ML19305A780

Predicted Steady State Thermal-Hydraulic Performance of Fuel & Blanket Assemblies in Plant Heterogeneous Core, Revision 1

ML19305A780
Person / Time
Site:	Clinch River
Issue date:	01/31/1979
From:	Bach C, Corelli M WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
To:
Shared Package
ML19305A779	List: ML19305A778 ML19305A780
References
5737A-650A-(S14, WARD-D-0210, WARD-D-0210-R1, WARD-D-210, WARD-D-210-R1, NUDOCS 8003180333
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• e m a aa,nt Clinch Rwer Breeder Reactor Plant Nuclear lsland .

PREDICTED STEADY STATE THERPJAL-HYDRAULIC i PERFORMANCE OF THE FUEL AND O BLANKET ASSEMBLIES IN THE i

CRBRP HETEROGENEOUS CORE M. D. CARELLI C.W. BACH l

JANUARY 1979 l Prepared for the Project Management Corporation as part of the U.S. Energy Research and Development l

j Administration Liquid Metal Fast Breeder Reactor j Demonstration Program l Any Further Distribution by any Holder of this Document or of the Data Therein to inird Parties Representing Foreign Interest, Foreign Governments, Foreign Companies and Foreign Subsidiaries or Foreign l Divisions of U.S. Companies Should be Coordinated with the Director,

! Division of Reactor Research and Development, U.S. Energy Research and Development Administration O W Westinghouse Electric Corporation ADVANCED REACTORS DIVISION 6270-2 80X 158 M ADISO N. PENNSY LV ANI A 15663 8003100

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i O WARD-D-0210, REV. 1 U

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[ PREDICTED STEADY STATE THERMAL-HYDRAULIC PERFORMANCE OF THE FUEL AND BLANKET ASSEMBLIES IN THE CRBRP HETEROGENEOUS CORE L

1 January 1979 l

l PREPARED BY:

M.D. Carelli, i Core T&H Analysis c.w dd i C. W. Bach, Core T&H Analysis APPROVED BY:

R. A. Markley, f Manager, Core T&H Analysis n

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! LIST OF CONTRIBUTORS l i

j .A. Biancheria e

! C. W. ' Bach I

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M. D. Carelli K. D. Daschke

. -0. R. Forsyth I

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, A. J. Friedland i ' K. B. Kittredge, l T. S. Roth i

E. C. Schwegler

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D. R. Sundquist D. R. Spencer D. O. Tomlin 1

M. L..Travis

J. M. Willis I.

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ACKNOWLEDGEMENTS The authors gratefully appreciate the assistance throughout the entire work of J. L. Robertson, W. R. Rymer and R. A. Zebley in the endless computer runs, data elaboration and display. Also, many thanks are due to S. J. Hoover for her skill and patience in translating our " chicken scratch" into this report.

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_ () ' . TABLE OF CONTENTS P_ag_e

SUMMARY

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1. INTRODUCTION 3

2. CORE ORIFICING 10 2.1 Introduction 10 2.2 Orificing Philosophy and Approach 11 2.3 Calculation of Equivalent Limiting Temperatures 13 2.4 Results 16

3. UNCERTAINTY ANALYSIS 19 3.1 Rod and Mixed Mean Temperature Uncertainties 19 3.1.1 Introduction 19 3.1.2 ' Rod Temperature Engineering Uncertainty Factors 21 3.1.3 Nuclear Uncertainties 23 3.1.4 Plenum Pressure Uncartainties 26 3.2 Duct Temperature Uncertainties 28' O 3.2.1 Introduction 28 O 3.2.2 Fuel Assemblies 30 3.2.2.1 Heat Generation 30 3.2.2.2 Assembly Flow 31 3.2.2.3 Coolant Enthalpy Rise (AH) 32 3.2.2.4 Film 34

! 3.2.2.5 Duct 35 3.2.2.6 Interstitial Gap 36 3.2.3 Inner and Radial Blanket Assemblies 37

. 3.2.3.1 Heat Generation 37 3.2.3.2 Assembly Flow 38 3.2.3.3 Coolant Enthalpy Rise (AH) 38 3.2.3.4 .F ilm 39 3.2.3.5 Duct 39 3.2.3.6 Interstitial Gap 39 3.2.4 Primary Control Assemblies 39 3.2.4.1 Heat Generation 40 3.2.4.2 Assembly Flow 40 3.2.4.3 Bundle / Bypass Flow Split 40 3.2.4.4 Bundle Enthalpy Rise 41 i

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TABLE OF CONTENTS (Continued) ll Page 3.2.4.5 Bypass Enthalpy Rise 41 3.2.4.6 Film 41 3.2.4.7 Duct 41 3.2.4.8 Interstitial G;p 42 4 STEADY STATE THERMAL PERFORMANCE 43 4.1 Plant Conditions 43 4.2 Linear Power 46 4.3 Assemblies Mixed Mean Temperatures 47 4.4 Rod Lifetime Cladding Temperature / Pressure Histories 49 4.5 Duct Temperature and Related Analyses 51

5. POWER-TO-MELT ANALYSES 56 5.1 Introduction 56 5.2 Fuel Assemblies Power-to-melt Analyses 56 5.2.1 EBR-II Uncertainties 58 5.2.2 Results 63 5.3 Inner Blanket Assemblies Power-to-melt Analyses 67

6. CORE ASSEMBLIES PRESSURE DROPS 70 lh 6.1 Introduction 70 6.2 Pressure Orop Correlations 71 6.3 Re_sults 78

7. CONCLUSIONS 84 REFERENCES 86 5737A-650A-( S1421') iv 9

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SUMMARY

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The steady state themal-hydraulic performance of the fuel and blanket assenblies in the heterogeneous configuration of the CRBRP core is reported here.

A'nalyses were conducted' for the projected first and second core, i.e., a total of four cycles of-CRBRP operation; additionally, since radial blanket assemblies in the outemost row have a five year continuous residence prior to refueling, they were analyzed for their entire lifetime.

Relevant themal-hydraulic parameters calculated for each fuel and blanket assenbly throughout their lifetime were: maximum cladding temperature and fission gas pressure in the hot rod, mixed mean exit temperature, linear power rating. 'Three-dimensional duct temperatures on a core-wide basis were

. calculated at beginning of the first cycle and end of the fourth cycle, thus bracketing the entire behavior. Detailed pressure drops were calculated throughout the core, utilizing the most recently available experimental data (m) and sophisticated analytical tools. The effect of engineering uncertainties (hot channel / hot spot f actors) at various levels of confidence were accounted for in the analyses. Power-to-melt analyses were conducted for the highest power rod in the fuel and blanket assemblies to verify that the criterion of no-incipient melting, considering uncertainty f actors at the 3o level of confidence and a reactor power equal to 115% of the nominal rated power, was actually satisfied. A programmed startup was identified to assure compliance with the above criterion in the fuel assemblies at beginning-of-life.

Two novel concepts and analytical methods were introduced and implenented for the first' time in the thermofluids design of LMFBR cores. The first was a

.cmprehensive, integrated method of orificing the fuel and blanket assemb lies. Coolant flow was allocated to simultaneously satisfy perfomance,

. design and safety constraints such as attainment of burnup/ lifetime objectives, compliance with transient limitations, minimization of temperature level and radial gradient. in the coolant flow impinging on the upper internals, limitation-on the total number of discriminators, and consistency with cooling requirements of other reactor components.

