on-site standby power. There is very little likelihood that there will ever be a need for removing decay heat by natural circulation, as there should always be forced circulation in at least one of the loops.
Nevertheless, the capability for decay heat removal by natural circulation is a desirable feature of LMFBRs in general and CRBRP specifically. The inherent safety of the plant is enhanced since no electric power is required to provide adequate circulation in the heat transport or steam generator systems following a scram (i.e., adequate flows are provided even if all six main sodium pump pony motors and the three steam generator recirculation pumps are de-energized).
From the time that design work first began on the CRBRP, it has been a requirement that the plant have the capability to remove decay heat by natural convection. This requirement influenced component and piping arrangemants, as well as certain component hydraulic requirements. The relative elevations of the reactor, IHX and steam generator system components were selected to provide the necessary thermal driving head for adequate loop flows.
In addition, requirements on pump coastdown and stopped rotor characteristics and 1
on primary system cold leg check valve pressure drop versus flow characteristics were established and included in their respective equipment specifications.
In 1976, a preliminary but extensive and detailed evaluation of the natural circulation capability in the plant was made and reported in Reference 1.
The conclusions from that evaluation were that the arrangement of the components will provide the thermal driving heads necessary to promote adequate flows in the natural circulation mode, that the specified pump coastdown characteristics satisfy the minimum flow decay requirement and that the heat transport systems (HTS) and steam generator system provide a desirable margin in decay heat removal by their natural circulation capability.
Since that evaluation was~ performed, the basic core design has changed from one having a relatively " homogeneous" central core region of fuel assemblies surrounded by blanket assemblies to a relatively " heterogeneous" design which includes blanket assemblies interspersed in the central fuel region in l
addition to the outer or radial blanket region. The purpose of this report is to provide results of current analyses that were performed to assess the effect of that design change as well as changes in the modeling of plant related aspects on the natural circulation capability. The results of those analyses are given in Sections 3.3 and 4.3 for the plant and core analyses, respectively. Appendix A includes graphs of various other plant parameters f rom that analysis. Appendix B provides a comparison of the early (1977) heterogeneous core results with the results for the homogeneous core.
Appendix C provides a discussion of other f actors which can potentially affect the magnitude of the core temperatures seen during the natural circulation event.
2
~
2.0 ACCEPTANCE CRITERIA The acceptability of the respoase of the plant to the transition to and operation in the natural circulation decay heat removal mode is judged solely on the resulting temperatures seen in the core during the transient. Even though there is experimental evidence that at low powers and flows stable boiling can exist within the core, the criterion currently used by CRBRP for acceptable natural circulation flows through the reactor are those flows which will limit the maximum core hot spot temperatures (in fuel and blanket assemblies) to less than the sodium saturation temperature. For the CRBRP, this temperature is approximately 1720 F at the top of the active core.
Simply stated, the objective is to preclude boiling anywhere in the core during a natural circulation event.
In order for this criterion to be satisfied during an extended period of operation in the natural circulation mode, the plant response must be such that the intermediate flows are higher than the primary flows and there must in turn be a substantial water recirculation flow through the evaporators in the natural circulation mode. To maintain the maximum thermal driving head in the primary system, an intermediate to primary flow ratio greater than one is required to assure that the IHX thermal center is driven to the top of the tube bundle. Sufficient evaporator recirculation flows are required to force the thermal center in the evaporators to the top of the unit.
3
3.0 PLANT ANAL YSIS
3.1 DESCRIPTION
OF NATURA'. CIRCULATION EVENT The reactor is assumed to be operating at full power (975 MWt) when it is postulated tnat there is a simultaneous loss of power to the main feedwater pumps, steam generator system recirculation pumps and the sodium coolant pumps (iacluding power to the pony motors on both the primary and intermediate pumps). This results in a reactor scram 0.6 seconds af ter the loss of power.
The main coolant pumps coast down and stop. The primary pumps stop in this analysis in 131 seconds. Flow in all sodium loops is then maintained exclusively by the thermal driving heads. Recirculation flow in the steam generator system is likewise maintained by the natural circulation thermal driving head in the recirculation (evaporator / drum) loop. The steam drums in all three loops are furnished with auxiliary feedwater from the turbine driven auxiliary feedwater purap. This flow is available in less than 30 seconds but actual injection into the drums does not occur by design until the level in the drum is 17 inches below the normal drum water level. The final heat sink is supplied by steam venting through the steam generator auxiliary heat removal system (SGAHRS) vent valves, and the protected air cooled condensers (PACCs) controlled on steam drum pressure. The SGAHRS vent valves remove the majority of the decay heat early in the event until the PACCs can handle the heat rejection load and control the steam drum pressure without the assistance of the vent valves.
The transient conditions for the reactor and heat transport systems during the first 30 seconds, are very much like an ordinary plant trip. Following that time, the flows are substantially lower than those produced by the pony motor driven pumps, which are normally operating following plant trips.
The critical period for the natural circulation decay heat removal mode is the first 2 to 5 minutes af ter scram when the flows are low (because the pumps have stopped) and the decay powers are still relatively high. The analysis of the natural circulation event has, therefore, focused on this time period.
4
3.2 DESCRIPTION
OF WHOLE PLANT ANALYSIS METHOD The code used to evaluate the response of the plant to events leading to the removal of decay heat by natural circulation is DEMO (Reference 2).
It is not the intent of this report to describe in detail the calculations used in this plant transient simulation. However, the basic techniques for computing flows and temperatures are summarized below. As the acceptance criterion for the adequacy of the natural circulation capability of the CRBRP is based on reactor temperatures, DEMO is basically used to calculate the reactor flow and inlet temperature as a function of the dynamic response of the plant to the natural circulation event. Calculations are made of average assembly temperatures in order that the core contribution to the total primary thermal head can be calculated; however, the detailed core analysis performed to calculate hot rod temperatures is done using the FORE 2M code (Section 4.0).
In DEM0, the primary and intermediate heat transport system hydraulic flows are calculated using conservation of fluid momentum. The flow equations consider the ef'ects of pump characteristics, fluid iner tia, component pressure losses and fluid thermal heads.
Each component and the required interconnecting piping are nodalized to provide a temperature distribution and an elevation profile about the heat transport loop (including the reactor). The net thermal head for each component or pipe section is computed by the product of each nodal density and its change in elevation. The total loop thermal driving head is then the summation of all compcnent and piping thermal heads.
I During the past four years, the CRBRP project has conducted a natural circulation verification program aimed at verifying the analysis methods (codes) and input data used for the analysis of natural circulation events.
During that time, sensitivity studies have been conducted which were used to identify those parameters or characteristics which are important to the results. The plant analysis results provided in this report have incorporated f
modifications, in some areas, to the models described in DEMO Rev. 4 (Reference 2), as a result of those studies. Some of the more important modeling and data assumptions are discussed below.
I 5
i
3.2.1 MODELS 3.2.1.1 PUMP C0ASTDOWN The pump coastdown characteristic is important to the natural circulation event because of its impact on the flow history up to and including the time that peak temperatures are seen in the core. The basic equations describing the dynamics of the pump are the same as those given in Reference 2 and used in the earlier natural circulation assessment (Reference 1). However, constants in the equations have been modified based or CRBRP prototype pump water test data which were obtained during 1981 Coastdown tests performed during the water test phase demonstrated a significantly longer time to stop than the earlier equations predicted. This has a marked effect on the flow history during the first two minutes of the transient. Thus it has a strong effect on the peak temperatures seen in the fuel and blanket assemblies. The important equations used are:
A.
Pump Head-Flow Relationship H = 458
- li {1.2655 - 0.0377 (U/N) + 0.1079 (U/N)2 2
- 0.3892 (U/N)3 + 0.0971 (U/N)4}
for016/5 1 2.053 2
H = 458
- R {2.371 - 1.155 (U/R)}
f or Q/5 > 2.053 l
where Q = Q/33,700 gpm N = N/lll6 rpm 6
4 B.
Pumping Torque T = R (0.6500 + 0.2738 (4/K) + 0.2512 (4/R)2 2
p
- 0.1325 (4/R)3 + 0.0051 (4/R)4}
where Q = Q/33,700 gpm W = N/lll6 rpm T = T /19,900 f t. Ib.
-l p
p 1
C.
Pump Loss (Friction) Torques i
T = 0.00383 + 0.010715 + 0.01406N L
f or 0.268 < W 11.0 2
T = 0.00268 + 0.07N t
f or 0.01 < N 1 0.268 i
2 T = 0.01 - 73.13R L
for W i 0.01 where i = N/lil6 rpm T = T /19,900 ft. Ib.
L L
2 The pump inertia was determined to be 28,050 lb-f t,
The agreement between the predicted coastdown using these equations and the water test data is shown in Figure 3-1.
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As discussed in Appendix C, the time to reach 10% speed following pump trip, is 38 seconds with the modified equations versus 30 seconds for the equations used in the Reference 1 analysis; and the time to stop is 131 seconds for the primary pump rather than 55 seconds as shown in Reference 1.
The impact of this extended coastdown is to delay the initiation of natural circulation flow and, therefore, the addition of the pump locked rotor resistance to the loop impedance. This allows time for a reduction in the decay heat as well as the reactor sensible heat.
3.2.1.2 IN-REACTOR MODELS The reactor models used in DEM0 provide sufficient detail for use in calculating the primary system natural head and dynamic pressure drops as well as calculating the core inlet conditions and average channel temperatures.
The details (equations used) for the in-reactor models are given in Reference 2.
These models include the following:
A.
The nuclear power is simulated by a neutron kinetics equation employing 6 delay neutron groups and a prompt jump approximation.
I B.
The total reactivity is the sum of the control reactivity, the reactivity due to sodium expansion and the reactivity due to Doppler.
C.
Reactivity and reactivity rate insertion are computed with taoular functions.
D.
Decay heats for each region (fuel, inner blanket and radial blanket) are based on separate calculations and read into DEMO from tables (i.e., DEMO does not calculate decay heats directly).
E.
The time variation of thermal power for each channel is simulated by the combination of the neutron power and decay power.
F.
The average channel of the active core (fuel assemblies) is represented by 7 axial nodes and 5 radial nodes (2 fuel, one 9
cladding, one sodium coolant and one duct). The above core region (top of upper axial blanket to top of fuel assembly) employs 5 axial nodes.
1 G.
The average inner and radial blanket channels are represented by 7 axial nodes and 6 radial nodes (4 fuel, one cladding and one coolant).
H.
The total flow through the various regions (fuel, inner blanket, radial blanket and bypass) are based on balancing the dynamic losses and static heads through each region. Thus the code calculates flow redistribution between the four regions.
I.
The upper internals structure (UIS) has been explicitly modeled. The abrupt drop in the power to flow ratio immediately af ter scram results in an overcooling of the core and the insertion of relatively cold sodium into the upper core assembly regions and on into the UIS structure. When colder fluid enters the UIS chimneys, there is a suppression of the thermal head and the chimney flow fraction of the total reactor flow decreases, thus forcing more flow through the radial gap between the tops of the assemblies and the bottom of the UIS skirt. This has the effect of increasing the total resistance of the primary flow path. Once the temperature out of the core increases, and is seen in the UIS, the UIS thermal head recovers and the flow increases through the chimneys.
