ML20069C757
| ML20069C757 | |
| Person / Time | |
|---|---|
| Site: | Clinch River |
| Issue date: | 07/15/1982 |
| From: | Longenecker J ENERGY, DEPT. OF |
| To: | Check P Office of Nuclear Reactor Regulation |
| References | |
| HQ:S:82:070, HQ:S:82:70, NUDOCS 8207210215 | |
| Download: ML20069C757 (31) | |
Text
-
Department of Energy Washington, D.C. 20545 Docket No. 50-537 HQ:S:82:070 JUL 151982 Mr. Paul S. Check, Director CRBR Program Office Office of Nuclear Reactor Regulation U.S. Nuclear Regualtory Commission Washington, D.C. 20555
Dear Mr. Check:
RESPONSES TO REQUEST FOR ADDITIONAL DOCUMENTATION
Reference:
RAPIFAX from T. King, " Documentation Desired as a Result of June 22, 1982 Meeting on CRBR PSAR Chapter 4.4," dated June 24, 1982.
This letter transmits PSAR Chapter 4 pages which have been modified and provide the information requested in the referenced RAPIFAX (Enclosed). These pages will be incorporated into Amendment 69 to the PSAR, scheduled for submittal in July.
Sincerely, C,s 3.
A J n R. Longen ker Acting Director, Office of the Clinch River Breeder Reactor Plant Project Office of Nuclear Energy Enclosures cc: Service List Standard Distribution Licensing Distribution g(
$2ggoo 050 A
_______.___._j
r
.,$2250'j;*ie' h A n i k - _ _ _
" ' ' ~
is j,.; A n e e %,,
c10)
. c) <Ie1
,k.. -.
D_0CUMEf]TATION DESIRED AS,,A._ RL5till 01 6/22/82 MFFTING ON CRBR PSAR CHAPTER 4.4
- 1) Modify PSAR or provide a separate letter clarifying the Projects plans t
~
- \\
on 2-loop operation (f.c.. not planned for first core operation and A.-
not included as part of first core operating license request).
t t
?)
In paragraphs 8 and 10 of Section 4.4.1 of the PSAR, reference should be made to Section 4.2.2.l.3 of the PSAR for the conditions which must he met in orificing flow to the reactor vessel internals.
a 3)
At the end of the 3rd paragraph of Section 4.4.3.2.1 add 4 statement that
+
the uncertainties and confidence levels of the hot channel facters and the affccts of a non-IIncar application of the hat channel factors will be evaluated in the FSAR.
l i
4)
Document the margin provided to the 30-year lifetime components regarding their steady st. ate design temperature (i.e.. What is their predicted l
temperature, including uncertainty. versus their design temperature?).
t 51 ) for core replaceable components document rational as to why PEOC +2C l
temperatures are used for steady state and anticipated transient analysis whercar. THDV 4 39~ temperatures are used for unlikely and extremely 1
unlikely event analysis.
This should address whether or not cladding failure is considered a safety issue.
i.
l 6)
Document the maximum'.flavi thru primary control assemblies (essuming
~
primary pumps are at their maximum speed) and the required tiow for I
floatation.
1 O
i i
L i
l q'
.<25D 15 8 37 m *
- = - - "
i ;
i I
2
~
7).. Reference experimental data used for flow distribution c.alculations f'
i p ;)
~~
at low flow (Section 4.4.7.5).
i 7
l l
- 8) Change paragraphs 0,6 and 7 of Section 4.4.1 '.o state that no melting is allowed (i.n., not just no centerline melting).
I
- 9) Provide a clarification on the Project position on fuc1 lifetim'e j
(i.e., operating license is only for 80,000 MWD /MT burnup).
- 10) Clarify khat parameters were used as guidelines and what parameters were used as design limits in developing the reactor vessel ctxnponent
' flow allocations.
Address such ' item, as SELT, DELT, TELT. Im,1550 F transient temperature limit., no boiling 1,imit, 900 F vessel temperature.
l i
etc.
l i
j-yx,9 -(.. ! -
l n.
a e
l s
i t
'4 O
e l
I
c) Recutrement - Transmit the applied loads f rom the reactor core assemblies and the Upper internals Structure to the Core Barrel including upward vertical loads.
Bases - Transfer loads to the Core Support Structure, the primary structural support and positional ref erence for the Core Former Structure.
l l
d) Enquirrment - Provide a structural attachment for the Fixed Radial l
Shield.
Bases - Maintain lateral restraint at the upper end of the Fixed Radial Shield.
e) Reautrement - Provide a temporary vertical support for the Upper.
Internals Structure.
Bases - React the dead weight of the Upper Internals Structure during Installation of reactor components.
4.2.2.1.3 Design Loading The loading conditions to which the reactor internal structures may be subjected are categorized into Normal, Upset, Emergency, Faulted, and Design Conditions as defined in Section ill NG & NB-3000 of the ASME Boiler and Pressure Vessel Code.
Table 4.2-21 provides for the 30-year life reactor internals components the design temperatures versus the predicted steady-state temperatures (including uncertaintles) at the maximum temperature point of the components.
Design loading conditions are given for the two principal groups of ' reactor internals components, the upper internal structure and the lower internals structure, The only structural component of the lower internals structures is the core support structure.
Thus the temperature, pressure and static loads for the lower internals which follow are. stated for the core support structure.
4.2-153a Amend. 69
. July 1982
TABLE 4.2-21
~
DESIGN TEMPERATURES VS PREDICTED STEADY-STATE TEMPERATURES FOR PERMANENT REACTOR INTERNALS COMPONENTS Design Predicted'Maximug)
Minimum Component Temperature S.S. Tem erature Margin
( F)
( F)
( F)
Core Support Structure
)
Core Plate 775 750 25 Module Liner 775 750(2) 25 Core Barrel 1060 1010 50 Bypass Flow Module 775 750(2) 25 Fixed Radial Shield 950 932 18 Horizontal Baf fle Assy.
FT&SA Support Block 775 750(2) 25 HBA Base Piate 1020 1015(2) 5 Core Former Structure Lower Ring 928 928 Cylinder 937 937 Upper Ring 1076 1076 Lowee inlet Module 775 755 20 Upper internals Structure 1220 1191
,29 Notes:
(1) All values shown include a 25 uncertainty.
(2) Coolant temperature.
Actual component temperature siIghtly lower.
I l
t 4.2-365
~
Amend. 69
~
July 1982
W-WWF lDdMJTLAM 4
No f uel melting is allowed in the fuel assemblies at 115% overpower conditions (*),
including dasign and experimental uncertainties at 3cr conf idence l evel.
Consequently, the linear power rating will not exceed the limiting power-to-melt under the aforementioned conditions.
l 5.
No f uel melting is allowed in the. blanket assemblies at 115% overpower condi ti ons(* ), including design and experimental uncertainties at 3o-confidence l evel.
The blanket management scheme will theref ore be arranged not to exceed the limiting power-to-melt under the aforementioned conditions.
l 6.
No absorber melting is allowed in the control assemblies at 115%
overpower conditions (*), including design and experimental uncertainties at 3c' confidence level.
7.
The sodlum temperature exiting the core assemblies will be consistent with the limitations reported in Section 4.2.2.1.3.2 to assure the structural Integrity of the upper internals structure during its prescribed lif etime.
l 8.
Mixing in the inlet and outlet plena will mitigate the ef fects of thermal transients on the internal structures, such that the components structural requi rements are met.
9.
Adequate cooling shall be provided to the shielding, core barrel and core f ormer components to yield a thermal environment capable of assuring their structural Integrity as speci fied in Section 4.2.2.1.3.
Suf ficient flow shall be provided to the reactor vessel thermal liner to limit the vessel wall temperature below 9000F during normal operation.