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The second was the determination of duct temperatures accounting for uncertainties directly in the temperature calculations, rather than by <

superpositior. over nominal temperatures. Additionally, both positive and negative uncertainties were considered, i.e., hot channel factors greater and lower than unity, since under particular combinations they will yield the worst cross-duct gradients.

During the course of these studies, room for significant improvement of the thermal-hydraulic performance was identified in the reduction of the orificing resistance with consequent reactor flow increase in excess of 5%. Even though the analytical procedure and techniques have been developed, the results here reported only partially reflect this improvement. Chiefly for schedular reasons it was decided not to redo all the thermal performance analyses, since the predicted performance will be enveloping the actual conditions once the aforementioned improvements are put into effect, e.g., assembly temperatures will be lower than predicted here. Advantage of the possibility of increasing the total reactor flow will be taken in the final, as-built design analyses.

However, even neglecting this improvement, the thermal-hydraulic performance of the CRBRP heterogeneous core, as reported here, is indeed adequate and satisfactory.

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. /N 1.- INTRODUCTION

N ,I The design of the Clinch River Breeder Reactor Plant (CRBRP) core has evolved over the years and a v.ry brief discussion is necessary to put into proper perspective how the core thermal-hydraulic design and analysis has played a key role in such evolution.

The original core layout was of the homogeneous type with fuel assemblies (in two.different enrichment zones) surrounded by radial blanket assemblies. The steady state core thermal-hydraulic performance analyses were reported in References 1 and 2. Subsequently, a modified core layout (see Figure 1 and Table I) was proposed to satisfy, with margin, the original objective of a breeding ratio equal to or larger than 1.2. Various other performance and safety improvements resulted from adoption of this core, which was of the heterogeneous type, where fuel and blanket (called inner blanket) assemblies are inter-mixed in the inner region of the core and are surrounded by two rows of outer (also called radial) blanket assemblies. Fuel and inner blanket

-assemblies were alternating at each yearly refueling in six core positions.

The fuel assemblies consisted of only one enrichment. In parallel to the thermal-hydraulic performance predictions documented in this report, an extensive effort was conducted in totally reviewing and updating the CRBR hot channel factors values as well as providing their rationale and methodology of application (3) Only part of the hot channel factors recommendctions from Reference 3 were actually factored into the T&H performance predictions because the two efforts were being conducted in parallel. Nuclear uncertainties were, however, specifically derived for the heterogeneous core.

Full consideration of the new hot channel factors will be factored into the final, as-built design analyses.

It is obvious that the thermal-hydraulic performance predictions will differ for different core layouts. For example, the heterogeneous cores have higher -

linear power rating in the fuel assemblies than the homogeneous configuration due to the reduced number of fuel assemblies utilized. Two fuel enrichment zones (which affect the orificing) were adopted in the homogeneous and only

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one in the heterogeneous configuration. Radial blanket assemblies were shuffled in the former, while no shuffling of either inner or radial blanket is performed in the heterogeneous core. The projected lifetime in the homogeneous core was three years for the fuel and six years for the radial blanket assemblies, while in the heterogeneous scheme, the fuel and inner blanket assemblies have a two year lifetime, the first row radial blanket assemblies have four years, and the second row radial blanket five years (obviously the alternating fuel / inner blanket assemblies have only one year).

Equilibrium core conditions were predominantly considered in the homogeneous core analyses, while first and second cores are analyzed for the heterogeneous.

The aforementioned differences are, of course, fundamental. However, an additional reason exists in making the predicted T&H performance of the two cores different, a reason which is just as important as the different core configuration. Over the years, the knowledge of the core T&H design has obviously increased by the automatic " hands on" learning process, where new apprnaches, constraints, potential problems as well as areas of improvement have been discovered. This continuous learning process has led to shif ts in design philosophy, improvement of analytical tools, development of new computer codes, and above all, elaboration and implementation of new concepts, all towards the goal of optimizing the CRBRP core thermal-hydrau'.ic design.

Tnerefore, it has to be kept in mind that if, for example, the homogeneous core would be analyzed today, much different results than previously reported would be obtained. To visualize the impact of the continuous improvement of the core T&H analytical approach, following is a brief summary of the key differences in performing the two designs.

The most striking difference between the homogeneous and heterogeneous core thermofluids design was the adoption of a new philosophy in core orificing.

For the homogeneous, the criterion in orificing fuel and blanket assemblies was essentially to equalize the end-of-life (E0L) maximum cladding temperature, somewhat tempered by a zoning arrangement assigning relatively more flow to the high burnup fuel assemblies (2) .

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[N The idea, which germinated in the homogeneous core study, that not only

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cladding temperature,' but also burnup, had to be taken into account in orificing, was fully exploited in the heterogeneous core analyses. By this time, it was realized that, by far, the most important step in the core thermofluids design was the core orificing. Thus, flow allocation to the core assemblies simultaneously satisfied not only lifetime /burnup goals, but also transient ~ limitations and limitations (as much as practical) posed by the upper internals structure on the assemblies exit temperature and temperature gradient. Simultaneous consideration of these different constraints and optimization of the core flow allocation was possible through use of the newly deveicped OCTOPUS code (4). Limiting temperatures (see Section 2) were calculated for all core assemblies and fuel, inner blanket and radial blanket assemblies were orificed simultaneously. Limiting temperatures were also calculated for both first and second core and a combination of the most restrictive ones was selected for orificing.

Analyses were conducted in greater depth in the areas of characterization of plant operating conditions and uncertainties evaluation. Expected plant O operating conditions were calculated 'or various times in the plant life V (rather than at 30 years only as for the homogeneous core), and associated uncertainties were calculated more rigorously. Nuclear uncertainties were evaluated in more detail for the heterogeneous core, accounting for local effects due to the presence of control rods and for lifetime effects. Some of the engineering uncertainties were also further evaluated as previously discussed.

Power-to-melt analyses conducted fer the heterogeneous core introduced, for the first time in the core design, the concept of a programmed startup to increase the power-to-melt capability of the fuel assemblies, due to the fuel restructuring effect oc. curring very early in life. An elaborated power-to-melt analysis for the inner blanket assemblies was performed following realization that the cladding temperature has a very significant effect on cladding swelling, hence on fuel / cladding gap size, hence gap conductance, fuel temperature and finally on power-to-melt. This prompted examination of the hot (maximum cladding temperature) rod in addition to the peak '(maximum power) rod, which was the only one analyzed in previous studies.

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One of the codes developed during the heterogeneous core thermal-hydraulic design was the TRITON code (5) which evaluates duct temperatures by rigorously accounting for the ef fect of inter-assembly heat transfer.

Additionally, for the first time duct temperatures were calculated accounting for the effect of uncertainties. TRITON uncertainty calculatio" are performed such that assembly and duct temperatures are evaluated accounting simultaneously for the effect of uncertainties and inter-assembly heat transfer. This represents a marked improvement over the usual method of calculating nominal temperature first and then superimposing the effect of uncertainties, as discussed in Section 3.2.1 of this report. The concept of abandoning the superposition approach is by no means limited to duct temperature calculations and, in principle, it can be extended to all other temperature calculations, thus opening a new area of improvement.