In fact, once it recovers fully, colder sodium (f rom the bottom of the upper plenum) is asperated back through the UIS radial gap and the chimney flow actually exceeds the reactor flow.
It is assumed that all chimneys can be represented by a single chimney (simulated with 10 axial nodes) and that the available heat for warming the relatively cold sodium entering the chimneys early in the transient comes from the UIS structure and surrounding sodium within the UIS shroud.
J.
Two distinct modes of mixing are used to simulate the mixing in the reactor upper plenum. The first mode is a single node mixing with
-metal heat capacity. The second mode is a stratified-flow model where the reactor exit temperature is related to the sodium 10
temperature exiting the core by a first order lag. The selection between the two modes of mixing is based on the jet height of the sodium coolant exiting from the core.
If the jet height is greater than 20 feet, the single mode mixing is used.
if the jet height is less than 20 feet the stratified flow model is used. The crossover point occurs af ter a time delay which is equal to the time required 3
to fill a preset volume (837 f t ) by the jet exiting from the core, K.
The lower or inlet plenum is based on a single mixing node with metal heat capacity used to represent the lower plenum structure.
The modifications to DEMO 4 required to represent the heterogeneous core involved the addition of an inner blanket, changes in the power frac' ions, flow fractions and decay heats.
3.2.1.3 PRIMARY HEAT TRANSPORT SYSTEM MODELS The models for the check valve and the IHX used in the natural circulation evaluation for the heterogeneous core were identical to those used in the analyses reported in the initial preliminary a'sessment for the homogeneous core (Reference 1). The IHX calculations are based on one-dimensional models. The 26-foot long tube bundle is represented by 45 axial nodes. The check valve model accounts for the closing of the disc as the flow decreases.
The current assessment has added modeling to account for heat capacities of piping, pump and IHX plena structure. There are 88 axial piping nodes in each primary loop. The details of these models are given in Reference 3.
The exchange of energy between these structures and the flowing sodium modifies the spatial temperature distribution as the transient progresses and thus, A separa' e study (Reference 3) showed affects the calculated natural heads.
t that the effect on flows (and thus, core temperatures) was small, d,an all other models remain unchanged. However, the thermal inertia of these structures represent a real effect and therefore has been included in this analysis.
11
3.2.1.4 INTERMEDIATE HEAT TRANSPORT SYSTEM AND STEAM GENERATOR SYSTEM MODELS s
The hydraulic modeling for the intermediate heat transport system (including the sodium side of the steam generator modules) is identical to that of the
' primary system as described in Section 3.2.1.3.
There are 104 and 128 axial nodes in the S and L intermediate loops respectively. The steam generator modules have seven axial nodes representing the 46 feet of active length.
^
The steam drum simulation is based on a homogeneous model with mass and energy balances maintained throughout the transient. The auxiliary feedwater sparger has been modeled by assuming that the feedwater is heated to saturation in the steam space of the steam drum. This effect was simulated by reducing the enthalpy of the steam / water mixture entering the drum by an amount equivalent i
to that required to accomplish the heating of the auxiliary feedwater to the saturation temperature. Steam flows through the superheater are based on drum pressure and SGAHRS vent valve characteristics. The recirculation flows through the evaporator are calculated based on the natural heads around the system and the associated drum / recirculation loop / evaporator pressure drops which account for the locked rotor drop through the recirculation pump, two phase pressure drop multipliers in the evaporators and evaporator to drum piping including the check valve in this line.
3.2.2 DATA 3.2.2.1 DECAY POWER The DEMO reactor core model used in the analysis is a four region model.
It calculates the conditions for an average fuel assembly, average inner blanket assembly, average ' outer blanket assembly and a bypass channel. The data used for the core decay power (fraction of operating power at time of scram) as a function of time af ter shutdown is shown in Table 3.1.
Decay heat contributors evaluated in the analysis are fission products, actinides, stainless steel activation and the transuranics. All data sets used included t
uncertainties.
The data presented are based on FFTF-grade fuel with a five-year shelf life prior to use in the reactor. Uncertainties on fission
[.
' products and transuranics are based on ENDF/B-IV and ZPPR-7 analyses. The uncertainty associated with actinides and stainless steel activation is +20%.
12 s
in making this natural circulation assessment, the core conditions used were those existing at end-of-cycle 4 (EOC4).
-3.2.2.2 ROD WORTHS s
The negative reactivity inserted upon scro affects the delayed neutron power and thus when combined with the decay power, the total power as a function of m
' ~ time after scram.
In the OEMO analysis, a negative reactivity of $24.92 was ps -
. inserted following the scram which occurred 0.6 seconds following the pump trips. This conservative value for negative reactivity reduces the power to flow ratio and hence reduces the thermal head early in the transient. As will be discussed in Section 4.2.12, the core temperatures were calculated with minimum design values.
It should be pointed out that the difference in peak temperatures between insertion of both the primary and secondary rods (which are inserted 1.4 seconds af ter the pump trips) and insertion of only the secondary rods is not large. Figures 5.6 through 5.8 of Reference 1 show this effect.
In this analysis, which employs updated pump equations resulting in a much longer pump coastdown, the differences in peak temperatures due to the dif ferences in the amount of negative reactivity insertion would be expected to be e'ven smaller. The temporal changes in neutron power and total power are shown in Figure 3.2.
3.2.2.3 IllkTIAL CONDITIONS Initial conditions used for the
- ',u
- re given in Table 3.2.
Also shown for comparison purposes are the Ant;
.jd aulic Design Values (THDV) for the l
plant. The initial conditions used for the analysis of the natural circulation event are basically the thermal hydraulic design conditions except that the hot and cold leg temperatures have been set approximately 20*F higher to account for possible operational uncertainties includiag instrumentation and control errors.
It should be noted that the temperatures used are considerably higher than plant expected conditiens since they are basically an adjustment on the THDV conditions. The THDVs were established for the purpose of sizing components (IHXs, SG5 modules, pumps, piping, etc.) and thus represent extremes. For 13
TABLE 3.1 AND RE ACT3R CORE DECAY P' NER AS A FRA;r ION OF h/UNCE FTAINTIE S AVERAGE ASSE4BLY OPERATING 80WER - VALUES END OF OYCLE P/ F(0)
TIME AFTER 4WERAGE AV ER AGE AVERAGE REACTOR FUEL IN DER BLANKET RADI AL BLANKET
. CORE ASSEN BLY SauTDOWN ASS EN8Ly 4SSEMBLY (SE-ONOS) 7.246E-32 6.515E ~2 8.702E-02 7 616E-CZ 1 2 0, cr E-3 7 211E-;!
5 666E-02 8.024E-02 8.244E-02 i.30: 0f E+00 1 134CI E+( 0 6 75tE-02
- 7. 60 C E-02
- 7. 45 3E-C2
- 7. 7 09E-0 2
- 6. 732E-; 2 2 0310f E+60 6.3',E-02
- 7. 3 0 5 E '. 2
- 7. 5 7 CE-3 2
- 6. 60 6E-: 2 2 4240r En a 2 5940r E+tG 5 177E-02 7.156E-02
- 6. 857 E-02 7.147E-02 6.214E-2 4.1250CE+0c 5 926E-02
- 6. 70 7E-32
- 7. L Q 4E-0 2 6.080E-;2 5.rSoE-02
- 6. 92. C t E + 3. 0
- 6. 55 S E-52 E. 8 6 (E-0 2 5.943E-t 2 5 565E-02
- 5. 37: 7f E+t c
- 6. 40 3 E-02 E.716E-0 2 5.805E-:2
6.3770fE+,0 5 39EE-02 l
6.~99E-02
- 6. 4 2 SE-C 2 5.525E-:2 1.
' ' 3 f E+ - 1 5.25'E-02 6.278E-02 5.382E- !
5 545E-32 1.19 4 3f E+ 1 5.11dE-02 i
- - - ~ _
TABLE 3.1 AVERAGE ASSI1BLY AND RE ACTOR CCRE 3ECAy P0hER AS A FRA TION OF END OF CYCLE OPERATING POWER - VALUES b/UNCERTAINIIES P/F(0)
TIME AFTER AVE 4GE AVERAGE AVEFAGE REACTOR SHUTDOWN FUIL INNER BLANKET RADIAL 8LANKET CORE (SE30N05) 45SEMBLY ASS EM BLY AS S EM ELY 1 42 5 00 E+ C 1 4 377E-02
- 5. 63 7E-02
- 5. 9 8 EE-0 2 5.091E-;2 2.G310CE+C1 4.687E-02 5.483E-02
- 5. 8 3 4E-0 2 4.944E -: 2 2 4240C E,ri
- 4. 54 2E- 02
- 5. 3 29 E-02
- 5. 68 6E-0 2 4.797E-t!
2 8940rE+ii 4.395E-02 5.177E-72 5.538E-02 4.649E : 2 3.455cCE+si 4.249E-02 5.(25E-02
- 4. 72 5E o 2 5.10 CE-0 2
- 4. 2 0 9E-; 2
- 5. 97 3 0C E+ 1 3 513E-02
- 4. 5 79 E-C 2
- 4. 95 7E-0 2 4.0 64E-; 2 7.31700E+fi 3 67:E-32
- 4. 4 34 E-0 2 4.ei!E-02 3.921E-C2 E.3773rE+01 3 53:E-02 4.292E _2
- 4. 67 EE-0 2
- 3. 781E-t 2 1.*9:ME+ 2 3.393E-02
+.15 5E-0 2 4.541E-02 3.644E-:Z 1.1943CE+/2
- 3. 262E- 02 4 022E-C2
- 4. 41 CE-0 2
- 3. 513E-; 2 l
1.4250fE+62 3.13eE-02 3 996E-?2 4.28!E-n2 3.389E-:2 1.7ai?rE+ 2
- 3. 0 21E- 02 3 77 EE-22 4.16 EE-0 2 3.271E-;2
- 2. 0311r E+ ' 2 2.911E-C2 3 663E-;2 4.C51E-02 3.16CE-;2 e
_. ~___.__._ _
E
TABLE 3.1 A FRA3T ION OF CORE DECAY POLER AS OPERATING POWER - VALUES N/UNCEFTAINTIE S AND RE ACTOR AWERAGE ASSEMBLY END OF 3rCLE P/ F(t)
AVEFAGE REACTOR SWERAGE C0fiE TIME AFTER INfE R BLANKET R ACI AL 8LANKET AWERAGE FUEL ASSEMBLY SHUT 00WN ASSI M ELY ASSEMBLY (SE00NDS) 3.G 56E-: 2 2 942E-02 3 555E-02 2.SCBE-02 2.957E-:2 2.4240fE*C2 2 8 36E-02 3 452E-32 2 711E-02 2.862E-:2
- 2. 894 0C E+ t 2 3 733E-02
- 2. 676E-s !
3.52 E-02 3 152E-02 2 43oE-02 2.582E-:2 l
4.92500E+;2
- 3. 414E-0 2
- 3. C 50 E-J2 2.35'E-02 5.8793rE+'2 2.30fE-02 2.487E-:2 2 945 E-02 2.26:E-G2
- 7. 017 00 E+ s 2 3.18 2E-02 2 389E-;2 2e 83 EE-C2 2.16fE-02 6.3770 ret:2 3 059E-02 2 2 89E-: 2 2 722E-22 2.375E-C2 2 187E-12 1.00~0rE+.3 2 929E-02 2 60 5E-32
- 1. 9 8 E- 02 1 19 4 0f E+. 3 2 79 EE-02 2 08 3E-f 2 2.I,e AE-02 1.SS4E-02 1 4250tE*C3 2 66 E-02 1.978E.2
- 2. 36 2E-02 1.TeCE-02 1.873E-I!