Adequate cooling shall be provided f or the Fuel Transfer and Storage Assembly to preserve the structural Integrity of stored fuel assemblies.
l
- 10. Adequate heat removal by forced and free convection from heat producing reactor components shall be assured for all operating condi tions.
l
- 11. During operating conditions, fuel, blanket and control assemblies total. pressure drop along with the rest of the primary system pressure drop will be within the primary pump head capability at the corresponding flow.
- 12. Coolant velocities shall be less (unless test data support higher acceptable velocities) than the following limits dictated by cavitation and/or corrosion / erosion considerations: 30 ft/sec f or non-replaceable components; 40 f t/sec f or replaceable components in the high coolant temperature region (exit); 50 f t/sec for replaceable components in the low coolant temperature region (Inlet).
t
(*)This definition means a power equal to 115% of rated power conditions.
4.4-2 Amend. 69
. July 1982
l
- 13. The control assemblies flow rate will be such as to assure adequate margin against flotation in case the driveline becomes accidentally di sconnected (see Section 4.2.3.1.3).
l 14.
Assemblies orificing will be designed to be consistent with the requirement that the lower shield in the f uel, blanket and control assembliss will have suf ficient solid volume fraction to limit radiation damage to the core support structure and to assure its prescribed lifetime.
l
- 15. The thermal-hydraulic design of the control assemblies will be such as to satisfy the scram insertion requirements during the reactor l if etime (see Section 4.2.3.1.3).
l
- 16. The sodium temperature shall be less than its boiling point during normal operation and anticipated and unlikely transient conditions.
l
- 17. The reactor will meet the aforementioned design bases operating over a range of power and flow rates, including power ranges and flow l,
variati ons, from 0 to 100% of nominal conditions.
l 18.
Adequate design margins (see Section 4.4.3.2) will be provided to account f or design, f abrication, operational uncertainties and tolerances to ensure meeting the aforementioned limitations.
The semi-statistical hot channel factors approach will be adopted in combining Individual fuel, blanket and control assembly uncertai nti es.
l 19.
As explained in Section 4.4.3.3.1, plant T&H design conditions are considered in perf ormance evaluations of permanent plant com-ponents(+),
e.g., vessel, internals, heat exchangers.
Therefore, these conditions shall be considered in evaluation of Itans 7 through 10,16 through 18.
On the other hand, plant expected operating conditions are adopted in steady state perf ormance and design evaluations of replaceable components.such as the reactor assemblies.
Therefore, plant expected operating conditions sh'all be considered in l
evaluation of items 1 through 6,11 through 15,17 and 18.
l l
~
4.4-3
- - - ~ - -
Amend._69 July 1982 O O E M E
I
W enrig@%QrteP>
r 4.4.2 Descriotion 4.4.2.1 Summarv Comoarlson This section presents a comparison of general and core assemblies design parameters f or the CRBRP and FFTF reactors.
l.
CRBRP AND FFTF GENERAL PARAMETER COMPARlSON Units CRBRP**
EEIE*
Design Lif e Yrs.
30 20 Reactor Power (Thermal)
MWt 975 400 Primary Coolant Sodium Sodium Primary Coolant Design Flow Rate 106 lbm/hr 41.45 17.28 Cool' ant Temperature:
Reactor Vessel Inlet OF 730 600 Reactor Vessel Outlet OF 995 858 Reactor Vessel Temperature Rise OF 265 2 58 Pressure Drop:
Reactor inlet Nozzle-to-Outlet Nozzle (4) psi 123 110 Lower inlet Module to Assembly Outlet Nozzle psi 116 101 Primary Pump Design (static) psi 160.3 182.5 Number of Primary Loops 3
3 Suppressor Plate Yes Yes Cover Gas Argon Argon Cover Gas Pressure (nominal) psig 0.36 0.36 Allowable Overpower percent 15 15
(+) Permanent plant components are those components which: ' 1) will b'e designed f or 30-year lif e; and 2) cannot be easily replaced.
- FFTF Initial Condition
- CRBRP T&H Design Val ues CRBRP value includes uncertaintles; FFTF value is nomin'al..
l l
t l
4.4-4 i
I Amend. 69
. July 1982
iuse 4 LuA=Usu)J LU,u4J #5)
_Outfet Plenum All f uel, blanket, control, and a portion of the radial shield assembly flow discharges into the upper internals structure.
The coolant first enters a mixing chamber bef ore entering the chimneys (Figure 4.4-8).
The chimneys duct the flow vertically upward and discharge the flow into the upper region of the vessel outlet plenum. The flow is directed into the upper region of the plenum to minimize flow stratification in this region during a reactor trip transient.
The flow from some of the removable radial shields which are located outside of the peripheral skirt of the upper internal structure discharge directly into the outlet plenum.
- Also, 14% of total reactor flow from the f uel, blanket, control and radial shield assemblies bypasses the chimneys through the gap between the top of the core assemblies and the skirt of the upper Internals structure and discharges directly into the outlet plenum.
The coolant leaves the reactor vessel outlet plenum through three 36-Inch dinneter outlet nozzles.
4.4.2.5 Fuel and Blanket Assemblies Orlficing 4.4.2.5.1 OrfficInc Philosochv. Acoroach and Constraints Core ori f icing, i.e., flow allocation to the various f uel and blanket assemblies is an important step in the core thermal-hydraulic design.
Since the assembly temperatures are directly dependent on the amount of flow and since the flow allocation is the only thermal-hydraulic design parameter which can be varied, within certain limits, by the designer, it logically follows that the core T&H design and perf ormance is only as " good" as the core orificing.
Therefore, much attention in the CRBRP core T&H design has been pl aced on core a-i f icing.
Orlficing analyses do not provide the final design results.
Following the orificing, T&H perf ormance parameters of the core assemblies are' predicted.
Using these predicted perf ormance parameters, actual design ~ calculations are
~
conducted to assess the adequacy of the design.
If all the design constraints were already f actored in the orificing, no f u~ther Iteration would be necessary.
Although exact prognostication and correct representation of all the constraints is not always possible, a priori consideration of the design constraints as orificing guidelines nevertheless serves as a useful means in enhancing the ef ficiency of the analysis process.
This was the approach adopted in CRBRP core T&H analyses where a systematic orificing analysis was developed, which accounted for lifetime /burnup, transient, upper internals temperature constraints. This new approach represented a change in philosophy and a significant improvement over the previous maximum temperature equaliza-tion method.
Characteristic features of this approach are determination of the limiting temperatures (see Section 4.4.2.5.2) for all types of assemblies and simultaneous orificing of the f uel and blanket assemblies.
Final ly, both first and second core conditions were Investigated in determining the orific-Ing constraints and the most restrictive in either core was used in deriving the ori ficing configuration.
This guaranteed, a priori, that the thermal-hydraulic perf ormance would satisfy the constraints considered in both cores.
4.4-10 Amend._ _6_9 July 1982
r l
The following orificing constraints (Reference 1) are satisfied in selecting the flow orificing for the CRBRP f uel, inner blanket and radial blanket assemblies:
^
o Maximum cladding temperature must be compatible with lifetime and burnup objectives, which can be expressed in terms of maximum allowable inelastic cladding strain and cladding cumulative damage function (CDF);
Maximum coolant temperature conditions must be such as to assure, with o
adequate margin, that no bolling occurs during the worst emergency transient (e.g., the three-loop natural circulation event), accounting f or uncertainties at the 3 level confidence; Maximum assemblies mixed mean outlet temperature and radial o
temperature gradient at the assemblies exit must be compatible with upper internals structure (UIS) limitations; Maximum of eight discrimination zones (fuel plus inner blanket) are o
al lowed; Flow allocation to fuel, inner blanket and radial blanket assemblies o
must not exceed 94.0% of the total reactor flow to account for cooling requirements of other reactor components.