The concept of positive and negative uncertainties in duct temperature calculations was also first introduced in the heterogeneous core analyses reported here. The worst cross-duct gradient in a given assembly would occur when the cdjacent assembly towards the core centerline (thus, next to the hot side of the considered assembly) is at higher temperature than nominal (positive uncertainties), while the assembly away from the core center and next to the cold side is at lower temperatures (negative uncertainties).

Mixed mean temperatures were calculated as for the homogeneous core assuming adiabatic boundaries; however, in the present analysis, these calculations were paralleled by other calculations performed with the TRITON code, where the effect of inter-assembly heat transfer is duly taken into account.

In future analyses the designer will consider, even though not necessarily adopt, such concepts as: a) factoring of uncertainties in all temperature calculations through the integral rather than the superposition approach; b) extension of the positive / negative uncertainties approach to other than duct temperature calculations; c) full consideration of the interassembly heat transfer effect in all of those analyses (e.g., cladding temperature calculations) presently conducted under the assumption of adiabatic assembly l

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boundaries; d) extension of power-to-melt analyses to other rods than hot and

[V] peak which could conceivably be slightly worse due to a particular combination of the various factors; and e) complete implementation of the hot channel f actors and methodology recomended in Reference 3. These are just examples, since new avenues are continuously discovered as the design proceeds. One of these new avenues, which proved to be quite fruitful, became evident when during the present studies a critical look at how to distribute the orificing resistance in radial blanket assemblies between the assembly itself and the lower ' inlet module (LIM) prompted an overall review of the required orificing, which eventually led to mcdified flow conditions. Since no system tem pe ratures , ' s pec i fi ca l ly reac to- inlet temperature, corresponding to the modified reactor flow (and AT) conditions had yet been calculated, most of the studies reported here, and expecially the themal perfomance predictions in Section 4, were conducted for the plant conditions evaluated at the inception of the heterogeneous core analyses. Pressure drop calculations (Section 6) were, however, partially performed for the new, modified conditions. For this reason, plus the very significant effect on the entire core T&H design, and the f act that these conditions will be the basis of the next round of design analyses, they wil; be briefly discussed here.

As previously mentioned, shuffling of radial blanket assemblies was contemplated in the homogeneous design, thus their orificing was located in the LIM; since shuffling was not considered for the hetcrogeneous core, all the orificing was located in the assembly. Even thougn shuffling is not the reference approach, there is no reason why this desigi option should be precluded for the entire life of the CRBRP (the LIM is a pemanent, 30-year component). - Therefore, the radial blanket orificing was apportioned between assembly and LIM to allow shuffling, if so desired at a later time. More importantly, during the homogeneous design, it was decided to size the orificing resistance in zone 1 (highest flow zone) to give a pressure drop of the order of 30 psi. Since, in principle, the zone 1 pressure drop is zero or more realistically, a few psi to allow for final flow trimming, the allowance of 30 psi parasitic pressure drop was designed to accommodate future cores with higher rod bundle pressure drop, e.g., gridded and/or carbide cores.

Analyses of the heterogeneous core showed that the flow requirements were v

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quite similar to those of the homogeneous configuration, while on the other hand, emphasis on the CRBRP was shif ted from providing for insertion of future, hypothetical cores to improving the performance and design margins of the cores actually analyzed and designed. From these premises, it followed that the parasitic pressure drop in zone 1 orifice (and in the entire reactor, since zone 1 is the highest pressure drop flow path) was actually hindering the overall core performance. In fact, once the overall core hydraulic resistance is lowered, the pump operating point will shif t to the right along the characteristic head / flow curve with the end result that the pump will deliver more flow at a lower head. Obviously, this will improve the core performance, since higher flows mean lower temperatures, thus, either longer lifetime and higher burnup (lower fuel cycle cost) or increased margins to limiting conditions (enhanced safety). A computer code, named CATFISH (6) ,

was developed and is now operational. CATFISH is a hydraulic code which models the entire primary tystem; it considers all the hydraulic resistances in the core plus inlet and outlet plenum, primary system piping, check valve, IHX, etc. It also models all the various reactor flow paths (fuel and blanket orificing zones, primary and secondary control assemblies, radial shield assemblies, vessel, leakage), and ties this entire, complex hydraulic network with the pump head / flow characteristit. urve. Thus, for any specified set of resistances CATFISH calculates the pump htad, the total reactor flow, the flow in each assembly and the pressure drop acro.s each subcomponent. CATFISH also has the capability of calculating the above parameters under nominal conditions or accounting for uncertainties, either positive (increased resistanc.:s, which will yield the minimum reactor flow) or negative (decreased resistances, yielding the maximum reactor flow). It was found that by reducing the parasitic orificing pressure drop, and still remaining within the presently specified " window" of pump operating characteristics, the total reactor flow could be significantly incraased, with flow through the fuel and blanket assemblies increasing in the 5-10% range. A 1% increase in flow corresponds to e5 F decrease in maximum cladding temperature, which cnrresponds to an allowable burnup increase of e2500 mwd / ton and fuel cycle cost savings of <$12.5M. Consequently, a 5% flow increas 2 will yield an increased lifetime in each core of approximately 86 full pwer days or 4 calendar months and cost savings of #$60M over the CRBRP lit atime.

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Due to schedular requirements, predictions of the heterogeneous core thermal performance could not include conclusions of the CATFISH code development and of related analyses, which had been conducted in parallel. Additionally, new plant system temperatures were not yet available, as previously mentioned.

Thus, the performance predictions reported in Sections 2 through 5 still reflect the presence of parasitic core orificing resistance and the core assemblies feature lower flows and higher temperatures than otherwise achievable. These predictions envelope the actual operating conditions, once the core hydraulic resistance is reduced and " CATFISH predicted" flows are factored into the design. The fact that the CRi3RP core assemblies still exhibit excellent performance, satisfying all design constraints and limitations as will be shown in this report, points out the soundness of the CRBRP core thermofluids design. Since, however, room exists for significant performance improvement and considerable cost savings, reduction of parasitic resistance and related analyses will be an integral part of the core thermofluids analyses to be conducted for final, as-built conditions. A quantitative comparison of the reactor hydraulics for the "old" and "new" plant flow conditions will be found in Section 6.

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M _2. CORE ORIFICING V

. 2.1 Introduction Cnre' orificing, i.e., flow allocation to the various fuel and blanket assemblies is the single most 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 design parameter which can be varied at will, within certain limits, by the designer, it logically follows that the core T&H design and performance is only as " good" as the core orificing. Therefore, more and more attention in the CRBRP core T&H design has been placed on core orificing, which for the heterogeneous core was the item receiving primary emphasis.

Previous experience has indicated that a successful orificing should account "a priori" for all the various aspects to be considered through the design, in order to avoid time consuming and costly iterations. Thus, going from the homogeneous-to the heterogeneous design, a systematic orificing approach was

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developed, which accounted for lifetime /burnup, transient, upper internals V

temperature constraints. This new approach represented a drastic change in philosophy and quite a significant improvement over the previous and commonly accepted maximum temperature equalization method.

Limiting temperatures were determined (see Section 2.3) for all fuel and blanket assemblies, which were subsequently orificed simultaneously, thus providing the-best flow allocation and utilization. Finally, both first and second core conditions were investigated in determining the orificing constraints and the most restrictive in-either core was used in deriving the orificing configuration. This guaranteed, a priori, that the thermal-hydraulic performance would have satisified the considered constraints in both cores. The following sections, discussing in detail the orificing philosophy, procedure and results, will clarify and enlighten what briefly was discussed above.