1.7310fD 3 2 5 2 EE-02 2.238E-02
- 1. 692E- 02 1.769E-:2 2.0314fE+r3 2 3 85E-02 2.115E-32 1.59fE-t2 1 6 69E ~ 2 2.4240rE*f3 2 253E-02 l
- 1. 997 E ~ 2
- 1. 5 0 7E- 0 2 1 574E-: 2 2.6940rE+'3 2 128E-02
- 1. 885 E-;2 1.42;E-02 3.455crE+'3
TABLE 3.1 CORE DECAY PONER AS A FRA3r ION OF AND RE ACT3R AVER AGE ASSEqsty OPERAT 7NG POWE R - VALUES W/UNCE FTAINTIE S END OF CYCLE P/P(3)
TIME AFTER AVERAGE AVERAGE AVEFAGE REACTO R SquTDOWN FUEL IN hE R 3 L ANKET RADI AL BLANKET CORE (SECONDS) 1SSEMBEY A S SI M ELy ASSEM ELY 4 1250tE*C3 L.335E-02
- 1. 7 8 0 E-12 2.011E-02 1 485 -t2 E
4 9250tE+03
- 1. 2 6 3E- 02 1 685E-02
- 1. 9 0 6E-0 2 1 40 3E-! !
5 875 0CE+0 3 1 193E-02 1 60 CE-02
- 1. 459 E-J2
- 1. 40 CE-;2
- 1. 5 9 E E -0 2 1.144E-;!
2 16 3 0f E+ G 4 8 26?E-D3 1.19 2E-0 2
- 1. 3 7 7E -0 2 9.47 4E. 3 4 32 0fE+r4 5 904E-03 1.028E-C2 1.19tE-C2 8.016E : 3 6.4AC00E+.4 5 173E-13 9 297E-03
- 1. t 8 7E-0 2
- 7. 2 03E-i 3
- 8. 643 0f; E+ C 4 5 683E-03 S. 566E
'3
- 1. t G 2E-0 2 6.E34E-C3 1.29630E+(5
- 5. ? 3.E-3 3
- 7. 4 6T E-3
- 6. 63C E-03
- 7. 75 7E -0 3 5.267E-;3 l
2.5920CE+:5
- 3. 3 3 3 E- 0 3
- 5. 3 6 7E-a 3 6.17 EE-0 3
- .462E-23 l
3.4560f E "5 3 388E-33
- 4. 498 E-43 5.G7;E-03 3.903E-3 4.32'0fE+i5
- 3. 2 5 5 E- 0 3 3 536E-03 4.25;E-03 3.481E-:3 1
- 3. 332E-T 3
- 3. 631E-0 3 3.158E 3
5.1540CE+'5 3.04EE-03 Mi
TABLE 3.1 P0hER AS A FRA;T ION OF AND RE ACTOR CORE DECAY MUNCERT AINTIE S AVERAGE ASSEMBLYEND OF OYCLE OPERAT ING POWER - VALUES P/P(0)
REACTOR AVE RAGE COE A VE E A GE RADIAL 8LANKET
%dERAGE INER 8L ANKET ASSEMBLY TIME AFTER FUEL A SSENELY SHUTOOWN
%SSEMBLY (SE00N053 2 697E-Is 2.79 1-03 2 644E-03 2 597E-03 2 392E-23 6 912 00 E+ C 5
- 2. 28 E-0 3 2.217E-03 2 45'.E 03 t.64: OC E605 i
I 8
h H
E
~
d
-Y S.h es I
.c Y
14
]
8 i
l l
i
/
S
/
l S
/
'/
l
-W V
.o 2
2 8
8 4
8 2
9 Det - 83eed IM113V3tl
+
Figure 3.2 19 4
E TABLE 3.2 INITIAL CONDITIONS USED FOR ANALYSIS OF NATURAL CIRCULATION EVENT 1
Thermal Hydraulic Parameter Value Design Value Reactor Total Power MW 974.8 975 6 Total Flow lbfhr 41.49 x 106 41.45 x 10 Inlet Temp
'F 748.2 730.
Outlet Temp F
1012.
995.
Primary System Loop Flow lb/hr 13.83 x 106 13.82 x 106 Pump Speed RPM 1077.
Pump Total Head psid 155.4 163.5 Hot Leg Temp
- F-1012.
995.
. Cold Leg Temp
'F 748.2 730.
4 4
i Int.rmediate System
'oop Flow lb/hr 12.78 x 106 12.78 x 106 Pump Speed RPM 951.7 (S) 994.0 (L)
Pump Head psid 117.3 (S) 133.3 (L) 126.
Hot Leg Temp
'F 951.1 936.
Cold Leg Temp
- F 671.0 651.
Steam Generator System Steam Flow lb/hr 1.07 x 106 1,11 x 106 SH Outlet Temp
'F 919.1 906.
2.22x10g..
155 SH Outlat Press psia 1534 Evap Flow lb/hr 2.22 x 106 Drum Pressure psia 1840.
1889.
t 20
example, the THDV conditions for the IHX yields an LMTD of 68.5 F but the expected IHX performance is such that the IHX at the end of life will, on a best estimate basis, require an LMTD of only 61*F.
In fact, the nominal expected reactor inlet and outlet temperatures for the CRBRP are approximately l
715*F and 957 F, respectively. The significance of this is that the peak temperatures seen in the reactor during the event have been calculated based on a reactor inlet temperature that is 33 F higher than expected values.
Thus, the analysis has been performed from a conservative set of initial conditions.
3.2.2.4 PRESSURE DROP CORRELATIONS The DEMO Rev. 4 code (Reference 2) pressure drops were basically calculated as a function of the square of the flow (except for the reactor, the check valve and the steam generator recirculation loop).
In the analyses used to provide the results given in this report, these correlations have been changed to l
account for Reynolds number dependent friction f actors which in most cases were determined by component testing. The pressure drop relationships used in the analysis are as follows:
A.
Primary loop. The three primary loops are essentially identical.
The correlation shown in Figure 3.3 includes the reactor to pump piping, the IHX, the piping between the IHX and the pump, the pump to reactor vessel piping and the reactor inlet and outlet nozzles. The pressure drop at 3% flow is 73% higher than it would be if the aP were assumed to be proportional to flow squared.
B.
Cold leg check valve. The pressure drop relationship for the check l
valve is shown in Figure 3.4.
The pressure drop through this component is basically a form loss; however, the correlation reflects l
the movement of the disc as the flow decreases. The disc begins to l
close when the flow reaches 20% and moves to a minimum hang angle at 3%.
The correlation used for the 24 inch CRBRP check valve pressure drop as a function of flow is a correlation developed from water tests of 21
PHTS (Piping, IHX, Reactor Inlet and Outlet Nozzles)
Pressure Drop Characteristic (Maximum) 1 fi!1A 1 : n. a.
l l
g
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5
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- .:n =
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- ==. n:
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n:3 gp.:1:un --
y rn:
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.= = m=
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n= m -=
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=
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- .n..n. :;.":n...=
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0 I $ 1 1
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Figure 3.3 22
a 16 inch check valve (FFTF prototype) which is very similar to the CRBRP valves. A nominal or expected aP correlation was developed along with an appropriate uncertainty. The analysis used the maximum pressure drop curve.
C.
Pump locked rotor resistance. When the pump reaches 0.1% speed, it is assumed to be stopped and the locked rotor resistance is inserted into the loop pressure drop. The relationship shown in Figure 3.5 is based on data from the CRBRP prototype pump water tests. At 3% flow, the correlation gives a aP of 0.1134 psi which is 42% higher than the value of 0.08 psi used in the Reference 1 analysis.
D.
Reactor core. The core pressure drop versus flow relationship used in this analysis is shown in Figure 3,6.
This pressure drop is the pressure loss due to friction and form losses through the lower inlet modules, and core assemblies.
It should be noted that the pressure drop shown is consistent with that which would give a total reactor nozzle to nozzle pressure drop of 123 psi at the design flow of 6
41.446 x 10 lb/hr. This is the maximum allowable reactor pressure drop. Actually, the core is being orificed such that at rated pump speed, the nominal core flow will be 1.1 times the design flow.
Thus, the pressure drops used in this analysis are conservative.
E.
Intermediate Heat Transport System.
In the intermediate system, individual pressure drop correlations for the IHX, piping and steam generator modules are used. The IHX and steam generator pressure drop correlations are based on key feature testing.
Since the piping lengths are different for the three intermediate loops, the pressure drops are different. As stated before, the S-loop simulates Loop 2 in the plant (the shortest intermediate loop) and the L-loop simulates Loop 3 in the plant (the longest intermediate loop). The pressure drop versus flow relationships for the 5 and L intermediate loops are shown in Figures 3.7 and 3.8.
The intermediate pump is hydraulically identical to the primary pump so the same correlation is used.
23
Primary Cold Leg Check Valve Pressure Drop Characteristic (Maximum) r_100 I
i
+::!!h:=iiiiE iill nu :
1 3
4, -
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h : iiHi l'il !!h iiF5:".
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n ti + t:4::. ;gs:n::- ~~ ,I "wn
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CRBRP Prototype Pump Locked Rotor Resistance (Nominal) 1000 I'ilI i!i 11!!i lihiti ti' i 8 i': tjiiji {i!!!!8}! hhF fi' iiii-lH! it' ' ii g .h ..,! ! : n" : Lih!.h: ,, t ;. ;;. ;;; .,t ' niHilli !!!!!!!" titiiliiiM tili-liti tili#iti El# ;" iiii :iii Rimi!!!!!iiihr,M f~ "bii!M ils liti M -E ".'l W "~ "~ 13 - n:l==:.M hu: = r: =3:=-.