Since the heterogeneous core contains a single f uel enrichment zone and because the nut.ser of required discriminators depends on the unique combinations of flow orificing and f uel enrichment zones, the maximum number of fuel plus inner blanket assembly orificing zones is equal to the total allowable number of discriminators (i.e., 8).
Inner blanket and f eel assemblies employ identical inlet nozzles.
Therefore, both must be considered in determining the total number of discriminator zones.
The outer 6lanket assemblies employ a unique inlet nozzle and, theref ore, are not considered in determining the total number of discriminators.
The two 6 corner positions (*)
which alternate between inner blanket and f uel assemblies during successive cycles, form a separate. discriminator zone which is included among the eight.
To put the lifetime /burnup and translent temperature constraints on the same l
basis and to provide quantitative, comparable orificing guidelines, the concept of equivalent limiting temperature is employed.
The equivalent limiting temperature is defined as that cladding temperature at a specified l
radial position (cladding ID in these analyses) and time in life (end-of-Ilfe) which must not be exceeded in order to satisfy the considered constraint.
(*)A map of the 600 core symmetry sector analyzed in the thermal-hydraulic studies and assemblies numbering scheme are shown in Figure 4.4-9.
C 4.4-11 i
Amend. 69
. July 1982
r Three equivalent limiting temperatures were defined to represent the lifetime /
burnup and transient constraints, i.e., SELT, DELT and TELT. They are defined as the end-of-life maximum cladding ID temperatures f or Plant Expected Operating conditions (see Section 4.4.3.3.1), considering uncertainty factors at the 2 level of confidence, such that accounting f or the assembly temperature / pressure lifetime history, the limiting value of the inelastic cladding strain (SELT), or cumulative damage function (DELT), or worst time-In-life transient coolant temperature (TELT) is not exceeded.
As it appears f rom the above definition, the equivalent limiting temperatures are calculated for each assembly.
In f act, all the various assemblies have Individually dif ferent lifetime histories of cladding temperature and fission gas pressure, and theref ore, the limiting equivalent temperatures are necessarily dif ferent f rom assembly to assembly to stay within a constraint common to all assemblies.
Calculations are performed for plant expected operating conditions, which are the conditions where the CRBRP is expected to operate on a probabilistic basis and the conditions used in the design of replaceable components such as the core assemblies (see Section 4.4.3.3.1).
As previously mentioned, both first and second core conditions have been considered in defining the core orificing, theref ore, the SELT, DELT and TELT have been calculated for both cores, in the case of the radial blanket assemblies, where the lifetime spans both cores, obviously only one set of limiting temperatures was calculated.
Using the DCTDPUS code, the assemblies minimum flow in the first and second core necessary to satisfy the most restrictive of the limiting conditions was calculated for each assembly.
Subsequently, the various assemblies.were grouped in zones and the orificing arrangement was selected such that the flow allocated to each assembly was at least equal to.the larger of the flow requirements in first and second core.
This assured meeting all constraints f or both cores.
Final ly, the excess flow, if any, is allocated among the fuel assemblies to minimize and equalize the assemblies exit temperature and temperature gradients.
4.4.2.5.2 Calculation of Eaulvalent Limitina Temoeratures Assemblies lifetime /burnup goals are achieved when both.the cladding inelastic strain and cladding CDF.are within the established limits 'during steady-state operation. The ductility strain guideline was set at 0.2% and the CDF guide-line for orificing analyses was set at 0.7 in the f uel assemblies and 0.5 in the blanket assemblies.
Since the CDF limit for steady-state plus transient operation is by definition 1.0, the margin for CDF transient accumulation was I
0.3 in the fuel assemblies and 0.5 in the blanket.
Both cumulative cladding strain and CDF depend on the rod cladding temperature / pressure historf. Thus, I
using a preliminary estimate of the assembly flow (but using the proper physics data), the hot rod (*) in each assembly at end-of-Ilfe was identified t
4.4-12 Amend. 69 July 1982 1
n
r using the subchannel analysis code COTEC.
Subsequently, the hot rod was followed throughout lifetime and the lifetime temperature / pressure history was calculated with the NICER code.
Uncertainty f actors (see Section 4.4.3.2) at the 2cr ievel of confidence were used in the cladding temperature / pressure calculations.
Based on the above lifetime histories, a strain equivalent limiting temperature (SELT) was calculated f or each assembly. The SELT serves as an analytical expression of an orificing guideline.
It represents the end-of-lif e temperature which, if maintained constant throughout lifetime, would cause an end-of-life cumulative strain of 0.2% for the particular assembly relative behavior of cladding temperature and pressure through l i f etime.
Accordingly, the SELT does not depend on any guessed value of assembly flow, but rather on the relative behavior through lifetime, which is only a f unction of the power generation changes during life.
Since the DELT is the equivalent end-of-life temperature corresponding to a l CDF of 0.7 or 0.5, the method employed in its determination was to extract it f rom a curve correlating the cladding ID temperature at EOL with the corresponding CDF.
Thus, at least ihree (in some instances more were necessary) lif etime temperature / pressure histories were generated for each assembly by varying the flow and the corresponding CDF was calculated.
Typical curves are reported in Figures 4.4-10 through 4.4-14 for the fuel and inner blanket assemblies (first and second cores) and radial blanket assemblies.
By interpolation, the DELT corresponding to the CDF constraint was then determined.
l Regarding the transient constraint, the design guideline is to provide adequate margin-to-sodium boiling throughout the assembly lifetime during the worst transient.
This was quantitatively translated into an orificing guide-line of 15500F'which was conservatively defined as the maximum coolant temperature allowable during a natural circulation transient in any assembly at any time in life accounting for uncertainty factors at the 3 level of confidence.
This guideline also assumes plant THDV conditions and e 7500F reactor inlet temperature.
(*)Each assembly is characterized by its hot rod at end-of-life, which is obviously the one with the highest strain and CDF.
l t
l l
4.4-13 i
Amend. 69 July 1982
et@n.rwpavbX L%EQF1m r
the first and second cores, minimum flows must be put on the same basis.
Cycle 4 was chosen as the standard basis since it will require the higher core flow fraction (fuel assemblies are in alternating row 6 positions).
When flow requirements for cycle 2 are translated to cycle 4 equivalent values, second core requirements are found to be slightly more restrictive in some outer fuel assembl ies, as shown i n Figure '4.4-17.
Cycle 5 flows are reported f or the transient limited second row radial blanket assemblies, since their TELT's are maximum at EOL.
Using the required minimum flows as guidelines, the CCTOPUS code selected, for a given number of orificing zones, that combination of assemblies grouping into orificing zones which among all the various possible combinations, yleided the minimum value of total core flow and was therefore the most offective.
As mentioned in Section 4.4.2.5.1, a maximum number of eight discriminators (and orificing zones) is allowed for the fuel and inner blanket asse'mbl i es.
Four orificing zones in the radial blanket assemblies were chosen, thus, the total number of core orificing zones resulted equal to 12.
The selected arrangement is reported in Figure 4.4-18, where the starred assemblies are the ones which determine the amount of flow allocated to the orificing zones (they are called zone driver assemblies, or drivers).
Also Indicated are the limiting assemblies in each orifice zone for first and second core; obviously the driver is the one with the more restrictive flow requirement (compare with Figure 4.4-17).
As shown in Figure 4.4-18, the orificing arrangement does not have a 300 symmetry because the control rod location and insertion pattern, hence the power generation, does not have a 300 symmetry.
For example, considering the assemblies around the row 7 corner control assemblies (see Figure 4.4-16),
first core conditions are limiting f or the f uel assemblies around the control assembly at the right of the figure, while second core conditions are prevalently limiting for the fuel assemblies surrounding the control.* assembly at the l ef t.