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2.2 Orificing Philosophy and Approach The following constraints ( } must be satisfied in selecting the flow O

orificing for the CRBRP fuel, inner blanket and radial blanket assemblies:

e 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).

e Maximum coolant temperature conditions must be such as to assure, with adequate margin, that no sodium boiling occurs during the worst emergency transient (e.g., the three-loop natural circulation event), accountig for uncertainties at the 3a level of confidence.

e Maximum assemblies mixed mean outlet temperature and radial temperature gradient at the assemblies exit must be compatible with upper internals structure (UIS) limitations.

e Maximum of eight discrimination zones (fuel plus inner blanket) are allowed.

e Flow allocation to fuel, inner blanket and radial blanket assemblies 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 fuel enrichment and because the number of required discriminators depends on the unique combinations of flow orificing and fuel 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 fuel assemblies employ identical inlet nozzles. Therefore, both must be considered in determining the total number of discriminator zones. The outer blanket assemblies, due to the high pressure drop requirement across the inlet nozzle, employ a unique inlet nozzle and, therefore, are not considered in determining the total number of discriminators. The row 6 corner positions which alternate between ir.ner blanket and fuel assemblies during successive cycles, form a separate discriminator zone which is included among the eight.

To put the lif etime/burnup and transient temperature constraints on the same quantitative basis, the concept of equivalent limiting temperature is O

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employed. The equivalent limiting temperature is defined as that cladding temperature at a specified radial position (cladding ID in these analyses) and time in life (end-of-life) which must not be exceeded in order to satisfy the considered constraint. 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 for Plant Expected Operating conditions, considering uncertainty factors at the 20 level of confidence, such that accounting for 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 from the above definition, the equivalent limiting temperatures are calculated for each assembly. In f act, all the various assemblies have individually different lifetime histories of cladding temperature and fission gas pressure, and therefore, the limiting equivalent temperatures are necessarily different from assembly-to-assembly to stay within a constraint common to all assemblies. Calculations are perfomed 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. The other set of plant conditions used in thermofluids analyses are plant thermal-hydraulic design

values.(THDV), which are the plant rated conditions and, being more conservative than plant expected conditions, are used in the design of pemanent components and in transient and safety analyses. Orificing of core assemblies is performed on the basis of plant expected conditions.

As mentioned in Section 2.1, both first and second core conditions have been considered in defining the core orificing, therefore, 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 OCTOPUS code (4), 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 n

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least equal to the larger of the flow requirements in first and second core.

This assured meeting of all constraints for both cores. Finally, the excess flow, if any, is supposed to be allocated among the fuel assemblies to minimize and equalize the . ssemblies exit temperature and temperature gradients.

2.3 Calculation of Equivalent Limiting Temperatures Assmblies lifetime /burnup goals are achieved when both the cladding inelastic strain and cladding CDF are within the established limits during steady state operation. 1he ductility strain limit was set at 0.2% and the CDF limit was set at 0.7 in the fuel 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 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 history. Thus, using a preliminary estimate of the assembly flow (but using the proper physics data), the hot rod (*} in each assmbly at end-of-life was identified using the subchannel analysis code COTEC(8) Subsequently, the hot rod so identified was followed throughout lifetime and the lifetime tmperature/ pressure history was calculated with the NICER code I9) . Uncertainty f actors (see Section 3) at the 2a level of confidence were used in the cladding temperature / pressure calculations. Based on the above lifetime histories, a strain equivalent limiting temperature (SELT) and a strain equivalent temperature (SET) were calculated for each assembly. A typical example of these calculations is reported in Table II.

The SET is defined as the end-of-life tmperature which, if maintained constant throughout lifetime, would cause the same end-of-life strain as the actual tm peratuis/ pressure history. The SELT can thus be defined as that SET which would cause, for the particular assembly relative behavior of cladding temperature and pressure through lifetime, and end-of-life cumulative strain of 0.2%. Thus, while the SET depends on the actual temperature / pressure

(*)Each assembly is characterized by its hot rod at end-of-life, which is obviously the one with the highest strain and CDF.

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values, therefore,'on the guessed value of. assembly flow, the SELT does not.

9 depend on the assembly flow, but. rather on the relative behavior through lifetime,-which is only a function of-the power generation changes during life.

_ . Since the DELT is the equivalent end-of-life temperature corresponding to a

.CDF 'of 0.7 'or 0.5, the method employed in -its determination was to extract it from;a curve correlating the cladding ID temperature at E0L with the corresponding CDF. Thus, ~ at least three (in some instances more were

- necessary)' l.i.fetime- temperature / pressure histories were generated for each assembly varying the flow and the corresponding CDF. was calculated. Typical curves f are reported in Figures 2'through 6 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.

'Regarding the transient constraint, the general guideline is to provide adequate margin-to-sodium noiling throughout the assembly lifetime during the worst transient. This was quantitatively translated into a value of 1550 F 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 f actors at the 3a level of confidence. This limit also assumes plant THDV conditions and a 750 F reactor inlet tenperature. The determination of the TELT for each core assembly then proceeded ~ as follows.

. From preliminary transient calculations performed using the FORE-2M-code QO) , the maximum coolant transient temperature at -the worst time in

{ife and the' corresponding steady state coolant temperature were obtained for the worst fuel, inner blanket and radial blanket assembly. As previously mentionec , the- transient considered was the natural circulation event which

' had proved to .be the most severe. Then, the temperature TM was calculated,

'which is' defined-as the maximum steady state coolant temperature at plant expected operating conditions and 2a hot channel factors' corresponding 'to a 1550 F transient maximum-coolant temperature at plant THDV conditions, 30 f"}

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5737A-650Al (S1421)- 14

hot channel factors and 7500F inlet. Calculation of Tg was quite straightforward from ratioing the transient temperature rise calculated by FORE-2M, fixing the maximum transient temperature at 15500F for all three types of assemblies and accounting for the differences in plant conditions and hot channel factors. Table III summarizes the FORE-2M and TM temperatures.

Rigorously speaking, the TM thus calculated are only valid for the three assemblies whose transient behavior was actually investigated; it was conservatively assumed that they apply to all other assemblies of the same t3 Since: a) the value of TM essentially depends on the magnitude of the temperature s ing from steady state to transient; b) higher swing implies higher TM ; and c) other assemblies will have lower steady state-transient swings than the worst representatives investigated through FORE-2M, it follows that adoption of the values of T M from Table III in calculation of the TELT's for assemblies of the same type is conservative.

Finally, using the lifetime temperature profile for each individual assembly, the TELT in each core assembly is calculated by the equation:

(Tg-Tin)

TELT = Tin + (Tcool - Tin}t',x' I cool - Tin)t",x" + (TID - Tcool l t",x" where Tin = reactor inlet temperature in first or second core; Tcool = hot subchannel coolant temperature; TID = hot spot cladding ID temperture t' = time in life when maximum transient temperature occurs; t" = E0L (i.e., EOC2, E0C4 or E0CS);

x' = axial position whe e maximum transient coolant temperature occurs; x" = axial position where maximum TID occurs at EOL.

Again, it should be noted that since the TELT's depend on temperature differences rather than absolute values, lifetime temperature profiles corresponding to first guestimates of assembly flows are perfectly adequate for an accurate evaluation of the TELT's.