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Reactor Plenum to Plenum Pressure Drop Characteristic i (Itaximum) 1000 ,.tj t { {. l,, j j .u=:
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F. Steam Generator System. Steam side pressure drops through the superheater are assumed to be proportional to flow squared. In the evaporator / steam drum recirculation loop, the locked rotor resistance of the recirculation pump is accounted for as is the check valve in the piping between evaporat Jr and the drum. It should be noted that the recirculation flow, under natural circulation conditions, is not very sensitive to uncertainties in recirculation loop pressure drops since a lower flow due to higher than expected pressure drops would result in higher voids within the evaporator which in turn provides a higher thermal head. The recirculation flow rates calculated are more than adequate to force the sodium side thermal center towards the top of the unit. 3.3 RESULTS Based on the criterion for acceptable natural circulation decay heat removal (i.e., preclude boiling in the core), the analysis results show that the time interval of interest for the event is the first few hundred seconds. Thus, the analysis results given in this report for the first 500 seconds of the transient covers the critical period of the event. Following that time, the thermal heads are_well established; the rate of change of flow is small in the PHTS, IHTS and steam generator recirculation loops; the temperatures in the hot legs decrease slowly and the transient is well behaved. Data from natural circulation tests conducted at FFTF and previous extended natural circulation analyses shows this same characteristic. The whole-plant analysis performed with DEM0 has as its principal outputs, the calculation of the total reactor flow and the temperature at the core assembly inlet as a function of time. These results are then used as input conditions for the core analysis discussed in Section 4. The calculated reactor flow and assembly inlet temperature results as well as a discussion of the major f actors which impact these results are presented below. 27
IHTS Pressure Drop Cnaracteristic (S-Loop, Maximum) ,, 1000 iifi!! ih t i I 2 N: N!i !!- . i..: :::. t:. g id i!!! P' :!N@'i5ii !!!! '.ili W - 1 !!:: E !I!!!!!!! 'i% H W Bmtsec 4 3 .3
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3.3.1 PRIMARY FLOW ' The primary system flow is a function of the pump speed until the pumping head becomes small relative to the thermal head. When the pump stops, of course, the flow is solely driven by' the natural or thermal head in the primary . system (PHTS loops plus reactor). The flow transient is shown in Figure 3.9. Figure 3.10 shows an expansion of the curve f rom time t = 40 seconds to 500 seconds. The rather abrupt change in flow seen at 130 seconds af ter scram corresponds to the time the pump stops. The dip in flow following t = 130 seconds reflects the increased system resistance due to the locked rotor impedance of the pump. The flow dips to a minimum of 2.3% at 144 seconds af ter scram followed by a steady, gradual increase to 2.9% at 500 seconds as the thermal head increases. It should be noted that there is a significant contrast between the reactor flow from this analysis and that shown in the report on the preliminary assessment for CRBRP (Reference 1). The pump coastdown calculation in the earlier analysis used a friction torque relationship known at the time to be very conservative. it resulted in a pump stop time of 55 seconds. The flow shown in Figure 3.11 demonstrates the effect of the longer coastdown (principally resulting from much lower f riction torques) as seen in prototype - pump tests. The quasi steady state flow shown in Figure 3.10 is somewhat lower than the earlier predictions due to higher system pressure drops. It should be noted that the thermal heads calculated in the two analyses are comparable. 3.3.2 FACTORS INFLUENCING REACTOR FLOWS Early in the transient, the sodium flows are determined by the dissipation of thc kinetic energy in the pumps and sodium. The primary pump speed to flow ratio is shown in Figure 3.12. When this ratin gets to approximately 0.5, the pump is producing zero head. This marks the transition to the energy disipation regime which occurs at about 120 seconds into the transient. From that time on, the flows are produced by the thermal head in the system..The corresponding curve for the intermediate system is shown in Figure 3.13. It has a distinctly different shape'due to the higher f raction of fluid' inertia in the intermediate system compared to the primary system. 30 l l _____._____.___m_
.m g u> o (, O LI) \\ a> .o O. - q i .o D3 Q .u x 3 o o 9' 0 . uiis : Il.g LtIMy 't .j-2 s 9 X b. ~ d .onN 9 $= ne m h u -? m a] L> t> __.,_a.-.....--r---r--~~~~ q t__ O a o o o o o O o._ r o N 0 33S/H8'l - M07dSSiH 11X3 d)ind 1.6%I!Ud.PXl-S Figure 3. 9 31
) H eL u h I .h ~ . a. W n U M \\ . i I ( ) E ~ 8 i, / p i. 1 .o l k k k b b E E 9 03S/H8~1 - M8"Id88VN 11X3 efid AWHied me 1-3 Figure 3.10 32 l -a
1\\ 10 g \\ LEGErlD: g fk Analysis with Homogeneous Core Current Analysis with Heterogeneous Core \\ 8 \\ y 7 p 5 \\ M 6 E \\ 2 8 \\ y g 5 \\ F \\ \\. E 4 x \\ 3 2 1 0 I I I I I I I l l 1 20 t-0 60 80 100 120 140 160 180 200 TIME (SEC) l
The primary system thermal head is shown in Figure 3.14. The PHTS thermal head is approximately 0.25 psi at the full power condition, dips to nearly 0.08 psi at about 18 seconds and then recovers to 0.4 psi at 400 seconds. The initial drop in the thermal head is due to the overcooling of the core immediately af ter-scram when the power to flow ratio drops precipitiously. The thermal head at the time the pumps stops is just under 0.26 psi. The thermal head calculated at each time step is determined by summing the temperature dependent sodium density in each node around the loop where an elevation change is involved. Thus, the temperature distribution through the core, in the above core region, in the UIS, in the piping system and in the I IHX all affect the thermal head at any point in time. This thermal head, in conjunction with the pressure drop relationships around the loop, establishes the flow. The distribution of the system pressure drops is markedly different under natural circulation than it is at the full power condition. In the primary system, the system pressure drops are significantly affected by the stopped pump and the change in the check valve disc position. In the intermediate system, the stopped pump provides the major impact on the loop pressure drop. 3.3.3 REACTOR INLET TEMPERATURE Due to the relatively long transport time in the HTS, the reactor inlet temperature (Figure 3.15) remains unchanged for the first 20 seconds of the transient.. Thereafter it drops slowly to about 680 F at 500 seconds into the transient reflecting the change in the IHX primary outlet temperature, the transport delays and the effect of the heat capacity in the piping between the IHX and the reactor vessel inlet nozzle. The IHX primary outlet nozzle temperature transient is also shown in Figure 3.15. The relatively rapid down transient seen early.in the event is typical of all trip transients and reflects the collapse of the IHX cold end temperatures on each other (primary outlet to intermediate. inlet at the bottom of the tube bundle) when the primary and intermediate flows coast down. The primary IHX outlet temperature reaches 685 F at about 90 seconds into the transient. However, this temperature is not seen at the reactor vessel inlet until about 425 seconds after the trip due to the transport delay and energy exchange between the sodium and piping. 34
+ a I," ~ I_ 8 T i ~ ~ b W - ag ws ~ H ~ E 4 og ~ E m [ l> i i n s s s x n 9 O O O O O chi el1V8 Me73/033ds 18d S Figure 3.12 + 33
d I" ~ U 8 ~ b W ~ $'l ws ~ p N E E k ~ l 7 ~ co 8 i- ~ t> g m E R i I O W T 9 9 O O O O O O LpI 13V83 - M07d/033dS dH0d 1NI S Figure 3.13 + 36
+ a { a 8 1 1 3_ h m d .W 95 g WE i~ I. .t. 5_ \\ 1 I i 8 s i i, ) d e x e e O O O O O O Let VISd - t A3MlH31 OV31 IVN d881 IBd S Figure 3.14 + 37 +
l u l h f' s 4, 1 g ' s
- /
s s o I l s o l r l l h I s l I h r
- i
- sl 5
Y 1 .u n1 s .g g S 52 s s ~ g-! W hs _gE g g-M l C-EE ~ " hit s caps S l ~ 'Nh 2_ i i 8 / / l ,,./ ..A. W .d h k h .i *030 - 3WuYH3dG10310103 + Figure 3.15 38
The temperature at the fuel assembly inlet shows a very slow response to the reactor vessel inlet transient. For example, at 400 seconds into the transient, the lower blanket inlet temperature has only decreased 23*F while the reactor vessel inlet temperature has decreased approximately 60*F. This response reflects the large amount of stored energy in the reactor vessel inlet plenum and surrounding structures. 3.3.4 INTERMCDIATE HEAT TRANSPORT SYSTEM EFFECTS ON PRIMARY FLOW The intermediate system affects the primary system through its effect on the time dependent temperature distribution along the tube bundle in the IHX. As the flows coast down, the temperature difference between the primary and intermediate sodium within the IHX collapses. As the transient progresses, the thermal center undergoes a shif t towards the top of the unit if the intermediate mass flow rate remains higher than the primary system flow. This condition, in f act, persists throughout the transient. For example, as seen in Figure 3.16, the IHTS flow in the S-loop is 165 lb/sec at the time the IHTS pump stops (t = 130 seconds) compared to about 100 lb/sec in the same PHTS loop. At 400 seconds into the event, the intermediate to primary mass flow rate ratio is approximately 1.4. A comparison of the S-loop primary, intermediate and steam generator recirculation flows is provided in Figure 3.17. Because of the long transport delays associated with the IHTS cold leg piping, there is virtually no change in the IHX intermediate inlet temperature for the first 200 seconds in the S-loop (Figure 3.15) and considerably longer in the L-loops. (The S-loop represents Loop 2 in the plant which has the shortest piping runs between the IHX and the steam generator, and the L-loop simulates Loop 3, which has the longest piping runs.) Thus during the time frame in which the peak core temperatures are seen, temperature changes in the IHTS resulting from the collapse of the evaporator outlet temperatures onto the incoming water temperature have little or no effect on primary outlet temperatures at the IHX. i i l v'; < l 39 \\ l
f a i i' ,l 1 / \\ .d . ag W r-( ( k l I / li 8._ i - y / / i, I E 9 9 t a g g g .O h k k k k b OS 339/HG7 - MeldSSYH del 1NI 8 Figure 3.16 ~ 40 t
3 o Q h Y o i g ga d d 8 e.3 uw .m S U u t j ~ ,5u i .C III E s@ S ] sq 2_ 1 r j 8 f l ( 1 l 0 --- =~ ~~ ~ e. m. w n. =. o. o o o o o o o WillNI de Nell3VHJ - Sme 13 Figure 3.17 41 I
d 4.0 CORE ANALYSIS i
4.1 DESCRIPTION
OF CORE RELATED NATURAL CIRCULATION PHENOMENA For the core natural convection cooling mode, the effect of dynamically i approaching low flow with worst case decay heat loads results in a power-to-flow ratio greater than one. Consequently, core temperatures increase and natural convection phenomena such as inter-and intra-assembly flow redistribution due to different thermal heads and hydraulic characteristics of the core assemblies become important. In general, the core thermal head j becomes significant relative to the form and friction losses across the core below approximately 5% of full flow. Coupled with the flow redistribution, significant heat redistribut'an on an inter-and intra-assembly basis occurs throughout the core due to large temperature differentials and an increased heat transport time (low power assemblies can have a transport time of over 20 j seconds). These effects (i.e., natural convection flow and heat j redistribution) are found to significantly reduce maximum core temperatures (compared to analyses which neglect them) in core system tests as demonstrated in the EBR-il experiments (Reference 4) and FFTF natural circulation tests i (Paference 5). The ARD Blanket Heat Transfer Tests (Reference 6) and ORNL-THORS Fuel Bundle Heat Transfer Tests (Reference 7) demonstrated the significant beneficial effects of intra-assembly heat and flow redistribution at low flows. Independent studies outside the CRBRP Project have been publi:hed which also show a significant decrease in predicted maximum core temperatures due to reactor flow redistribution during natural circulation conditions. For example, Brookhaven National Laboratory (Agrawal, et al., Reference 8), using the SSC-L code, predicted localized flow fraction increases as large as'20% in the hot fuel assembly and 40% in the hot blanket assembly for the CRBRP during i natural convection. Temperature reductions on the order of 16% and 22% '(sl30*F and 210*F) were shown for the hot fuel and blanket assemblies, respectively, relative to the maximum temperatures predicted without flow redistribution. Similar results due to inter-assembly flow redistribution only were found in Reference 9 using ts.e CURL-L code. For both of these u aforementioned studies, inter-assembly heat transfer as well as intra-assembly 4 h 42
i flow redistribution and heat conduction effects were neglected. Inclusion of l these effects would further reduce the maximum core temperatures (References 8 and 9). In summary, n& Wral convection cooling of the core is a core design transient where low power /high temperature conditions exist. Due to the long coolant l transport time and low pressure drop for the core while descending into and l operating in this mode, core inter-assembly and intra-assembly flow / heat i redistribution: a) becomes significant with regard to accurately predicting temperatures; and b) significantly decreases the maximum hot rod temperatures in all core regions, when compared to analyses which neglect these effects. l The CRBRP core temperature predictions, described in Section 4.3, neglected l the aforementioned intra-and inter-assembly affects and therefore are conservative.