The minimum anount of core flow necessary to satisfy the various constraints and the grouping of the core assemblies into 12 orificing zones was equal to 93.07% of the total reactor flow of cycle 4 conditions.
Since 94% of the total reactor flow is allocated to the fuel and blanket assemblies and since 93.07% is the minimum required to meet the conservat!vely selected con-straints, it follows that slightly less than 1% of the total reactor flow is available to be allocated as deemed desirable by the designer.
Usually, if a significant amount of excess flow is available, this is distributed among the f uel assemblies to minimize / equalize the assemblies mixed mean temperature and temperature gradient.
This was not, however, the procedure adopted in these studies since the amount of available excess flow is not enough to signifi-cantly influence the value of the outlet temperatures.
Additional ly, the l
relative assemblies power generation and the sophisticated orificing, which o
4.4-16 Amend._69 July 1982 l
are characteristics of this heterogeneous design, yielded maximum dif ferences in exit temperature (see Sections 4.4.3.3.3 and 4.4.3.3.5) between two adjacent assemblies (which generally occur at the f uel/ Inner blanket Interf ace in rows 6 through 8) within the UlS capability. Therefore, the excess flow l
was distributed roughly evenly among the various core orificing zones.
The final core flow allocation is reported in. Table 4.4-4, which shows the cycle-by-cycle yarlation of fIow in the various orifIcIng zones.
Both thermal-hydraulic design value (THDV) and plant expected operating condition (PEOC) flows are reported in Table 4.4-4.
Subsequent performance predictions and design calculations reported in Sections 4.4.3.3 and 4.2.1 demonstrated that the core orificing so determined was adequate and that design constraints and objectives were met, 4.4.2.6 Reactor Coolant Flow Distribution at low Reactor Flows t
The normal mode of CRBRP core heat removal upon reactor shutdown is by forced circulation f rom AC powered pony motors (which have emergency backup power l been designed to have the added capability of adequate cooling by means of f rom diesel generators) driving the primary pumps.
However, the CRBRP has natural circulation.
This inherent emergency coolant flow is provided by the thermal driving head developed by the thermal center of the IHX being elevated above that of the core (plus the respective elevation dif ferences in the intermediate loops and steam generator system).
At the N10% pony motor flow level af ter shutdown, insignificant flow redistri-bution occurs between the parallel flow core assemblies.
However, for the core natural convection cooling mode, the ef fect of dynamically 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 dif ferent thermal heads and hydraulic characteristics df the core
- assem-bites become important.
In general, the core thermal head becomes significant relative to the fccm and friction loss scross the core below 5% of full flow.
Coupled with the flow redistribution, significant heat redistribution on an Inter-and intra-assembly basis occurs throughout the core due to large tem-perature dif ferentials and an increased heat transport time (Iow power assem-biles can have a transport time of over 20 ssconds). These effects (i.e.,
natural convection flow and heat redistribution) are found to significantly reduce maximum core temperatures. This has been demonstrated in the EBR-ll and FFTF natural circulation experiments (Ref. 68 and 79).
In addition to the in-pile data, a large out-of-pile data base exists to characterize the flow behavior of the various corrponents over a wide range of operation, including low flow conditions.
A listing of the experimental data ref erences for flow distribution calculations is provided in Table 4.4-36.
Independent studies outside the CRBRP Project have been published which show a significant decrease in predicted maximum core temperatures due to reactor fIow redt strIbution durIng natural circuiatton conditions.
For exampie, Brookhaven National Laboratory ( Agrawal, et.al., in Ref. 69), using the SSC-L code, predicted localized flow Increases as large as 20% in the hot f uel
}
assembly and 40% in hot blanket assembly for the CRBRP during natural convection cooling.
Corresponding reductions in the predicted maximum 4.4-17 Amend. 69
-July 1982
W W W T W,56T M*M
)
transient coolant temperature on the order of 16 and 22% ( 1300F and 2100F) were shown f or the hot f uel and blanket assemblies, respectively, relative to the maximum temperatures predicted without flow redistribution.
Similar results were found in Ref erence 70 using the CURL-L code.
For these studies, Inter-assembly heat transf er as well as intra-assembly flow redistribution and heat conduction ef fects were neglected.
Inclusion of the.co effects would f urther reduce the maximum core temperatures.
Preliminary studios with CORINTH have been perf ormed to demonstrate the ef fect of inter-assembly flow redistribution f or the heterogeneous core design.
The ef fects of Inter-assembly heat transfer and intra-assembly flow and heat redistribution which wer e neglected are discussed later.
Figure 4.4-66 shows the results of these analyses f or the peak f uel, peak inner blanket and peak radial blanket assemblies.
Figure 4.4-67 shows results f or a typical ori ficing zone f or the f uel, inner blanket and radial blanket assemblies.
Condistent with other natural circulation studies, the flow increase to the hotter core regions is apparent.
This ef fect, along with the other natural convection phenomena, will significantly decrease the maximum hot rod temperatures in the core.
To assess the ef fect of all natural convective cooling phenomena (i.e., inter-and Intra-assembly flow redistribution and heat transfer) on the maximum transient coolant temperatures in the CRBRP core, the following system of three computer codes is used:
- 1) DEMO - predicts the overall plant-wide, dynamic natural circulation performance and defines the core boundary conditions;
- 2) COBRA-WC - predicts the detailed dynamic, core-wide perf ormance including all Inter-and intra-assembly flow and heat redistribution effects;
- 3) FORE-2M - predicts the localized hot rod dynamic temperatures including ef fects of localized od phenomena and uncertainties in nuclear / thermal-hydraulic / mechanical data.
A linkage between the COBRA-WC and FORE-2M codes has been developed to incorporate the Inter-and intra-assembly phenomena into the loca!! zed hot rod transient analyses by using the expression f or the heat transported to the coolant f or each axial node of the hot element modeled in FORE-2M.
Coupl ed w'th this, the axial mass flow rate f or each axial node is also input f rom COBRA-WC analyses.
The heat and axial mass flow rate for each axial node are based on nominal conditions in the COBRA-WC code. This is a conservative approach because these values are lower than those calculated for the hot channel temperature conditions and thus, result in a conservatively higher predicted hot channel temperature.
4.4-18
_ lcend._69 July 1982
Based on a 90%/10% Pu (239 + 241)/U-238 fission rate split, the weighted everage fission gas yield value may be calculated directly from the data presented in Table 4.4-15.
The value of the Xe + Kr fission yield in fuel rods resulted equal to 0.249.
For the blanket case, the fuel isotopic composition, and hence, the Isotopic fission rate, changes significantly with burnup (plutonium accumulation). For a fresh assembly in either the inner or outer blanket, about 90% of the fissions occur in U-238, and the renalning 10% occur in U-235. Therefore, the beginning-of-life fission gas yield is equal to 0.240.
At end-of-life, just prior to discharge, the breakdown of fissions is as follows:
In U-238, 33% for inner and 16% for outer blanket; in U-235, 2% for both inner and outer blankets; in Pu-239, 65% for inner and 82% for outer blanket.
Thus, the fission gas yleld calculsted from data in Table 4.4-15 is 0.247 in Inner blanket assemblies at EOL and 0.249 in outer blanket assemblies at EOL.
Conservatively, a nominal fission yield of 0.249 constant throughout life for both fuel and blanket assemblies was adopted.
The isotopic uncertainty in the ENDF/8-IV fission yields results In a tJ.5%(15) uncertainty in the rare gas (Xe+Kr) yleid from U-235, U-238 and Pu-239 fissions. Therefore, the 2e fission yleid adopted in plenum pressure calculations was equal to 0.266.
The substantial conservatism in calculating plenum pressures is discussed in Section 4.4.3.2.4, together with a quantitative evaluation of the over-estimation of plenum pressure for +wo typical blanket rods.