O 5737A-650A-(S1421) 15

The SELT, DELT and TELT thus calculated for each assembly are reported in

(

V Figure 7 (first core) and Figure 8 (second core). Since the radial blanket assemblies are refueled only once, their values are reported only in Figure

8. No SElf's are reported for the blanket assemblies, nor for the fuel

~

assemblies in the first core since very little cladding strain occurs, and therefore, their value was very high, thus being not limiting at all.

Actually, an inspection of Figure 8 shows that even in the second core no fuel assembly is strain limited. The limiting constraint in each assembly is given by the ELT with the minimum value. Fuel assemblies are generally CDF limited with the exception of the ones farther from the core center which are transient limited; inner blanket assemblies are transient limited; radial blanket assemblies are CDF limited if flanked by at least two fuel assemblies, otherwise, they are transient limited.

2.4 Results The OCTOPUS code was used to calculate the flow rate needed in each assembly to produce an end-of-life cladding ID temperature equal to the most limiting of the ELT's, as discussed in the previous section. Figure 9 reports such Q flow rates for the first and second cores (cycles' 2 and 4, respectively).

Obviously the minimum flow in each assembly necessary to satisfy all considered constraints in both cores is equal to the most restrictive of the two flow requirements. The higher of the two values reported in Figure 9 is not necessarily the most restrictive. In fact, due to the alternating of fuel and inner blanket assemblies in the six row 6 positions, the total core flow varies slightly from cycle-to-cycle. The replacement of three inner blanket assemblies by fuel assemblies at the beginning-of-cycle 2 results in 0.8% flow reduction (*) in each of the remaining assemblies. At the beginning-of-cycle 4, six inner blanket assemblies are replaced by six fuel assemblies with a 1.6% resultant flow reduction (*) in the remaining assemblies. Thus, when comparing the flow requirements for the first and second cores, minimum flows

(*)This preliminary estimate was subsequently proved to be conservative by CATFISH analyses.

yv 5737A-650A-(S1421) 16 l-l-

1 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 their cycle 4 values, second core requirements are found to be slightly more restrictive in some outer fuel assemblies, as shown in Figure

9. Cycle 5 flows are reported for the transient limited second row radial blanket assemblies, since their TELT's are maximum at EOL.

Using the required minimum flows are guidelines, the OZONE code (which is now incorporated in OCTOPUS as a subroutine) selected, for a given number of orificing zones, that combination of assemblies grouping into orificing zones which among all the various possible combinations, yielded the minimum value of total core flow and was therefore the most effective. As mentioned in Section 2.2, a maximum number of eight discriminators (and orificing zones) is allowed for the fuel and inner blanket assemblies. The number of orificing zones in the radial blanket assemblies was selected as 4 for a total number of core orificing zones equal to 12. The selected arrangement is reported in Figure 10, where the starred assemblies are the ones which determine the amount of flow allocated to the orificing zones (they are actually 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 9).

As shown in Figure 10, the orificing arrangement does not present a 300 symmetry, a direct consequence of the fact that 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 8), first core conditions are limiting for the fuel 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 left.

The minimum amount 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 at cycle 4 conditions. Since 94% of the O

5737A-650A-(S1421) 17

r~'

total reactor flow is allocated to the. fuel and blanket assemblies and since s 93.07% is the minimum required, it follows that slightly less than 1% of the total reactor flow is available to be allocated as deemed necessary by the designer. The OCT9 PUS code is programmed to distribute the excess flow among the fuel assemblies to minimize / equalize the assemblies mixed mean temperature

and temperature gradient. This option, however, was not fully exercised in these studies since the amount of available excess flow is not enough to significantly influence the value of the outlet temperatures. Rather, the excess flow was distributed roughly evenly among the various core orificing zones to provide a little margin above what flow is strictly necessary. The final core flow allocation is reported in Table IV, which shows the cycle-by-cycle variation of flow in the various orificing zones. Both THDV and PEOC flows are reported in Table IV.

G' 4

t__/

5737A-650A-(S1421) 18 u.

L l

'3.0 -UNCERTAINTY ANALYSIS' tt']

V

-This section is divided into two main subsections: Section 3.1 dealing with the uncertainties traditionally adopted in calculation of rod cladding temperature, . assembly mixed mean temperature and gas plenum pressure; and Section' 3.2 dealing with uncertainties for duct temperature calculations which are introduced here for the first time. A' discussion of the uncertainties used -in thennal-hydraulic calculations does not, rigorously speaking, belong

- to a report devoted to describing the thermofluids design and performance predictions, but rather should be the subject of a separate report like Reference ~3. However, the uncertainties are reported in detail here since in one case (duct temperature) they are derived and used for the first time, in the other (calculation of other parameters) it was necessary to clearly spell out which uncertainties are actually used in the calculations.

Uncertainties for power-to-melt calculations are reported in Section 5.2.

3.1 Rod and Mixed Mean Temperature Uncertainties O

+ 1 V 3.1.1 Intrnduction During the previous thermal-hydraulic analyses conducted for the homogeneous core (2) , a preliminary study was performed to assess the impact of theoretical and experimental analyses uncertainties, instrumentation accuracy, manuf acturing tolerances, and physical properties and correlation uncertainties on the core thermal-hydraulic performance predictions. Hot channel / hot spot factors for fuel and radial blanket assemblies were determined to account quantitatively for the above uncertainties and semi-statistically combined. Separate groups of uncertainty factors were established for calculation of: a) local rod temperatures (coolant, cladding, fuel); b) assembly coolant mixed mean exit temperature; and. c) rod fission gas 1

plenum temperature ~ and rod burnup (which together with uncertainties on

' fission' gas yield, fission gas release and plenum volume, determine the

- overall' uncertainty on plenum pressure).

m l 5737A-650A-(S1421) 19 r

A set of uncertainty f actors for prediction of the heterogeneous core thermal-hydraulic performance was established along the same lines and is presented in

. Tables V through X. For convenience, the factors have been tabulated by assembly type (fuel, inner / radial blanket), intended application (rod temperatures, mixed mean temperatures, plenum pressure calculations), and type (engineering - table numbers with the "A" suffix; nuclear - table numbers with the "B" suffix). Note that 3o statistical values are presented which are the basis for transient and safety analyses. Uncertainties at 2a level are used in calculation of steady state thermal-hydraulic parameters, such as cladding temperature and pressure, which are input to replaceable core assembly lifetime analyses.

As previously mentioned, these factors represent an adaptation to the heterogeneous core configuration of the hot channel factors adopted in the thermal-hydraulic design of the homogeneous core. Of course, nuclear uncertainty factors were, however, properly determined for the power and burnup his'.ories characteristics of the heterogeneous core (ll) The significant features in the application of the nuclear uncertainty factors to the thermal-hydraulic analyses are reported in Section 3.1.3.

In parallel with the studies reported here, a thorough reevaluation of O

uncertainties analyses, their numerical values and methodology of application was performed. This reevaluation led to a revised set (3) , which was approved by the CR8RP Project Office in December 1978. It must, therefore, be emphasized that the performance predictions herein contained are not consistent with the uncertainty factors documented in Reference 3, but are actually more conservative. In fact, temperatures calculated using the f actors and method of Reference 3 are lower (*) than the values reported in Section 4, which were predicted utilizing the hot channel factors from Tables V through X.