4.2 DESCRIPTION
OF DETAILED CORE ANALYSIS METHOD 4.2.1 MODELS 4.2.1.1 SENSITIVITY TO PARAMETERS As stated in Section 2.0, the criterion for judging the natural circulation decay heat removal capability is based on the maximum core temperatures. In particular, the core hot channel temperatures are assessed relative to the coolant saturation temperature to assure that boiling does not occur. This section describes the results of sensitivity studies, which have been conducted during the past several years, of the hot channel temperature prediction sensitivity to various key parameters. Inter-and intra-assembly flow and heat redistribution effects which lower the hot rod temperature are not discussed in this section since these effects were conservatively neglected in the current design basis calculations; however, these phenomena are discussed in Appendix C. l One of the early studies of shutdown to natural circulation operation af ter sustained prior operation at full power (with the homogeneous core), demonstrated the sensitivity of the core thermal response to the rate of flow 43 1 h
coastdown of the primary pumps. Figures 4-1 and 4-2 show the effect of various rates of flow decay on the cladding temperature response of the hot rod in the fuel and blanket assemblies. FORE-2M (Reference 10) was used to calculate these temperatures using primary loop flow rates calculated with the plant simulation code DEMO. Dynamic flow and heat redistribution effects were neglected. The significant difference between the temperature response of the two types of hot rods is due to the larger stored heat and decay heat of the radial blanket rod. The bigger blanket rod has a time constant about a factor of 9 larger and about a 30% larger decay heat generation than the fuel rod. This latter effect is due to a relatively larger fraction of its operating power being from transuranium (2390, Np) isotope decay energy release. For the longer flow coastdowns, more stored energy is released f"om the rods during the higher flow portion of the transient, and thus, less heat is available later in the transient when the maximum temperature is attained. The CRBRP flow coastdown is designed for a moderate value in the range of Cases B and C. Recent studies at ARD show that fuel type, rod diameter and linear power rating (Reference 11) are very significant since, as with flow coastdown, rod inertial effects influence the transient temperatures. The effect of fuel-cladding gap conductance and shutdown power generation on core natural convection performance was also studied (Reference 16). Figure 4-3 shows the variation in performance for the fuel and blanket hot rods with gap conductance. The large diameter blanket rods have large stored heat effects, and thus, their performance is a strong function of the fuel-cladding gap conductance, whereas the fuel, which has a relatively small stored heat, exhibits a low sensitivity. Figure 4-4 shows the effect of control system shutdown worth. The curve shows that the fuel hot rod temperatures become less sensitive to shutdown worth beyond 46. Early in the shutdown, the power source is from both delayed neutrons and decay heat, and it is the delayed neutron portion that is affected by the scram worth. Af ter the first several hundred seconds following shutdown, the delayed neutron effect becomes-small and the shutdown power is basically from fission product and transuranium isotope decay. Thus, the effect of shutdown system worth on blanket hot rod temperatures would be less than that for the fuel hot rod. As 44
17M o a 1500 C y / E / 3., nm / / / t J f / rar :O n Ac4 g (0Att00w4 (SE C04015 a J / in Fou n0. n Futt ne. 5 um / / a a M n / u g 3 g E / / c c a i 3 7 o SS = Q \\ s %. // 300 I I I I I O le les 150 200 250 300 Tiut (StCl Fig. 4.1 ilatural circulation temperatures for fuel assembly hot rod. n00 - a 16 00 - N N {1500 -
- p. 7""" "" % %
y / / u / /,* ~~\\ 3 / / T / ,5,, 10 00 q/ / / 3 / i / / .s / l 000 eg / \\ / 1 \\ / \\ / n00 \\/ I I TWf 70 alACM ttf C04013 e C0ASf00w4 IM Futt FL Ow 1% Ftit t FL OW 3I00 - / 8 m a A N 33 e a m 0 to a 8 u ist tu 2s0 300 Tiet (StC1 Fig. 4.2 Natural circulation temperatures for radial 45 blanket assembly hot rod.
300 W [ 200 W wA..r... E a a I I FutL Nat R00 e-i i e I. E I e p p r e s 5 3 = 5 -see 3 19 80 98 TOTAL SCRAh4 WORTH ($1 Pf f.Cthf AGE Of N0tn4AL CAlt GAP C090VCTAm** v 4 tut Fig. 4.4 Effect of shutdown system worth on Fig. 4.3 Effect of gap conductance variation fuel hot rod maximum temperatures on the maximum cladding temperature for three-loop natural circulation for a natural circulation event. transient. 46
1 would be expected, the thermal performance is also a very strong function of shutdown power. Core temperatures would vary approximately 8*F for each one l percent variation in the power. j Many localized phenomena influence the core hot rod transient performance. These include fuel restructuring, dynamic gap conductance variations, localized cladding hot spots due to rod spacer arrangement, uncertainties, etc. All of these ef fects are modeled with the FORE-2M code. An example of how these phenomena affect the initial operating conditions of the rods is demonstrated by Figure 4-5 where the fuel temperature is shown with and without the fuel restructuring effect being considered. It can be seen that the maximum fuel temperature is reduced over 300*F by restructuring. This is due to the increased density of the fuel in the hotter inner region of the fuel, and subsequent development of a central void in the center of the pellets for regions where high specific power generation exists. Taking credit for the temperature decrease in the fuel under operating conditions, ~ lowers the subsequent stored energy level of the fuel during the transient. Models incorporated into FORE-2M for predicting the restructuring phenomena have been verified against LIFE computer code predictions (Reference 12) which in turn have been verified against experimental data. 4.2.1.2 CONSERVATISMS OF ANALYSIS This section describes the conservatisms used in the core transient nuclear and thermal-hydraulic predictions of Section 4.3. Worst case conditions were selected for the core hot channel analyses based on numerous sensitivity and l parametric studies. These studies and the associated considerations are covered in the following discussion of the conservatisms. Parametric studies were performed with the Reference 10 version of FORE-2M to substantiate worst case conditions for the nuclear power variation calculation. Conservatively, this model used a reduced level of core detail (from the 7 radial by 7 axial node capability of the code) where the total core wide Doppler for the fuel and blanket regions was included in the fuel region. A base case nuclear model was established using all significant factors that could affect the calculation. All other feedbacks which are 47
.1 _... _ ~ ~ _ _ _ _ ( FUEL / CLAD CAP y 5000 : I ' s N \\ \\ UNREsTRUCTURED \\ 4000 n RESTRUCTURED-o; cc 3000 g s I e 3 i RESTRUCTURED < RADIUS e j h } 2000 l, j CENTR AL VOID RADIUS h i' I I I I 1000 0.000 0.025 .0.050 0.075 0.100 0.125 R ADIUS, (IN.) Fig. 4b Typical 0xide Fuel Rod Radial Temperature Pr0 file Showing Effect Of Restructuring (14 kw/ft) a 48
negative (such as fuel expansion, cladding expansion, ccre radial expansion and bowing) were conservatively neglected. Results of tnis base case and the effect of each worst cise parameter are given by Table 4-1. Here the j variation in the neutronic power is shown for each significant parameter. This demonstrates the importance of each condition in establishing the base In addition to these conservative studies, the base case was repeated l case. using the current FORE-2M capability which considers all core regions. With i j the same base case conditions, at 2 seconds into the transient a maximum neutronic power variation of 0.1276 was found as compared to the value of 0.1327 in Table 4-1. Likewise, a case using nominal data was run with this model and a maximum neutronic power variation of 0.1265. This confirms the conservatism of the base case nuclear model defined by Table 4.1 and used in analysis discussed in Section 4.3. i Although the above nuclear base case conditions were used for evaluating the neutronic power variation, they do not necessarily form the worst case condition for the hot rod temperature calculations. For this type analysis, conditions like the quickest flow coastdown provide the worst case. A ] mismatch of conditions (between the neutronic and temperature calculations) was thus conservatively selected to calculate worst case transient hot rod temperatures. With the benefit of the information found through the various sensitivity and parametric studies, hot rod models are constructed which reflect the following l conservatisms: A. All power, flow and localized rod phenomena uncertainties in the core (both direct and statistical) which affect the thermal conditions on the hot channel; i i B. Worst time in core life with regard to temperature levei, nuclear feedback and decay power generation; .p C. Worst case plant flow and inlet temperature conditions; -49
_ _ _ = O TABLE 4.1 PARAMETRIC CASES TO DETERMINE WORST CASE FOR NEUTRONIC POWER VARIATI0ii DURING 4 - THE NATURAL CIRCULATION EVENT i MAXIMUM VARIATION IN NEUTRONIC POWER, CASE CONDITIONS P/P o,n i 1 o Minimum fuel C 0.1327 p (BaseCase) o Longest flow coastdown** o Maximum Doppler o Zero decay heat o Maximum fuel / cladding gap conductance o Zero sodium coolant density feedback ._..__._.2 ~~ o Quickest flow coastdown** 0.1322
- c. _ _ _
o Other conditions same as Case 1 [ = Maximum so.ium coolant density 0.1326 d i 3 o feedback o Other conditions same as Case 1 4 o Maximum decay heai; 0.1322 i o Other conditions same as Case 1 '7 5 o Maximum fuel C 0.1325 r p o Other conditions same as Case 1 ] 6j o Minimum fuel / cladding gap 0.1304 conductance o Other conditions same as. Case 1 j 7 o Minimum Doppler 0, $48 o Other conditions same as Case 1 ] i
- Valua quoted at 2 seconds into transient for comparative purposes.