4.4.2.9 Thermal Effects of Ooerational Transients Current design practice is that LMFBR components must mesh the required conditions of ASME Code Section lli (Ref. 43) and RDT Standard C-16-1T,(Ref.
44). Transient reactor design events are divided into categorius of normal, upset, emergency and f aulted according to their IikelIhood:of. occurrence.
Table 4.4-16 gives: a)'the definitions for the various incidents; and b) the allowable severity with respect to structural consequences.
Note that the RDT Standard respective terminology for the events are:
normal operation, anticipated fault, unlikely fault and extremely unlikely fault.
Table 4.4-17 presents a summary of preliminary design criteria (Limits and Guidelines) for emergency and f aulted events to assure that the core operates safely over its design lifetime and meets the requirements of the ASPE Code and RDT Standard.
The frequency of occurrence and classification of events is established by the designer based on Industrial and nuclear experience and also the special characteristics and differences in LMFBR design (as compared with an LWR for example).
Under normal steady state operating conditions, the cladding is loaded due to the Internal gas pressure.
Fission gases are released from the fuel with burnup, and thus, the Internal pressure continually increases over the rod's t
4.4-42 Amend. 69 b
The primary pump head flow characteristics and reference operating points are presented in Section 5.3.2.3.1 and 5.3.3.3 and Figures 5.3-19, 5.3-20 and 5.3-21.
The primary pump flow coastdown is presented in Figure 5.3-22.
The Intermediate pumps are identical to the primary pumps with the exceptions noted in Section 5.4.2.3.1 with operating characteristics shown in Figure 5.4-3.
4.4.3 Evaluation 4.4.3.I Reactor Hvdraulic!;
The total reactor flow rate Is one of the primary parameters that affect the thermal performance of the CRBRP.
The hydraulic analyses include the effects of uncertainties such as:
instrumentation errors, correlation uncertaintles, experimental accuracy, manufacturing tolerances and primary loop temperature l and 4low uncertelnties.
The method used to perform the steady-state hydraulic analysis consists essentially of identifying all possible flow paths in the reactor, establish-Ing a hydraulic network and solving the network by use of such codes as l CATFISH and HAFMAT.
Solution of the network will provide reactor flow rate and flow distribution within the reactor for certain specified plant operating conditions, which in the case of the. CATFISH code are the pump head / flow characteristics curve.
The CATFISH code includes pressure drop analytical correlations obtained from the results of the out of file tests reported in Tables 4.4-36.
The coolant flow distribution is determined by the geometry of the regions through which sodium flows.
Their hydraulic impedance establishes the reactor pressure drop and pressure distribution. These paths include Inlet and outlet nozzles, Inlet and outlet plena, core support structure modules, annulus between radial shielding and core barrel, annulus between vessel and core barrel, annulus between vessel and vessel liner r,nd the core assemblies upper Internal structures region.
Because of their importance, the resistance and hydraulle characteristics of the main flow paths ere determined by scale model l tests.
The tests conducted for CRBRP are discussed in Sect!cn 4.4.4, Testing and Verification.
Prior to the availability of data from these tests, the I results from similar tests in the FFTF Development Program are used where applicable.
Also see Section 4.4.2.7 for a discussion on hydraulic Impedance l correlations.
In addition to the main flow path, leakage flow paths exist in the CRBRP; these are taken into account in the flow distribution studies, but no credit Is taken for Ieakage fIow when satisfying cooling requirements.
Seals between the core support structure and the core inlet module Iiner, between various parts of the hydraulic balance system, etc., form flow paths for leakage.
The design objective of the seals is to minimize leakage.
Where possible, the piston ring type seal developed in FFTF will be used and others of different design will be evaluated experimentally with the Intent to minimize leakage.
t 4.4-45 Amend. 69
~ ~ ~
JtTI D 982
4.4.3.2 Uncertainties Analvsts (*)
4.4.3.2.1 Introeuction The impact of theoretical and experimental analyses uncertainties, instrumen-tation accuracy, manuf acturing tolerances, physical properties and correla-tions uncertainties must be considered in~ predicting the reactor thermal-hydraulic performance to ensure the safe and reliable operation of the CRBRP core and to guarantee that proper margins are provided so as not to exceed the design limits and requirements.
Hot channel / spot f actors for all core assemblies have been determined to account quantitatively for the above uncertainties.
Consistent with previous studies, the semi-statistical hot spot analysis method is used for the CRBRP core assembiles; i.e., random variables are combined statistically and.
together with the direct bias uncertainties they characterize a hot channel /
spot as the one affected by the simultaneous occurrence of all uncertainties.
Predicted hot channel / spot temperatures are the ones to be compared with the required limits.
The preliminary uncertainties analysis made certain simplifying assumptions, such as the overall tanperature dif ference is a linear f unction of Individual variables, statistical uncertainties.are normally distributed, and a large number of samples are implicit in the data base. The effect of these assump-tions have been investigated in a detailed study (Ref. 19) which showed that the overall uncertainty analysis approach adopted in these analyses is con-servative.
A full evaluation of the adopted uncertainties, of the confidence levels of the hot channel factors and of the ef fects of non-linear application of the hot channel factors, will be performed for the FSAR.
Use of the semi-statistical method requires the separation of the variables which cause the hot spot temperatures into two principal groups, one of statistical origin and the other non-statistical. The two categories are defined below.
A non-statistical (or direct) uncertainty is de' fined as a variabis,-the exact value of which cannot be predicted in advance, but which
(*) The Information specified in the Standard Format and Content for Section 4.4.3.2 " Influence of Power Distribution", is included in Section 4.4.3.3 to enable the inclusion of this major area of T&H analysis as Section 4.4.3.2.
t 4.4-46 Amend. 69
. July 1982
-cu,<<-
.y iv usu.,u,v v
,s, 4.4.3.3 Steady-State Performance Predictions Reported in this section are the analyses performed to characterize the steady-state thermal behavior of the CRBRP core together with highlights of the results.
For a much more detailed report of the results, see Sections 4 and 5 of Reference 3.
4.4.3.3.1 Plant Conditions Two sets of plant conditions are used in the thermal-hydraulic design, i. e.,
plant thermal-hydraulic design value (THDV) conditions and plant expected operating conditions (PEOC).
The THDV condftlons (730 F inlet /995 F outlet temperature; total reactor flow 41.446 x 10 lb/hr) are the Clinch River rated plant conditions and are used in: a) analyzing permanent components which have the same 30-year lifetime as the plant; b) transient and saf ety analyses, since they are more conservative than the plant expected conditions and represent the " worst bound" of plant conditions.
The plant expected operating conditions represent the plant conditions at which the CRBR is expected to operate accounting for the operating conditions of the heat transport systems, such as pump characteristics, reactor and primary loop pressure drop uncer-tainties, fouling and plugging of heat exchangers, etc.
During actual reactor operation, the long-term damage accumulated by the f uel and blanket assembly components is expected to correspond.to the damage which would be calculated using time averaged nominal temperatures.
However, in assessing the ef f ects of steady-state operation and anticipated faults (normal and upset condi-tions), fuel and blanket assembly component temperatures are based on maximum expected plant operating conditions (PEOC) and upper 2a levels.
At this level, there is a 97.5% probability that the corresponding temperatures are not exceeded. This is conservative since the calculated damage accumulation generally increases with temperature.
For the unlikely and extremely unlikely events (emergency and faulted conditions) an upper limit on plant conditions (Thermal Hydraulic Design Val ues - THDV) and the upper 3a vncertainty level ir used, simply to add additional conservatism for the saf ety analyses.
At this level, the probability of exceeding the calculated temperature Is u0.1%.
~
The above designated use of plant conditions and uncertainties derives from the pranise that stochastic f ailures are not a safety issue and the plant is capable of operation with limited fuel rod cladding failures.