(*) Calculations for a typical fuel, inner blanket and radial blanket assembly hot rod, 20, plant expected conditions at end-of-life showed the present analysis yielding a maximum cladding ID temperature approximately 540F, 250Fand80F,respectively,higherthanobtainedusingReference3 uncertainty f actors and analyt. cal procedures.

5737A-650A-(S1421) 20 9

.f~N)- The major differences between the engineering uncertainty factors adopted here t

U for rod temperature calculations and those recomended in Reference 3 are briefly sumarized in Section 3.1.2. Uncertainties applied in power-to-melt analyses are consistent with Reference 3 and are discussed in Section 5. A discussion of how uncertainties are conservatively accounted for in plenum pressure calculations is reported in Section 3.1.4.

3.1.2 Rod Temperature Engineering Uncertainty Factors The hot channel / hot spot factors representing engineering uncertainties in rod temperature calculations and adopted in these analyses differ from those recomended in Reference 3 as follows:

1 e pellet-cladding eccentricity hot channel factor. Both a direct and statistical component were evaluated in Reference 3 as well as a lifetime dependency, while only a statistical component unchanged throughout~ life was considered here. Note that the latter is conservative when calculating end-of-life temperatures, whicie are the ones of utmost importance in steady state analyses, since Reference 3 recommends to O consider no end-of-life uncertainty factor, neither direct nor statistical, due to pellet-cladding eccentricity; e assembly inlet flow maldistribution factor. Detailed analyses conducted in Reference 3 recomended a split of this factor into a direct and a statistical component, in lieu of the entirely. direct factor adopted here; and again, the latter approach is more conservative; e subchannel flow area. A reevaluation of the analyses performed to establish this factor led to a reduction in its value; e cladding circumferential temperature variation. Recent modifications to the FATHOM-360 code (12) allowed calculation

.in Reference 3 of individual hot spot factors representing separately the effect of the wire wrap on the film and cladding temperature drop. Previous practice was to consider l only one factor to be applied to the temperature difference between the bulk coolant and the radial cladding position considered. Of the two factors reported in Reference 3, the cladding hot spot is less than unity, as it physically should f be, since it represents the beneficial effect of circumferential heat transfer through the cladding. The film hot spot, on the other hand, accounts for the presence of the i (%. .

\v]

5737A-650A-(S1421) 21

wire wrap in the channel. The approach recommended in Reference 3 is, therefore, more adherent to reality than used here, however, the two methods yield the same overall effect on cladding temperature, in f act, the overall f actor reported in Tables V-A and VIII-A applies to the total temperature drop through the cladding and film; the Reference 3 approach features a higher hot spot f actor to be applied to the film AT alone, which is balanced by a less than unity hot spot f actor to be applied to the cladding AT; e cladding thickness and conductivity. A critical reevaluation conducted in Reference 3 of the correlation adopted in design, its deviation f rom experimental data and the effect of irradiation led to the conclusion that the 5% under-estimation directly built in the correlation is more than adequate to account for irradiation swelling and cladding thickness effects on temperatures and, therefore, no hot channel f actor is recommended in Reference 3. Again, the present analysis is conservative by adopting a 1.12 hot spot f actor; e fissile fuel maldistribution. The value recommended in Reference 3 (1.052) is higher than the value used here (1.03),

by a f actor of ,G; this is due to the use of a rectangular distribution, rather than nonnal as adopted in the preliminary hot channel f actors assessment; e reactor AT and inlet temperature variation. This is discussed in Section 4.1; e flow distribution calculational uncertainty. There are several differences between the values adopted in this study and those recomended in Reference 3. In the fuel and inner blanket assemblies case, the values used for rod temperature calculations are similar, while the values for plenum tenperature calculation adopted in these studies were conservatively assumed to be higher. In the radial blanket case, the values both for rod and plenum temperature calculations, adopted in these studies are lower than recomended in Reference 3. These discrepancies are attributable to the f act that comparison / calibration /

verification of the COTEC code is an on-going task as more data become available. For example, since issuance of Reference 3 and inception of these studies, data have been obtained from the ARD full scale 61-rod blanket heat transfer test in sodium for flow and skewed power conditions prototypic of the radial blanket assemblies.(13) Comparison of experimental data with COTEC predictions indicated that COTEC consistently predicts higher temperatures, both at the hot rod maximum cladding temperature axial position and in the unheated plenum region, than obtained in the tests.

Preliminary evalur.cion of this data justify reducing the maldistribution calculation uncertainty f actor for the radial 22 O

5737A-650A-(S1421)

fy  : blankets from 1.2 (rod temperature) and 1.3 (plenum 3 '

) temperature) to 1.1 for both. A conservative approach was also taken for the inner blanket; since experimental data for flatter power distributions typical.of inner blanket assemblies were not yet available, the higher uncertainty factors were retained in this study, even though every indication exists that they can be reduced.

As mentioned, elaboration of data from the blanket heat transfer test was concurrent with this analysis; thus, determination of limiting temperatures and flows as discussed in Section 2 was based on the higher value (1.2) of the maldistribution calculational factor. Since such a procedure is conservative, resulting in lower limiting temperatures, a reevaluation of the orificing calculations was not warranted, but rather the updated value-was adopted in the performance predictions reported in Section 4.

Neither in Reference 3 nor here have results been incorporated from 61-rodthesaltrecently obtained injection fue(l in water bundle 14), test 61-rod the KFK data from the MIT heating in sodium (15) and the J0YO 91-rod heating in sodium (16),

Additionally, the 5:1 air ficw tests on a blanket assembly conducted at the Westinghouse Research Laboratory (17) have just been completed, the ARD 61-rod bundle blanket heat transfer test is in progress and the ORNL 61-rod fuel assembly bundle sodium heat transfer test is planned in CY 1979. As these additional test data are factored into the code

~V calibration, the basis for the statistical component of the flow distribution calculational uncertainty will be improved as well as the direct component (simulation bias) will be reduced or eliminated altogether since the forthcoming data are'from larger bundle, more prototypic heated rod experiments in sodium.

In conclusion, as previously mentioned, the overall effect of using the hot channel factors in Tables V through X rather than those recommended in Reference 3, is a more conservative evaluation of the core steady state thermal performance reported in Section 4.

3.1.3 Nuclear Uncertainties

\

The total power or burnup uncertainty is composed of nuclear design methods uncertainties and/or biases (based on comparisons of calculations and

-measurements of isotopic fission and capture rates and gamma-heating in ZPPR f criticals), plus CRBR design uncertainties relating primarily to absolute power normalization and fissile content variations, and a general class of

~C-l 5737A-650A-(S1421) 23

l l

l modeling uncertainties. In the fuel, the power uncertainty is broken down into a statistical part which can be cambined in quadrature (root-mean square) with other statistical uncertainties, and a non-statistica'i bias and uncertainty which is applied directly to envelope the upper limits of the peak power density. Due to the limited scope of the available blanket data, only a non-statistical uncertainty is developed. Uncertainties are provided for the fuel, inner and radial blanket assemblies. Where a basis exists, a spatial distribution of the uncertainty is provided (e.g. , adjacent to, and removed from the influence of inserted control rods, and by assembly row in the outer blanket). Otherwise, the uncertainty is developed for the peak power locations and should be assumed to be applicable throughout the region for the stated time-in-life. For a detailed discussion, and definition of all the nuclear power uncertainties and individual factors, see Reference 11.