- Flow coastdowns specified for pump design limits.
t l' I 1 50 l
D. Worst case Plant Protection System (PPS) scram setting and delays; and E. Worst case control system scram worths. For transient evaluations, 3a uncertainties are applied for each source of uncertainty above. Core related uncertainties of Item A are combined by what is termed the " semi-statistical" method. This has been accepted as a conservative approach to core hot rod / channel design analysis since it was originally formulated by Chelemer and Tong (Reference 13) in 1962 for PWRs. Items B, C, D and E are conservatively assumed to occur at the same time in the worst case directions. As a net result, many low probability events (3a) which are not related are assumed to occur simultaneously. Thus the thermal predictions which are calculated provide a very conservative envelope of transient temperatures for core design and safety evaluation purposes. With regard to item A, Tables 4.2 and 4.3 show that twenty different uncertainty parameters are considered for the fuel hot rod; similar uncertainty definitions were formulated for the blanket rods. These factors are typically termed " hot channel / spot factors." A detailed description of these f actors can be found in Chapter 4.0 of the PSAR and in Reference 14. A summary of all the above and other apparent conservatisms are tabulated in Table 4.4. These have all been applied to calculate the conservative FORE-2M hot rod temperature predictions described in Section 4.3. 4.2.1.3 LOCAL HOT R0D MODELING Much of the-local hot rod modeling effort involves establishing worst case (i.e., conservative) conditions to be analyzed. This was performed through sensitivity and parametric studies as described in Sections 4.2.1.1 and 4.2.1.2. Once the conditions to be analyzed were established on a worst case basis, the design transient hot rod evaluations described in Section 4.2.3 were performed with the FORE-2M computer code, which has undergone extensive verification under the LMFBR Natural Circulation Verification Program. 4 51
l Table 4.2 CRBR FUEL ASSEMBLIES R0D TEMPERATURE NUCLEAR UNCERTAINTY FACTORS, WITH AND WITHOUT CONTROL ASSEMBLY INFLUENCE I )- COOLANT HEAT FLUX DIRECT 1.02(I) (*)1.02(1.10)II) Physics Modeling 1.02(2) 1.02(2) Control Rod Banking 54) I4) 1.0 ZPPR-7 Flux Tilt 1.0 STATISTICAL (3o)I ) Nuclear Data 1.07 1.07 1.01(3) 1.01(3) Criticality Fissile Fuel Maldistribution 1.03 1.03 If assembly is influenced by adjacent control rod, replace with: COOLANT HEAT FLUX E H" Peak Power Position" " Top of Ccre" BOL E0L BOL EOL BOL EOL i Adjacent 1.04 1.02 1.03 1.02 1.15 1.15 ( I Far Side 1.01 1.02 .95 1.02 1.30 1.15 [
- 1) Physics Modeling i
s Adjacent 1.04 1.02 1.04 1.02 1.01 1.02
- 2) Control Rod Banking Far Side 1.02 1.02 1.02 1.02 1.01 1.02 Adjacent 1.04 1.04 1.04 1.04 1.0 1.01
- 3) Criticality Far Side 1.01 1.01 1.01 1.01 1.03 1.01 9, 10, 13, 14, 15, 16, 17, 23, 25, 37, 38, 41, 42, 43, 44, 45, 51, 53
- 4) ZPPR-7 Flux Tilt - Assy's.
(0.97 0 BOL, 1.0 0 E0L). Assy's. 8,11,19, 36, 39,' 47, 65, 68,101,104, (0.99 0 BOL,1.0 0 E0L). Assy's. 62, 98, (0.99 @ BOL, 1.0 0 E0L). I Parentr.esized (*)Non-parenthesized value applies at the peak power position (i.e., core midplane). value applies at the core lower /uppcr axial blanket interface except as superseded by note (1). (o) Engineering Uncertainty Factors are given on Table 4.3. 9
TABLE 4.3 t CR3R FUEL ASSEMBLIES R0D TEMPERATURE ENGINEERING UNCERTAINTY FACTORS COOLANT FILM CLADDING GAP FUEL HEAT FLUX IO) DIRECT Power Level Measurement and Control System Dead Band 1.03 1.03 Inlet Flow Maldistribution 1.05 Flow Distribution' Calculational . -----------l.022 Uncertainty'(Simulation Bias) 1.03 Cladding Circumferential Tem-1.0(f) 1.7(*) 1.0(,)
- perature Variation STATISTICAL ~(3a)(U)
Wire Wrap Orientation 1.01 Subchannel Flow Area 1.028 1.0 Film Heat Transfer Coefficient 1.12 i El Pellet-Cladding Eccentricity 1.15 1.15 i Cladding Thickness and Con-1.12 ductivity 1.48(*) Gap Conductance 1.10 Fuel Conductivity 7 Coolant Properties 1.01 i flow Di'stribution Calculational Uncertainty- (Calibration) 1.054----------l.015 (*) For cladding midwall temperature calculations. Applies to the nominal temperature drop between cladding midwall and bulk coolant. (t) For fuel temperature calculations. -i -(4) Applies to BOL c:nditions. ~ ~ ' ' ~ ~ ~ ~ i '(o) Nuclear Uncertainty Factors are given on Table 4.2. r j, k s s 1 - ~..
IABLE 4.4
SUMMARY
OF ASSUMPTIONS USED IN TRANSIENT 1101 R0D ANALYSIS o Conservative plant initial conditions (see Section 3.2.2.3). o Worst case doppler coefficient including uncertainties. o Minimum control rod shutdown worth (one stuck rod) of $5.82. o 3a' hot channel factors (maximum cladding temperatures are those at the cladding inner diameter under the wire wrap). l 0 Highest power an1. temperature hot rods at worst time in life. o Worst end of uncertainty range used for properties and fuel / clad gap conductance for both the power calculation and the temperature calculation. o Maximum decay heat loads including 3a uncertainties and time in life effects.
- o. 0.2 second delay af ter trip signal before rod insertion begins, o No credit taken for inter-and intra-assembly flow and heat redistribution.
o All above uncertainties assumed to occur simultaneously. I I 54
FORE-2M is a coupled thermal-hydraulics / modified point-kinetics digital computer code which calculates significant reactor core parameters under steady state conditions or as functions of time during transients. Variable inlet coolant flow rates and temperatures are considered. The code calculates the reactor power, the individual reactivity feedbacks, and the temperature of coolant, cladding, fuel, structure and additional material for up to seven axial positions in the core. Various plant protection system trip function can be simulated, and the control rod shutdown worth prescribed as a function nf time after the trip signal. By specifying appropriate hot channel / spot f actors, the transient behavior of an average, peak and hot rod can be analyzed. The heat of fusion accompanying fuel melting and the spatial / time variation of the fuel-cladding gap coefficient (due to changes in gap size) can be considered. The feedback reactivity includes contributions due to the Doppler ef fect, coolant density changes and dimensional changes (including bowing and radial expansion). A detailed description of the code, the equations programmed and modeling features can be found in References 10 and 15. The FORE-2M thermal-hydraulic model for the CRBRP fuel and blanket rods uses seven axial nodes. As maqy as ten radial nodes can be used in the fuel; however, seven have been typically used with no loss of calculational accuracy. Average cladding temperature plus inner and outer cladding temperatures are calculated f or each of the seven axial sections. Likewise, a coolant temperature is calculated for the average and the outlet of each axial section. Studies have been performed using up to 21 axial nodes and essentially no change was found in the accuracy of the calculation relative to that for a 7 axial node evaluation. FORE-2M neglects axial heat conduction in the fuel, cladding and coolant based on studies which indicate that this assumption has insignificant effect on the temperature predictions. An under-the-wire cladding temperature calculation is made for the hot rod at the maximum temperature position. 55 l 1
4.2.2_ DATA FOR HOT P^D ANALYSIS 4.2.2.1 CORE INITIAL CONDITIONS Figure 4-6 shows a 60* symmetry sector of the CRBRP heterogeneous core and the numbering scheme for the assemblies. Details of the lifetime variation in the power, flow orificing, and steady state temperatures of each assembly can be found in PSAR Chapter 4.4. The maximum temperature hot rod (during a natural circulation transient) for each core region (i.e., fuel, inner blanket and radial blanket) are found in the following assemblies with the tabulated maximum initial power (including 3a uncertainties), time in life when maximum power exists, and coolant flow: Peak Rod Max. Linear Assembly Time in Power Flow Region Assembly Life (Kw/FT) (Ib/hr) Fuel 52 BOC) 12.6 151,088 Inner Blanket 99 E0C4 18.4 87,447 Radial Blanket-203 E0C4 12.9 47,569 The hottest rods in these worst case assemblies were analyzed with the conservatisms described in Section 4.2.1, to predict the results described in -Section 4.3. 4.2.2.2 PLANT INPUT DATA Two parameters were used f rom the DEMO plant systems analysis code.in the FORE-2M hot rod calculation. These are the transient variation in core inlet temperature and in reactor inlet flow rate. The conservative basis for these - FORE-2M input parameters is described in Section 3.0. The maximum design value including uncertainties is conservatively used for the inlet temperature. .56
( Jo, \\ /,,7\\ ' a'ti:Sli";t"a'"ne'a',!'s" t 3 j y y s-un c.o m\\ /"\\ /,,,\\ \\ /m\\ / 3,, \\ 3., ( \\ " k " / " /,, \\ O I " / / /,' O/ \\'"/ \\ /,, ,,\\ /,,\\,,,/,,,\\,,,/,,3 \\ (h,, ,,, / / \\' ") \\ \\,, /,o, \\ /m\\ ' 98 3C 68 ,3 C/A 87 76 202 32 ,00 2, C/A 34 6, C/A M 3
- 62' 30 59 4
29 ,28
- FUEL / INNER BLANKET ALTERNATIflG POSITIONS s1 [
0 ASSEMBLY CONTAINING MAXIMUM CLADDINr TEMP DURING NATURAL CIRCULATION TRANSIENT Fig. 4.6 CRBRP Heterogeneous Core - 60 Synnetry Sector and Asseriblies Nunbering Scheme %6N 68 57
f 4.2.2.3 DECAY HEAT Decay heat predictions include contributions f rom fission product, transuranic, actinide and neutron activation isotope decay energy release. Isotopic fission and capture rates in the CRBRP fuel, inner blanket and radial blanket assemblies are used in the prediction of the reactor system decay heat. Verification of the data for fission product isotope decay energy 239 235 238 release due to fast neutron fission in Pu ,U , and U has been provided by detailed comparison to experimental measurement data provided by the DOE sponsored national program to develop the Evaluated Nuclear Data File (ENDF/B). The transuranium isotope decay energy release is provided by decay data from Nuclear Dats Sheets or Table of Isotopes and exists as reference data. Verification of isotopic neutron fission rates and neutron capture rates is provided through an extensive critical experiment program supporting CRBRP nuclear design and conducted at ANL. Time dependent 3a uncertainties associated with fuel and blanket assemblies are utilized which are dependent The on the type of assembly, i.e., fuel, inner blanket, or radial blanket. decay heat uncertainties from scram to two (2) hours after scram are summarized as follows: (a) Scram (b) 2 minutes (c)2 hours Fuel A3sembly 32.4% 13.7% 10.3% Inner Blanket Assembly 28.6% 9.1% 5.6% Radial Blanket Assembly 27.6% 10.1% 7.2% These values are for end-of-equilibrium cycle conditions at rated power operation. These decay heats are conservatively combined with beginning of cycle power generation for fuel hot rod predictions. The total shutdown power generated in the respective hot rods is due to this decay heat source plus the delayed neutron power generation. Normalizing the transient power from both sources to the total initial operating power of the hot rod gives the relationship S8 ) ~
P I'} = P (t) ~ j,P(} D(*} T g D + b P P o o g g where: initial total operating power of rod P = g P (t) transient total rod power. = T P (t) transient neutronic power. = g P (t) tr nsient decay heat power. = D Power due to decay heat sourt.es at full power. P (o) = 0 which is applied in FORE-2M for the transient hot rod power variation. Details of the conservative calculation of the neutronic power component, P (t), are N described in Section 4.?.l. 4.3 RESULTS l l 4.3.1 FUEL, INNER BLANKET AND RADIAL BLANKET HOT R0D TEMPERATURES Figures 4-7 to 4-9 show both the 3o and nominal cladding temperatures attained for the limiting heterogeneous core design hot rods. These are the l highest core temperatures at the hottest location of the hottest rod for the fuel, inner blanket and radial blanket assemblies. The maximum temperatures reached and the time they are reached for each limiting hot rod type are tabulated below: l Maximum Temperature (*F) versus Time of Occurrence (sec) Nominal 3a Assembly Temp Time Temp Time F/A-52 1299 178 1565 180 1/B-99 1229 222 1544 239 R/B-203 1279 275 1556 289 59
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rea-E 75 3,12 the ame ame IMe I R = im8 30 in. p g "' fi. ,/ 's.,,, E '^'---.: E" (fioraina1 isme M B ,_ me I me me se me.e me me me.e sme TI8E IMC.I FIGl'P.E 4-8: TEliPERATURE RESP 0flSE OF HIGIEST TEllTERATURE R0D Ill IBA-99 F0?. CRBRP f1ATURAL CIRCULATIGU C00LIf1G TRAllSIEllT 4 4&> w { % '$ e 9 - a *** b
FFME e6 3.12 PEaPV-110L8 4 tme tsuke t l I -two 3a g I e.0 13D m 'sa " %.~~ ew g f s,,, ggno ( r:om na-- a N ~ A l g N.*#,.- o g gissa N Ee ,_ use.o a.sd B - see.o I 2 ses.o me u us.e am.e me me an.e T!8E IMC.) FIGURE 4-9: T f!PERATURE RESPC:iSE OF ;!1CIEST TE!:PEPJ.TLRE RCD II; RJA-203 FOR CR3R!' ilATURAL CIRCULATIC.. ~ LI!!G TRA:45 ::.!