To support safe operation with f ailed f uel, all the saf ety analyses described in Chapter 15 af this PSAR are based on continued and extended plant operation with 1% f aller j
l fuel.
l The primary heat transport system principal parameters (Inlet, outlet tempera-ture and AT) are evaluated, together with the asso: lated uncertainties.
The j
results of this study for the heterogeneous core, which comprised a Monte Carlo type analysis, are reported in Table 4.4-28.
some significant features are: 1) the consideration of the progressive fouling of the heat exchangers during the plant 30-year lif etime, which af fects the predicted values of the plant operating conditions (rather than conservatively assuming end-of-life
- fouling, i.e., after thirty years operation); and 2) a comprehensive account-Ing of all uncertainties affecting plant operation.
Plant expected operating conditions are adopted in core thermofluids analyses of replaceable compo-nents, such as the core assembiles, chiefly In determining the f uel rod para-l meters (cladding temperature, fission gas pressure) which are the basis for 4.4-57 Amend. 69
~ ~ -
Mip982 g-.,
evaluating the structural behavior and for assessing whether lif etime/burnup
~
objectives are actually met.
Plant expected operating conditions and associated uncertainties adopted in the thermal performance analyses are reported in Table 4.4-29.
Following is a brief discussion of the rationale in determining the values reported in Table 4.4-29 f rom the ones in Tabl e 4.4-28.
First, the mean values of Table 4.4-28 are chosen as the nominal values of Table 4.4-29, thus, conservatively including the bias factor directly into the nominal values.
Since the most critical time for core h
i i
I i
j i
i I
l t
4.4-57a Amend. 69 l
. July 1982 i
K g; 1 (62-0$06) L8,04j #92 For the bl anket assemblies, inner blanket assembly 99 was investigated as the blanket assembly having the highest power in the fIrst fIve years of CRBRP operation.
Assembly 99 reaches its maximum power in the second core, at end-of-cycl e 4 (see Fi gure 4.4-33).
Inner blanket assemblies envelope with respect to power-to-melt conditions the longer residence time radial bienket assembiles.
Both the hot and the peak rods were investigated, since the peak pin has the highest lincar power, while the hot pin has the highest cladding tmporature.
The cladding temperature has, in f act, a very significant ef fect on cladding swelling, hence on f uel/ cladding gap size, hence gap conductance, fuel temp-erature and f inal ly on power-to-mel t.
Thus, both the hot pin and the peak pin need to be investigated.
Analysis of the hot pin was obviously not necessary for the f uel assemblies, since their critical time in life is at beginning-of-lif e, rather than end-of-lif e as f or the bl anket assemblies.
Finally because the maximum power in blanket assemblies occur at end-of-life, the programmed start-up cannot af fect the power-to-melt in the blanket.
The axial positions where the cladding ternperature and the linear power rating are maximum were Investigated in addition to intermediate positions between the two above.
Also considered were: a) when the blanket pins go through a f ull overpower f actor of 1.15 at EOL; and b) when the reactor power is increased to 115% of rated power from the top of the allowed variation, i. e.,
with an overpower factor of 1.15/1.03, it was f ound that the no-melting criterion is f ully satisfied in the worst The peak pin has 0.4% less margin than the hot pin.
When the overpower case.
excursion is a full 15% the margin is 0.4% less than for the case when the reactor power is ramped f rom 1.03% of the rated power.
Substantial conser-vatism was implicit in the analyses (e.g., in cladding swelling evaluation, adopting a direct combination of nuclear uncertaintles), thus, removal of the impiicit conservatism and f actorIng of experimental data When avalIable, would substanti al ly improve the power-tcr-melt margln.
4.4.3.3.7 Control Assemblies Thermal-Hydraulle Performance The CRBRP has two control systems: primary and secondary control rod system (PCRS and SCRS) with nine (9) and six (6) control assembiles, respectively.
Detailed design features of the systems are provided in Section 4.2.3 (Reactivity Control Systems).
The bases and methodology of the thermal-hydraulic analysis of the primary control assemblies f ollowed that used in the homogeneous core design, reported in Reference 13.
A summary of the principal operating permeter for the primary and secondary control assemblies are presented in Tables 4.4-32a and 4.4-32b, respectively.
Val ues reported in Tabl e 4.4-32 are f or the row 7 corner assembly, which is the thermally limiting PCA.
Key hydraulic perf ormance assessments relate to the assembly flow margin to control rod flotation and control rod scran dynamics.
The PCA E-Spec. re-quires that the control assembly design shalI assure that the control rod cannot be Iifted (or fIcated) from the f ully inserted position, under maximum l
assembly flowrate (and pressure drop) conditions, more than the distance causing a reduction in shutdown reactivity margin equal to the stuck rod 4.4-65 Amend._6_9 July 1982
margin.
This requirement shall apply to all 9 rods, either with the driveline connected to the control rod or to ref ueling conditions f or which the drive-line is disconnected and withdrawn to its ref ueling position.
Both experimental and analytical investigations were conducted to assure the PCA will not float under the worst possible conditions; the results of these investigations are summarized in Table 4.4-33.
Data, analytical results and margin-to-flotation were expressed in terms of both assembly flowrate and pressure drop across the absorber bundle.
Prototypic testing of the CRBR PCA provided experimental measurements of the PCA flotation characteristics; experimental uncertainties were directly superimposed over the observed values.
On the other hand,.the maximum flow through the PCA was calculated with the CATFISH code accounting f or all the various ef fects causing a flow variation in the PCA.
Specifically, the hydraulic resistance uncertainties in all core components were varied by their maximum value and, conservatively, the absolute variation in the PCA flow and P was taken as increasing the design value.
The three leading causes for en increase in the PCA flow were found to be: primary pumps at their maximum speed resulting in a maximum reactor flow equal to 115% of the rated THDV value; PCA orifice resistance at its minimum; and LIM containing the PCA at is minimum resistance allowable.
As reported in Tablo 4.4-33, the individually induced variations in the PCA flowrote and P were combined at various levels of conservatism, ranging from 25 and root sum of the squares to 3a and absolute sum combination.
Cor-respondingly, the flotation margin ranged from 15% to 3% in tenns of AP and f rom 9.5% to 5.5% in terms of flowrate.
in all cases a large amount of con-servati sm was included, f or example: a) by comparing the minimum experimental with the maximum predicted flotation characteristics, analytical and experi-mental uncertainties were superimposed rather than combined statistically; b) use of absolute rather than relative values of the PCA flow (and AP) variations does not take into account the variations causing decrease, rather than increase of the PCA flowrate and AP.
In spite of th!'s conservatism, a posit!ve margin to flotation resulted under the worst conditions, as shown in Tabl e 4.4-33 The secondary control rod system uses the concept of hydraulic scram ' assist l
design with a net hydraulic force in the 150-250 lbs. range on the control rod when f ully withdrawn f rom the core. The same magnitude of downward hydraulle force (in addition to the weight of the assembly) is also available under the abovementioned design conditions.
Thus, it is concluded that the secondary control rods do not float at 100% flow (even when disconnected).
Predicted control rod scram performance of the primary control rod system is reported in Section 4.2.3 (Reactivity Control Systems).
I Figures 4.4-54 and 4.4-55 show typical PCA absorber region temperature distri-butions under the minimum withdrawal and f ull withdrawal control rod cond!-
l tions, respectively.
4.4.3.3.8 RRS Thermal-Hvdraulle Analvses The steady-state duct temperatures at PE0V conditions were calculated for a t
300 sector of the RRS.
4.4-66 Amend. 69
. July 1982
~~--__
iu3e s w.usus; Lu,una a sz The region analyzed is partially rhown in Figure 4.4-56.