Of particular interest to the thermal-hydraulic design is the spatial dependence of the uncertainties:

a. Radial (i.e., adjacent to, and removed from the influence of an inserted control rod; locally dependent bias due to ZPPR-7 flux tilt; and by assembly row in the outer blanket),

b. Axial (i.e., at the peak power density near the core midplane, and at the core / upper axial blanket interface).

In the case of the radially distributed uncertainties, how and where the uncertainties are to be applied is shown in Tables V through X with the B suffix. The axial uncertainties are a result of the nuclear design calculations and modeling techniques. In particular, the peak and integrated power densities in the fuel are well predicted with the standard two-dimensional synthesis nuclear design techniques, while the " power density at the top of the core" is relatively poorly predicted due to difficulties in simultaneously modeling the behavior in the region while preserving the integral and peak (core midplane) power in two dimensions. In addition, the accuracy of few group diffusion theory is poorer in the presence of the steep 24 O

5737A-650A-(S1421)

(] flux _ gradient and in the region of the fuel / upper axial blanket material discontinuity. These f actors are reflected in the larger " power density at top of core" uncertainty.

The temperature distribution along the entire fuel rod is necessary for the evaluation of relevant thermal-hydraulic parameters (i.e., fuel temperature, cladding temperature, fission gas generation and release, etc.). As noted above, only two values of the spatially-dependent (in the axial direction) factors are provided. Therefore, it was necessary to make several assumptions and approximations. First, it was assumed that the value of uncertainties at the fuel / upper axial blanket (UAB) interface also apply at the fuel / lower axial blanket (LAB) interf ace. The actual shape of the axial variation of the heat flux uncertainty from the peak power position to the fuel /UAB and fuel / LAB interf aces is generally not a very critical item since: a) it does not affect the channel enthalpy rise (an uncertainty integrated over the rod length is provided); b) the heat flux is specified at those positions of most interest (i.e., top of the core for cladding temperature calculations, peak powr position for fuel temperatures). Thus, the most ininediate assumption, (v }j

/

i.e., a linear variation in both directions, was adopted. This was also a conservative assumption for plenum pressure calculations, as discussed in Section 3.1.4. In power-to-melt calculations it was, however, discovered (see Section 5.3) that due to the effect of cladding swelling, which is extremely sensitive to the cladding temperature value, the fuel was closest to melting not at a location corresponding to the peak power position. A more realistic definition of the heat flux uncertainty axial shape was therefore necessary.

An inverted chopped cosine curve fitting the heat flux uncertainties at the peak power and interf aces positions was therefore adopted in predicting the axial profiles of rod temperatures necessary as input to power-tc-melt evaluations. For both assumed distributions, linear and inverted chopped

cosine, the local values of the heat flux factor are routinely calculated by NICER.

Finally, the uncertainties at the fuel /UAB and fuel / LAB interfaces were assumed constant over the respective axial blankets.

[

v i 5737A-650A-(S1421) 25 l

3.1.4 Plenum Pressure Uncertainties The fission gas plenum pressure in CRBR fuel and blanket rods is calculated from the perfect gas law:

nRT p=

The physical parameters affecting the pressure value are, thereforeI*):

e fission gas generation, which depends on the rod burnup; e fission gas yield, which depends on the type of atom fissioned; e fission gas release to the plenum, which depends on the rod burnup, linear power rating and temperature; e plenum tmperature; and e p1enum voItune Uncertainties on the above parameters are accounted for as follows:

e uncertainties on burnup (see Tables VII and X; 2o level of O

confidence is adopted for performance calculations reported in Section 4);

e the adopted value for the fission gas d, regardless of the Recommended values are:

fissioned 0.2554 atm$,is0.2385 for U23 0.266.for U238, and 0.2506 for Pu239; e a comparison between fission gas release predictions and experimental data from EBR-II irradiation experiments was perfonned for the fuel assemblies; the +2c release correlation was adopted in the analyses. In the radial blanket case, since no pertinent experimental data are available, nominal fission gas release predictions fra the fuel assemblies calibrated model were increased by 15%.

e uncertainties on plenum temperature (20 level of confidence is adopted - see Tables VII and X); and

(*) Discussed here is only the component due to fission gases, which is by f ar overwhelming, especially at E0L. The pressure c m ponent due to initial gases is cmputed separately and added to the fission gas component to yield the total plenum pressure.

5737 A-650A-(S1421) 26

e minimum plenum volume was adopted in the calculations.

v It is evident from the above that while the resulting fission gas pressure is labeled as "20", in reality the level of confidence is much higher. In fact, 2a uncertainties on the various parameters are superimposed on each other; the uncertainty on.the plenum volume is a bounding value rather than a 2a; and, 238 the adopted. uncertainty on fissinn gas yield is much higher for U and 239 than for U235 (which is only a small fraction of the fuel, even more Pu so at E0L when the plenum' pressure value is most important and critical).

Predictions of fission gas plenum pressure have additional conservatism in the blanket assemblies case since: a) the same release model as for the fuel assemblies is assumed, which is by preliminary indications conservative; b) the release so calculated is multiplied by a 1.15 f actor, thus very closely approaching 100% release. Assumption of a linear, rather than inverted chopped cosine heat flux uncertainty distribution (which affects the fuel temperature, hence the gas release) as _ discussed in Section 3.1.3 is also conservative, since it tends to increase the fuel temperature and gas release in the central region of the rod where most of the gases are released.

(V)

A more proper and realistic way of calculating the 2a plenum pressure would be to individually vary each of the affecting parameters by the corresponding 2o

-level of uncertainty and calculate the induced change in plenum pressure.

Adding the root-mean-square of the various changes in plenum pressure to the 0a value will yield the true 20 value of the fission gas pressure. This calculation was perfonned for the hot pin in inner blanket assembly #99 and radial blanket assembly #201 at end-of-life (E0C4) to quantify the margin of conservatism implicit in the current evaluation.of plenum pressure. Factored

-in this analysis were also the very recently obtained fission gas release data from WBA-20 (the first fission gas release data for blanket assemblies) which indicated a very substantial over-estimation of fission gas release by the present high burnup fuel calibrated LIFE-III code model. It was found that for. the inner blanket rod using the r.m.s. method, the calculated 2a total plenum pressure was 175 psi, while the value used in design by accounting simultaneously for all uncertainties was 249 psi. For the radial blanket pin, the calculated 20 pressure was 188 psi versus a design adopted value of 273 m

I (v)

'5737A-650A-(S1421) 7' E l

psi. For comparison, the total plenum pressure assuming 100% fission gas release and simultaneously accounting for all the other uncertainties was 264  !

psi for the inner blanket and 310 psi for the radial blanket. Thus, the design adopted value is quite comparable with the 100% release value, a direct consequence of the over-prediction of the fission gas release model and the additional 15% increase. The above examples therefore show quite clearly the large amount of conservatism implicit in the plenum pressure values used throughout these analyses.