4.3.2 COMPARIS0N OF HOT R0D TEMPERATURES TO SATURATION TEMPERATURE When the reactor flow decreases, the static coolant pressure, and subsequently, the saturation temperature decreases. In addition to friction, expansion and contraction losses, the reactor cover gas pressure and the hydrostatic head of sodium above the core are,important parameters in determining the coolant pressure. Figure 4-10 shows typical values for CRBRP. As can be noted, temperatures in excess of 1900*F would not result in sodium boiling at 100% flow and even at low flow conditions (approaching zero) temperatures in excess of 1720*F can be sustained without boiling. This latter temperature is due primarily to the hydrostatic head of sodium that exir4s above the core hot spot locations and would apply at low flow conditions to both the fuel and the blanket type assemblies. Even on a 3a temperature basis it can be seen that over 150*F margin to coolant boiling exists and over 400'F margin exists on a nominal basis. 1 63 i s
o d. t 2000 1093 C O 1075 E E E w E" E k 1050 4 5 tg 1900 g d W 1025 H 2 Z O 9 P E 5 1000 g g p 1800 5, 4 2 975 g 5 5 O 8 Oo u 950 I I I I 1700 926 0 20 40 00 80 100 FRACTION OF FULL FLOW (%) Fig. 4.10 Coolant Saturation Temperature at Top of Active Fuel Position as a function of Reactor Flow for a Peak Fuel Assenbly (CRBR) t t 64 6
5.0 SINMARY AND CONCt.USIONS This report presents a summary of the current assessment of the capability of the CRBRP to remove decay heat by natural circulation. It effectively accounts for 1) the core design change from the homogeneous core to the heterogeneous core, 2) various plant system and model changes that have been shown to be important since the first assessment was made in 1976 and 3) updated input data including decay power, core pressure drop and test data f rom the CRBRP prototype pump water tests. The event analyzed is one in which it is assumed that while the plant is operating at rated power, there is a simultaneous loss of power to all motor driven pumps (includiag the pony motors on the sodium pumps) followed by a reactor scram. Forced circulation is provided during the coastdown of the main coolant pumps and thereaf ter, the flow is provided solely from the thermal or natural heads in the systems. This assessment confirms the conclusions of the preliminary evaluation of the natural circulation capability of the plant reported in 1977 (Reference 1). That is, the CRBRP can sustain a loss of power to the main coolant pumps as well as the steam generator recirculation pumps simultaneously with a plant trip with substantial margin to boiling in the core. The analysis indicates that even if no credit is taken for the intra-and inter-assembly flow and heat redistribution within the core, there is a significant margin to boiling (greater than 150*f at the core maximum hot spot). On the basis of these analyses, the CRBRP design provides adequate decay heat I removal capability by natural circulation. 65
6.0 REFERENCES
1. R. R. Lowrie and W. J. Severson, " Clinch River Breeder Reactor Plant, A Preliminary Evaluation of the CRBRP Natural Circulation Decay Heat Removal Capacility," CRBRP-ARD-0132, November 1977. 2. W. H. Alliston, et al., " Clinch River Breeder Reactor Plant, LMFBR Demo Plant Simulation Model (DEMO)," CRBRP-ARD-0005, February 1978. 3. H. P. Planchon and W. R. Laster * " Loop Heat Capacity Models and Their Effects for DEMO Natural Circu1ation Transient Analyses," WARD-NC-3045-2, September 1978. 4. R. M. Singer and J. L. Gillette, " Measurements of Subassembly and Core Temperature Distributions in an LMFBR," AIChE Symp. Ser., 73, No.164, pp. 97-104, (1977). 5. R. L. Stover, et al., "FFTF Natural Circulation Tests," Trans. Amer. Nucl. Soc., 39, pp. 702-703, (1981). 6. F. C. Engel, B. Minushkin, R. J. Atkins and R. A. Markley, " Characterization of Heat Transfer and Temperature Distributions in an Electrically Heated Model of an LMFBR Blanket Assembly," Nucl. Eng. Des., 62, pp. 335-347, (1980). 7. R. H. Morris and W. R. Nelson, " Single-Phase Sodium Tests in a 61-Pin Full Length Simulated LMFBR Fuel Assembly-Record of Phase 1 Experimental Data for THORS Bundle 9," 0RNL/TM-7315, August 1980. 8. A. K. Agrawal, et al., " Dynamic Simulation of LMFBR Plant Under Natural Circulation," ASME Paper 79-HT-6. 9. M. Khatib-Rahbar and K. B. Cady, " Establishment of Buoyancy-Induced Natural Circulation in Loop-Type LMFBRs," Trans. Amer. Nucl. Soc., 28, pp. 432-433, (1978). 10. J. V. Miller and R. D. Coffield, " Clinch River Breeder Reactor Plant, FORE-2M: A Modified Version of the FORE-Il Computer Program for the Analysis of LMFBR Transients," CRBRP-ARD-0142, May 1976. 11. R. D. Coffield, et al., " Generalized Transient Characteristics and Capabilities of Liquid Metal Fast Breeder Reactor Fuel and Blanket Rods," in Proceeding of the ANS/ASME/NRC International Topical Meeting on Nuclear Reactor Thermal-Hydraulics, NUREG/CP-0014. Vol. 3, October 1980, pp. 1581-1596.
- 12. " LIFE-III Fuel Element Performance Code," ERDA-77-56, July 1977.
13. H. Chelemer and L. S. Tong, " Engineering Hot Channel Factors for Open lattice Cores," Nucleonics, y, No. 9, pp. 68-73 (1962). 66
14. A. J. Friedland, "CRBRP Core Assemblies Hot Channel Factors Preliminary Analysis," CRBRP-ARD-0050, February 1980. 15. J. N. Fox, B. E. Lawler and H. R. Butz, " FORE-II: A Computational Program for the Analysis of Steady State and Transient Reactor Performance," GE AP-5273, September 1966. 16. R. D. Cof field, R. A. Markley and E. U. Khan, " Natural Convection Analyses and Verification f or LMFBR Cores," Nuc. Eng. Des., 62, ~~ pp. 181-198, (1980). t \\ 67
APPENDIX A SELECTED PLANT TRANSIENT DATA In addition to the analysis data given in the main body of the report, there are a number of other curves of interest in understanding the plant response to the natural circulation eve,nt. These are provided in this appendix and are as follows: Parameter Figure Primary Pump Speed S-Loop, % A.1 Intennediate Pump Speed S-Loop, % A.2 l lot Leg Temperatures, S-Loop A.3 Superheater Exit and Evaporator Inlet Temperatures, S-Loop A.4 Core Inlet Temperature A.5 j Natural Heads, S-and L-Loops A6 Steam Drum Pressure, S-Loop A.7 Steam Generator System Vent Flows, S-Loop A.8 Feedwater Flow, S-Loop A.9 e t 9 D / 4 \\ A.1 4 )
a b e W u n W k. W by r o it B \\ '. o 't 8 8.g i t .,v.. il' E 8 i, h, I> s-I .o O O o o O a 6_ o m o II l 631 0/0 - 033d8 eMnd led S Figure A.1 A.2 s w,
) k P k ~ E k W H. !E p ~ 3 a I / i, t o o o o R W o d i ICI o/0 - c334 dMfW J.N! 8 Figure A.2 A.3
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APPENDIX B COMPARISON OF RESULTS FOR THE HOMOGENEOUS CORE 1 When the change f rom the homogeneous core to the heterogeneous core was first being considered, analyses were performed in 1977 to assess the impact that the change would have on the natural circulation decay heat removal capability of the CRBRP. The reactor flows and inlet temperatures were calculated using the same models in DEMO that were used in the analysis for the homogeneous i core (Reference B.1) except that the power fractions and core pressure drops were changed. That analysis showed that there was little difference between the hot spot temperatures for the heterogeneous core compared to the homogeneous core. This was achieved through orificing the core at operating conditions to result in similar natural circulation temperatures, i B.1 DATA COMPARISON B.l.1 PLANT DATA Figure B.1 shows a comparison between the core pressure drop correlations used for natural circulation evaluation for the reactor with the homogeneous core (Reference B.1) and that used for the 1977 evaluation with the heterogeneous Core. Figure B.? shows a comparison of the resulting flows. Figure B.3 provides the j decay power used in the analysis. 1 B.l.2 HOT ROD FLOW-AND POWER A comparison of the initial conditions for the limiting homogeneous and heterogeneous core designs is provided in the table below for both fuel and - ' blanket type assemblies: i' Peak Rod Max. Linear Assembly Time in Power (30) Flow l . llomogeneous Life Kw/ft (1b/hr)- l /A #H BOL ' 12.2 -163,880 R/B #A-1 EOL 18.3 58,965 I 8.1
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i Heterogeneous F/A#52 80C1 12.6 151,088 j 1/8 #99 E0C4 18.4 87,447 R/B #203 E0C4 12.9 47,569 4 The hottest rods in these assembles were analyzed with the same methodology l described in Section 4.2.1 for both the homogensuus and heterogeneous core designs. B.1.3 HOT R0D DECAY POWER COMPARISONS The basis of the decay heat applied to the hot rods for the evaluation of the heterogeneous core was similar to that presented in Section 4.2.2.3. The . methodology for predicting the decay heat for the homogeneous core design was similar but instead of. applying time varying uncertainties, these earlier predictions used a constant +25% uncertainty for fission product riecay power 239 and1/39 decay power. Reference B.2 and a +10% uncertainty for Np provides details of the decay heat calculation methodology. A_ comparison between the homogeneous and heterogeneous core decay powers at 100 seconds after shutdown are tabulated below: Decay Power with 3a uncertainties Homogeneous 0100 seconds after shutdown, P/Po F/A #8 0.043 R/B #A-1 0.055 Hetergeneous i F/A #52 0.030 I/B #99 0.043 R/B #203 0.045 B.o
As can be noted, the decay power is 20 to 30% lower using the current decay heat methodology described in Section 4.2.2.3. 4 B.2 TRANSIENT COMPARISONS A comparison of the 3a maximum cladding temperatures calculated for i homogeneous core to those calculated in 1977 for the heterogeneous core are given'in the table below: Homogeneous Heterogeneous Core, (*F) Core,(*F) Maximum Fuel Rod 1577 1575 Maximum Blanket Rod 1580 1599 (Inner Blanket) 1623 (Outer Blanket) The conclusion reached in 1977 was that decay heat removal by natural circulation from the heterogeneous core would meet the acceptance criteria. As discussed in the main body of this report, the effect of changes in the prediction of decay power combined with an updated plant simulation (including a much longer pump coastdown) have resulted in predicted temperatures that are significantly lower. 2 B.3 REFERENCES B.1 R. R. Lowrie and W. J. Severson, " Clinch River Breeder Reactor Plant, A Preliminary Evaluation of the CRBRP Natural Circulation Decay Heat Removal Capability," CRBRP-ARD-0132, November 1977. B.2 C. A.. McGinnis, D. O. Tomlin and R. K. Disney, "CRBRP Decay Power Analysis," WARD-0-0090, January 1976. d: d I i B.6 -