The model consists of all 29 RRSA's in a 300 sector, plus a corresponding section of fixed radial shielding (FRS), core barrel (C8) and core 4.4-66a Amend. 69 July 1982
TABLE 4.4-17
SUMMARY
OF PRELIMlNARY DESIGN CRITERIA l
Event Classification Severity Level Criterton**
Emergency Minor incident The total cumulative' damage function is to be less than 1.0.
(Unlikely Faults)
The accumulated plastic and thermal creep straln' is to be less than 0.3%.
Faulted Major inctdent No ciadding melting (temgeraturelessthan 2475 F) and (Extremely UnlIkeIy
- No sodium bollIng (temperature less than saturation temperature at the existing pressure).
l
- Sodium bolling temperature is quoted as a guideline to establish that no cladding melting can occur.
- The emergency criteria are limits.
The faulted criteria are guidelines.
i s
4.4-103 Amend. 69
, July 1982 t
TABLE 4.4-32a PRIMARY CONTROL ASSEM3LY OPERATING PARAE TERS o Number of orificing zones 1
o PCA's total flow allocation (fraction reactor flow) 0.01 o Flowrote (PEOC, Ib/hr) 49,500 Flow split (bundle / total assembly flow) 0.62 o
o Maximum bundle flow velocity (f t/sec) 8 o Maximum hot rod midwall cladding temperature 1006 (PE00, 2a, OF) o Maximum fission gas pressure (27, psla) 3600 8 275 fpd o Maximum linear power rating (36+ overpower, Kw/ft) 16 Bottom 1.4 Top o Maximum absorber temperature (THDV nominal, OF) 3367 Maximum mixed mean exit temperature (THDV ner.ilnal, OF) 853 o
o Maximum exit gradient (nominal, CF) 246 I
i l
l l
i t
4.4-124 l
Amend. 69 l
"- ~
July 1982 o ',
~
-p
- e
' '4, / (9
,6
$s
)
y s\\h i
\\,
., n <,.
m TABLE 4.4-32b
., i i
(
a -
- 1
+
SECONDARY CONTPOL ASSEMBLY CPERATING PARAMETERS
~
.)
4 Flow Rate (THDV, Ib/hr)
N/
s Control Rod Flow
, 'j 9,130 b,
3
.s Bypass Flow
-)
-9,330
's t',
.y i ' '-, ',
i 8,4*j0 Total UpfIow z
Downflow 50,7)0.
I f 7
s' 69,!7,0 f,
Total AssembIy i
+
5 Hydraulic Scram Assist Force at Full Flow (!bs) 148 - 248 i
Peak Linear Power (kw/ft)
- 4. 0,
/
Outlet Temperature (THDV, Nominal, F)
'j s'
i
.i
+
1 Control Rod Bundle 829 N
,s.
Assembly 854 a.
,3 Maximum Cladding Midwall Temperature ( F)
I L
m Nominal, THDV 853
,q Hot Spot (THDV, 2a) i 895
(<>
.j Maximum Absorber Temperature ( F) is
r.,,1 l
Nominal, THDV
. 1054 y i ).
.\\
,(d 1
s l
Hot Spot (THDV, 36)
.i+ 1188 y
f f,
-}A'
' )'Q' h*
+
s i
/
'p s
s so l
n I
/
I
'y s
a
\\
l d
.%g,
a s
h r
4 l
5 4.4-124a 1
,(
'.j \\
Amend.g69 I
3 b
\\
.luIy 1982
,,. s 1
\\
., ~, /
L a
M-
-M
l 8 uSL i 404 VJiJJ Lu,U4J #UD TABLE 4.4-33
SUMMARY
OF FLOTATION EVALUATION Rod Bundle Pressure Drop Flow Rate i
< e (Pst)
_1Lb/Hr) d
- 1. Test observed flotation conditions 7.5 57,500
- 2. Above, accounting for test uncertainties 7.2 55,000
- 3. PCA operation conditions, nominal, PE00 5.94 49,600 s
- 4. Total ef fect of core components hydraulic resistance uncertaintles:
o r.s.s. combination 2a/3C 0.35/0.52 687/976 s
o Absolute sum combination 21/3e 0.73/1.06 1932/2696
- 5. Maximum design conditions (3+4):
'o r..s.s. combi nation 2a/3cr 6.29/6.46 50287/50576 o (.'bsoluto sum combinailon 2c/36 6.67/7.0 51532/52296
- 6. Margl'n-to-flotation (2-5):
'l,
o r.s.s. combinati on 2c/3c-0.91/0.74
4713/4424 (15/12)*
(9.5/9)*
i o Absolute sum combination 21/3a-0.53/0.2 3468/2704
~
l
/
(9/3)*
(7/5.5)*
l
- In percentage of nominal conditions.
l 1
4.4-125
~
'~
Amend. 69 JulyT981 1
,e
~s'
TABLE 4.4-36 EXPERIMENTAL DATA REFERENCES FOR FLOW DISTRIBUTION CALCULATIONS AT LOW FLOW (SECTION 4.4.2.6) 1.
PRESSURE DROP DATA - FLOW REDISTRIBUTION FOR FIGURES 4.4-66 & 67 1.
" Covered Pressure Drop Flow Test / Cross Flow Mixing Test", HEDL-TI-76049, November 1976.
(Availability: US/ DOE Technical Information Center).
2.
W. L. Thorne, " Pressure Drop Measurenents in FFTF Fuel Vibration Tests",
HEDL-TC-812, April 1977.
(Availability: US/ DOE Technical Information Center).
s 3.
W. L. Thorne, " Pressure Drop Measurements from Fuel Assembly Vibration Test ll", HEDL-TC-824, April 1977.
(Availability: US/ DOE Technical Information Center).
4.
P. M McConnell, " Clinch River Breeder Reactor Fuel Assembly inlet / Outlet Nozzle Flow Tests", HEDL-TME-77-8, February 1977.
(Availability: US/ DOE Technical information Center).
5.
H. M. Geiger, D. C. Meess and D. K. Schmidt, " Radial Blanket Flow Orificing Testing: Calibration. Tests", WARD-RB-3045-18, April 1977.
(Availability: US/ DOE Technical Information Center).
6.
F. C. Engel, R. A. Markley and A. A. Bishop, " Laminar, Transition and Turbulent Parallel Flow Pressure Drop Across Wire Wrap Spaced Rod Bundles", Nucl. Sci. Eno, 61, pp. 290-296 (1979); 24, p. 226 (1980).
7.
l. E.
Idel'chik, " Handbook of Hydraulic Resistance - Coefficients of Local Resistance and of Friction", AEC-TR-6630,1960.
8.
L. F. Moody, " Friction Factors for Pipe Flow", Trans.' Amer. Soc.~ Mech.
Eng, &&, pp. 671-684 (1944).
9.
R. G. White and M. D. Simmons, "CRBRP Fuel Assembly Flow and Vibration Tests in Water", HEDL-TME-78-106, October 1979.
10.
P. M. McConnell, " Interim Report on CRBRP Fuel Assembly Cavitation Tests in Water", HEDL-TC-967, October 1977.
11.
C. Chiu, et al., " Pressure Drop Measurements in LMFBR Wire Wrapped Blanket Assembl ies", MIT C00-2245-42TR Report (1977).
12.
" Radial Blanket Design and Development Quarterly Progress Report for l
Period Ending August 31, 1977", WARD-RB-3045-21 (1977).
13.
J. G. Akey and H. O. Lagally, " Primary Control Assembly Hydraulic Test",
WARD-D-0263, July 1980.
t i
4.4-126b tmend. 69
. July 1982
g TABLE 4.4-36 (Cont'd) 14.
D. R. Dickinson and F. H. Nunamaker, " Measurement of Outlet Plenum Velocity Profiles, Pressure Drops and Flow Splits in the IRFM of CRBRP",
HEDL-TC-1015, December 1979.
15.