APPENDIX C FACTORS AFFECTING THE NATURAL CIRCULATION RESULT 5 C.1 PRINCIPAL DiffERENLES BETWEEN THE PRELIMINARY EVALUATION AND THE CURRENT ASSESSMENT The current evaluation of the natural circulation capability of the CRBRP, when compared to the preliminary evaluation reported in Reference 1, has shown that the maximum hot rod temperatures are somewhat lower in spite of the fact that pressure drops used in the current analysis are higher than those used I before and average fuel assembly powers (at rated power) are higher. There are several reasons for this. The principal ones are as follows: i C.I.1 PUMP COASTDOWN Figure C.1 compares the pump speed versus time for the current analysis (heterogeneous core) with the earlier analysis which used the pump i characteristics given in Reference 1. The preliminary analysis used a very conservative friction torque. For example, at 20% speed the sum of the pumping and f riction torques (assuming a U/K = 1) is 0.04708 x the reference torque (T ) and the friction torque is only 11% of the total torque at that R speed. As stated earlier in the report, these values are based on data from the prototype pump water tests. The corresponding values for torque at 20% speed (with U/N = 1) for the earlier analysis gives a total torque of 0.0565 xT and the f riction torque is 29.4% of the total torque. Thus, the g coastdown is significantly affected and the predicted time to stop is 2.36-times as long as in the earlier analysis. This longer coastdown significantly affects the flow history f rom about 20 seconds af ter scram until the time the pump stops and the loop pressure drop increases due to the addition of the pump stopped rotor. resistance. C.I.2 DECAY POWER The current. values of decay power are lower than those used in the preliminary evaluation conducted in 1976. The maximum total decay power at t = 10 seconds dfler SCrdm is 5.5% rather than 6.1% used in the earlier analysis. The C.1
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corresponding values at t = 100 seconds are 3.65% versus 4.15%. It should be noted that in the earlier analysis (Ref erence 1) the maximum total decay power was 4.f;% when the pump stopped (at t = 55 seconds). With the updated decay power and revised pump coastdown, the decay power is 3.45% or a reduction of 25% in decay power at the time the pump stops. Therefore, the flow required to maintain acceptable core temperatures is redui.ed. C.I.3 STEAM GENERATOR /IHTS INTERACTION The evaluation conducted in 1976 showed a severe down transient at the l evaporator water inlet and consequently the sodium exit. This was the caused by a model which ef fectively introduced the 70 F auxiliary feedwater directly into the system at the drum exit. The total recirculation flow was a mix, then, of saturated water (trom the drum) and rather cold f eedwater. The effect of the down transient was to suppress the thermal head since the colder I sodium f lows upward f rom the evaporator outlet to the intermediate pump located at the high point in the system. Tne current analysis simulates the { change in the location and design of the auxiliary feedwater sparger. It is located in the steam space of the drum and equipped with spray nozzles to assure droplet sizes that can be effectively heated to saturation temperature bef ore f alling to the f ree surf ace in the drum. The effect is that down transient at the drum exit (and thus the evaporator inlet) will be eliminated and the drop in sodium temperature at the evaporator outlet is largely i mitigated. When combined with the longer coastdown of the intermediate pumps, this mitigation in evaporator outlet temperature results in a smoother reduction in intermediate flow which in turn affects the temperature distribution within the IHX. C.2 INTRA-AND INTER-ASSEMBLY FLOW AND HEAT REulSTRIBUTION EFFECTS As described in Section 4.1 core inter-and intra-assembly flow and heat redistribution under-natural circulation conditions a a known phenomena which significantly lower maximum core temperatures f rom those calculated without considering the effects. C. 3
The COBRA-WC code (Ref erence C.1) has been developed and verif ied to account for core inter-and intra-assembly flow and heat redistribution. It predicts the boundary conditions for a peak rod or cluster of rods in core assemblies given the reactor boundary conditions such as: a) total reactor flow, pressure drop and core inlet temperature from DEM0; b) individual core assembly powers; and c) individual core assembly thermal-hydraulic characteristics. Analyses of this type phenomena under natural convection cooling requires detailed core / reactor modeling because of the strong interaction between the fuel assemblies, blanket assemblies, control assemblies, plus other core regions and bypass flows which all act as highly coupled parallel flow paths with heat transfer between them, 4 A quantification of the affect of the flow and heat redistribution was made in connection with FFTF natural circulation acceptance testing using a methodology involving a system of computer codes (DEMO, COBRA-WC, FORE-2M). These studies were made with low decay power and consequently low core transient temperatures. As reported in Reference C-2, the following three cases were analyzed
- A.
Case 1 - Fuel assembly peak coolant channel transient flow and temperature calculations including both inter-and intra-assembly flow and heat redistribution under nominal conditions (no uncertainty factors applied); B.- Case 2 - Fuel assembly peak coolant channel transient flow and temperature calculations including both inter-and intra-assembly flow and heat redistribution for 3a uncertainty f actor conditions; C. Case 3 - Fuel assembly transient flow and temperature calculations i ' fwithout inter-assembly and intra-assembly flow and heat redistribution; 3a uncertainty f actor conditions. C.4
Typical results for these three cases are presented in Figure C-2 as normalized temperature differences relative to the steady state temperature difference in Case 1, i.e., U )~ I3 i = Case 1, 2 or 3 ,= c in 1 IUl Li lo) l in c where: T (t) = maximum hot channel coolant temperature; T (t) =' inlet temperature. As can be noted by comparing Cases 2 and 3, accounting for inter-and intra-assembly flow and heat redistribution effects significantly decreases the predicted transient coolant temperatures in the hot channel (i.e., by ~12%). It can also be seen that the uncertainty factors provide a significant conservatism in the predicted hot channel coolant temperatures (.i.e., Case 1 versus Case 2). The mitigating effect of the heat and flow redistribution would be substantially larger for the significantly more severe conditions (mainly end-of-cycle decay heat loads) and resultant localized core temperature peaking described in Section 4.3. It is anticipated that the temperature reduction for the hot fuel rods will be in the order of 100*F and that a similar 100-to 180*F decrease will be found for the hottest blanket rods. i Recently, f ull plant tests were conducted to determine the Fast Flux Test Facility natural circulation capability. Pretest predictions, reported in References C.3, C.4, and C.5 were made with 1) design data and methods as used in this report and 2) with a system of codes utilizing nominal data. - A comparison of' nominal predictions from the codes to the exrerimental data is shown by Figure C.3. Here it can be seen'that, even on a nominal basis, good 4 agreement was obtained with the predictions being somewhat cons _ervative in r.omparison to the data. Figure C-3 shows predictions which include i uncertainty f ar. tors would'be over 300 F hotter than the nominal prediction and measured data. When uncertainty factors are applied and i i C.5 { 1
inter / intra-assembly flow and heat redistribution are neglected, as in the approach presented in Section 4.3, the system of code predictions are extremely conservative relative to anticipated temperatures f or natural circulation cooling. C.3 REFERENCES C.1 T. L. Geory at al., " COBRA-WC: A Version of COBRA for Single-Phase Multiassem,y Thermal Hydraulic Transient Analysis," PNI.-3259, July 1980. C.2 R. D. Coffield, et al., " Buoyancy induced Flow and Heat Redistribution During LMFBR Core Decay Heat Removal," Proceedings of Specialists Meeting on Decay Heat Removal and Natural Convection in FBRs, Brookhaven National Laboratory, NY, February 1980. C.3 Letter, R. L. Copeland to D. G. Eisenhut, " Pre-Test Predictions of FFTF Natural Circulation," Docket No. 50-337, november 28, 1980. C.4 Letter, R. L. Copeland to D. G. Eisenhut, " Pre-Test Predictions of FFTF Natural Circulation Tests," Docket No. 50-337, February 26, 1981. C.5 R. L. Stover, et al., "FFTF Natural Circulation Tests," Trans. Amer._ Nucl. Soc., 39, pp. 702 and 703, (1981). C.6
J ki s 1.8 CASE 3 ],-. 1.G / CASE 2 su / "f 1 A / y at
- E CASE 1 3 E I2
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,/ -.' s W>4 0.8 e .J ) / 5O A:n > 1 'T (j G Ir,tes an<t Iniva Auernhty $ Of1F 7M R. fle 1.st s l..stion, f Jeaniinal COajl( A VrC i wG Case 2- .= DA enin e t Inc.a Anan.hty D (: finfide elns t ensi, 3 4 3 IIC I j O 4 (' a se 3 : fletfid e'"nb ut oon; 3 f) HC F"N l ' (' !" 4 N ' " ' ' " ' ' " " ' ^ " " 0 *2 b 0.0 I I I i 1 I I L1 0 40 00 120 160 200 TIME (SECONDS) (.486 1 Fig. C.2 Typical Effect of Uncertainties and Inter-/ Intra-Assembly Flow / Heat Redistribution on Coolant Temperatures Duri g Natural Convection Cooling C.7 7
1400 - Y 1300 1200;p 1100 2 9 o di s x = 4,s ., C ~ ' ~ h ? 25 II. W1000k. ,e - h l' 5 li m b O / /h 900{'., j ill. { '. ,A, l 1i..* y,,,, Pre-test predictions without inter-and intra-asembly 800f-)' flow and heat redistribution; 30 uncertainties Pre-test prediction with inter-and intra-C assembly flow and heat redistribution; nominal conditions s O - tieasured data (Reference C.5) 700-Post-test prediction with inter-and intra-A---- assembly flow and heat redistribution; nominal conditions i I-- l 1--- I - - l- - I I - I 600 O 100 200 TIME (SEC.) FIGURE C.3: MEASURED ATID PREDICTED SODIUM TEMPERATURES AT TOP 0F THE FUEL SECTI0il, TX1016, FOR R0W 2 F0TA (TEST INITIATED FROM 100% POWER /100% FLOW) C.b ~}}