P. M. McConnell, et al., " Inlet Plenum Feature Model Flow Tests of the CRBRP; Addendum Y-Resul ts", HEDL-TME-76-33, March 1976.
16.
P. M. McConnell, "CRBRP Removable Radial Shielding Orifice Pressure Drop Test Results, DRS 31.14.37", HEDL-TC-958, September 1977.
II.
- HEAT TRANSFER AND OTHER DATA USED FOR VAllDATION OF COBRA-WC (PNL-4128)
AND FORE-2M (Nucl. Eng. & Des.
- 68. No. 3. no. 323-336. Aoril 1982) CODES EDR FIGURE 4.4-68 1.
Bates, J. M., and E. U. Khan.
1980.
" Investigation of Combined Free and Forced Convection in a 2 x 6 Rod Bundle During controlled Flow Transients." AIChE Symoostum Series:
Heat Transf er - Orl ando.1980.
American Institute of Chemical Engineers, New York, New York, Series No.
199, Vol. 76, pp. 215-230.
2.
Bates, J. M., and E. U. Khan.
1980.
Investigation of Combined Free and Forced Convection in a 2 x 6 Rod Bundle During controlled Flow Transients.
PNL-3135, Pacific Northwest Laboratory, Richland, Washington.
3.
Chen, Y.
B.,
K.
lp and N. E. Todreas.
1974. Yalecity Measurements In Edge Subchannels of Wire-Wracoed LMFBR Fuel Assemblies.
C00-2245-11 TR, Department of Nuclear Engineering, Massachusetts Institute of Technology, Cambridge, Massachusetts.
4.
Chiu, C.,
et al.
1978.
Flow Sollt Measurements In LMFBR Blanket AssembIles.
C00-2245-41TR, Massachusetts Institute of Technology, Cambridge, Massachusetts.
5.
Engel, F. C., R. A. Markley, and B. Minushkin,1978.
" Buoyancy Effects on Sodium coolant Temperature Profiles Measured in an Electrically Heated Mockup of a 61-Rod Breeder Reactor Blanket Assembly." ASME paper 78-WA/HT-25, American Society of Mechanical Engineers, New York, New York.
~
6 Fontana, M. H. 1973.
Temoerature Distribution in the Duct Wall and at the Exit of a 19-Rod Simulated LMFBR Fuel Assembly (FFM Bundle 2A).
ORNL-4852, Oak Ridge National Laboratory, Oak Ridge, Tennessee.
7.
Gillette, J.
L., R. M. Singer, J. V. Tokar and J. W. Sullivan.
1979.
" Experimental Study of the Transition from Forced to Natural Circulation in EBR-11 at Low Power and Flow." Paper No. 79-HT-10, American Society of bbchanical Engineers, New York, New York.
4.4-126c Amend. 69
~-
July ~1982
~
TABLE 4.4-36 (Cont'd) 8.
C. W. Hoth.
1981.
" Fuel Open Test Assembly (FOTA) Data Obtained During the FFTF, 75% and 100% Natural Circulation Tests." HEDL-TC-1940, Hanford Engineering Development Laboratory, Richland, Washington.
9.
- Juneau, J., and E. U. Khan.
1979.
Analysis of Steadv-State Combined Forced and Free Convection Data In Rod Bundles.
FRA-TM-116, Argonne National Laboratory, Argonne, Illinois.
10.
Laf ay, J., B. Menant and J. BarroI s.
1975.
" Influence of Hellcal Wire-Wrap Spacer System in a 19-Rod Bundle." Presented to the 1975 Heat Transfer Conference, April 1975, San Francisco, California.
11.
Lorenz, J.
J., T. Ginsberg and R. A. Morris.
1974.
Exoerimental $lxing Studies and Velocity Measurements with a Simulated 91-Element LMFBR Fuel Assembly.
ANL-CT-74-09, Argonne National Laboratory, Argonne, Illinois.
12.
Lorenz, J.
J., D. R. Pedersen and R. D. Pierce. 1973.
Perloheral Flow Visualization Studies In a 91-Element Bundle.
ANL-RAS-73-14, Argonne National Laboratory, Argonne, Illinois.
13.
Milburg, M.
L., J. A. Hassberger and C. J. Boasso. 1977.
Natural Circulation Heat Transfer Testing with a Simulated Full-Scale LMFBR 217-Pin Electrically Heated Fuel Assembly. HEDL-TME-77-3, Hanford Engineering Development Laboratory, Richland, Washington.
14.
Morris, R. H. et al. 1980.
Single-Phase Sodlum Tests in 61-Pin Full-Length Simulated LMFBR Fuel Assembiv-Record of Phase 1 Exoerimental Data fer THORS Bundle 9 ORNL/TNh7313, Oak Ridge National Laboratory, Oak Ridge, Tennessee.
15.
Novendstern, E. H. 1972a.
Mixing Model for Wire-Wraoned Fuel Assemblies.
WARD-5915, Westinghouse Advanced Reactor D.lvision, Madison, Pennsylvania.
16.
Novendstern, E. H.1972b.
" Turbulent Flow Pressure Drop Model for Rod Assembl ies Util iz ing a Hel lcal Wire-Wrap Spacer System." Nuclear Engineering and Design. 22:19-27.
17.
Wang, S. and N. E. Todreas.1981.
" Computer Model for MIT Correlations for Friction Factors, Flow Splits and Mixing Parameters in LMFBR Wire-Wrapped Rod Assemblies." DOE /ET/37240-87TR, Massachusetts insti.ute of Technology, Cambridge, Massachusetts.
18.
Cheung, A. C.,
et al., " Verification of the CRBRP Natural Circulation Core Analyses Methodology with Data from FFTF Natural Circulation Tests",
CRBRP-ARD-0310, June 1982.
19.
Coffield, R. D., R. A. Markley and E. U. Khan, " Natural Circulation Analyses and Verification for LMFBR Cores", Nucl. Engrg. Des., 62, (1980) p p. 181 -198.
t l
4.4-126d Amend. 69
. July 1982 r
,,nna,
I Ugv I
LUL Usl IJ LUgUfj #JJ TABLE 4.4-36 (Cont'd) 20.
Bil lone, M. C., et al., "L IFE-I l l Fuel Element Perf ormance Code User's Manual", ERDA-77-56, July 1977.
21.
Henderson, J. M., S. A. Wood and D. D. Knight, " Sodium Boll ing and Mixed Oxide Fuel Thermal Behavior in FBR Undercooling Transient", W-1 SLSF i
expertemnts results.
22.
Freeman, D. D., "SEFOR Experimentsi Results and Applications to LMFBRs",
GEAP-13929, January 1979.
23.
Bishop, A.
A., R. D. Coffield and R. A. Markley, " Review of Pertinent Thermal-Hydraulic Data for LMFBR Core Natural Circulation Analyses'!,
AIChE Symoposium, Series 76, (1980), pp. 193-204.
24.
Coffield, R. D., et al., " Buoyancy-induced Flow and Heat Redistribution During LPFBR Core Decay Heat Removal", Proc. of Specialists Meeting on Decay Heat Removal and Natural Convection in FBRs., Brookhaven National Laboratory, New York, February 1980.
25.
Gillette, J.
L., D. Mohr, R. Singer and R. Smith, "A Flow Coastdown to Natural Convection Condiflons in EBR-II", ANS Trans. 22 (1975) p. 394.
26.
Singer, R. M. and J. L. Gillette, " Measurements of Subassembly and Core Temperature Distribution in an LMFBR", AIChE Symposium, Series 73, (1977), p. 97.
27.
Stover, R.
L., et al., "FFTF Natural Circulation Tests", Presented at ANS Winter Meeting at San Francisco, California, 1981.
t
[
]
4.4-126e Amend. 69
- __.--~J u'l V7 9 82 L