3.2 Duct Temperature Uncertainties 3.2.1 Introduction As mentioned, the core thermofluids analysis of the heterogeneous core represented the first instance where a significant effort was attempted to systematically define duct temperature uncertainties. They closely resemble and are consistent with the uncertainties discussed in Section 3.1 rather than the much more elaborate set and procedure developed in Reference 3. In addition to the fact that the latter was not yet available for design use when these analyses were being conducted, past history has taught that a fair amount of design experience is necessary before attempt.ng a very elaborate and much more complicated approach. Before future analyses are undertaken, the uncertainties proposed here should, therefore, be critically reviewed and eventually updated along the same lines as for rod temperature calculations for final as-built analyses. Since characterization of the thermal-hydraulic behavior of the control assemblies is necessary when calculating duct temperatures in surrounding assemblies (inter-assembly heat transfer is very pronounced at the control assemblies interfaces), uncertainties are reported here for all three types of core assemblies; i.e., fuel, inner and radial blanket and primary control assemblies. Due to the lack of data for the secondary control assemblies it was assumed that the same uncertainties as for the primary control assemblies apply.

Uncertainties are broken down in factors affecting the nominal values of assembly heat generation, assembly flow, coolant enthalpy rise, film heat transfer coefficient, duct and interstitial gap thermal conductivities, duct O

5737A-650A-(S1421) 28

4;'

thickness and. interstitial gap geometrical variations. In the control

- assembly, separate enthalpy rise uncertainties are' calculated for the bundle

- and the. bypass and additionally an uncertainty on the bundle / bypass flow split is considered.

The most recent version of the TRITON code is capable of calculating duct I' temperatures throughout the core under nominal and "with uncertainties" conditions.

A. major difference in accounting for uncertainties between duct temperature calculations and all the other calculations reported in Section 3.1 should be pointed out. In the latter, the nominal temperature differences (AT) across each caponent-are multiplied by the hot channel / hot spot f actors (super-position approach). In duct temperature calculations, a new and more realistic approach is adopted: the various uncertainties are input to TRITON and energy equations are solved accounting simultaneously for the effect of uncertainties and inter-assembly heat transfer (integral approach). The

! superiority of the integral approach in duct temperature calculations, where (n) inter-assably heat transfer is of paramount importance, is evident. Since

' TRITON also has the capability to calculate cladding temperatures, in future analyses the corservatism implicit in the superposition approach adopted for rod temperature calculations can be quantitatively evaluated by comparing tmperatures calculated by TRITON and current temperatures calculated via COTEC/ NICER without accounting for inter-assembly heat transfer.

Another major feature of this analysis is that both positive and negative uncertainties are considered, i.e., uncertainties either increasing or decreasing the nominal AT's. In f a;*, cross duct gradients are magnified when the considered assembly is flanked on the hot side (towards the core

centerline) by an assembly with positive uncertainties, i.e., running at i

higher tmperature than nominal, and on the cold side by an assembly with negative uncertainties and running colder than-under nominal conditions.

Caution must be exercised in calculating the values of_ the uncertainties when both positive and negative uncertainties in the same core are considered (e.g., fuel assemblies positive, blanket negative). In f act, once a certain if . %I uncertainty is selected, its value might remain frozen; for example, if the

7-5737A-650A-(S1421) 29

- * '" T k

w w - -t - -----vs- y+' r r C'T' W 5 7 j'

uncertainty on stainless steel conductivity is selected as positive in fuel assemblies calculations, it is obviously illogical to take a negative value in blanket assemblies calculations. In this analysis, it was decided that positive uncertainties have first priority, which is equivalent to say that absolute level of temperature is of more concern than relative gradients.

Consequently, while the positive uncertainties are all positive indeed, the r.egative set is a mixture of negative and positive values. Again, the concept of negative uncertainties is a novel one, with the potential for significantly impacting the overall core design, of which duct temperature is only an example.

Following are the values for the various uncertainties and their rationale and justification; they are grouped by assembly type (refer for a summary to Tables XI through XV).

l 3.2.2 Fuel Assemblies 3.2.2.1 Heat Generation Calculations in TRITON are performed using individ"al assembly powers directly multiplied by the heat generation uncertainty. Thus, each assembly in the core can have a different uncertainty, but within the assembly all the rods are assumed to have the same heat generation uncertainty. Therefore, uncertainties recommended (II) for assembly power are adopted. There are five different groups of assemblies in respect to the uncertainty associated

< with their power evaluation:

e Group I - Assemblies adjacent to inserted Row 7 corner control rod with a 3% bias at 80C due to ZPPR-7 flux tilt; e Group II - Assemblies adjacent to inserted Row 7 corner control rod w h a 1% down bias at 80C due to ZPPR-7 flux tilt; e Group III - Assemblies not affected by control rod but with a 3% ZPPR-7 bias; e Group IV - Assemblies not affected by control rod but with a 1% ZPPR-7 bias; and e Group V - Assemblies neither affected by control rod nor by ZPPR-7 bias.

30 O

5737A-650A-(S1421)

Going from positive to negative uncertainties, the following subfactors V maintain their initially chosen (positive)(*) value:

o Control Rod Banking - Since this factor affects the whole core, obviously it can have only one sign, either positive or negative; e ZPPR-7 Flux Tilt - This is a systematic bias obtained by comparing calculated values with ZPPR-7 data. By its very own nature, the same positive / negative consideration above applies; e Criticality - This uncertainty characterizes the prediction of the hot critical state of the reactor, thus the control rod insertion, thus the_ power distribution. Since it is related to the whole reactor, it cannot have locally variable values.

The various subfactors comprising the heat generation nuclear uncertainty are reported in Table XII. Except for the negative uncertainties and a different format, they are the same as in Table VI-B.

In addition to the nuclear uncertainties discussed above, a direct factor must n be considered to account for variations in power level measurement and control b system dead band. A value of 1.03 was used, as per Table V-A. This reactor-dependent factor remains at the same value for both positive and negative uncertainties.

3.2.2.2 Assembly Flow The individual assembly flow is divided in TRITON by this uncertainty. Among the various subfactors which affect the coolant enthalpy rise, the only one physically related to the amount of coolant actually flowing in the assembly is the inlet flow maldistribution. The value (1.05) from Table V-A is also

-adopted here. The negative uncertainty is the reciprocal, i.e., 0.95. This factor is treated as direct.

(*)The- reason why positive uncertainties are the primary choice is detailed in Section 4.5.

(a 5737A-650A-(S1421) 31

3.2.2.3 Coolant Enthalpy Rise (AH)

This uncertainty comprises all the subf actors which affect the coolant enthalpy rise, but are not directly attributable to assembly power or flow.

The AH uncertainty f actor is applied in TRITON to the temperature change over each axial calculation increment to provide imediate feedback and avoid calculaticnal instabilities. Subf actors included in the AH uncertainties are discussed below.

Flow Distribution Calculational Uncertainty (Simulation Bias)

The same value as recomended in Section 3.1.2 is adopted. Being a simulation bias, the negative uncertainty has to be the same as the positive.

Reactor AT Variation The same values as in Table V-A are adopted. Since this is a reactor-dependent uncertainty, there is no distinction between positive and negative value.

Wire Wrap Orientation This subf actor accounts for the f act that due to the swirl flow induced by the wire wrap in peripheral channels, the flow and temperature distribution in the assembly depends slightly .1 the relative orientation of the wire wrap and the power skew. Since there is no rotational discriminator in CRBR assemblies, six TRITON runs were conducted for a typical fuel assembly varying in steps of 60 the position of the wire wrap from the reference position (0 ) used in all TRITON analyses. The average coolant AT (at several axial positions) in the side channels of the hottest and coldest f ace for the five rotated wire wrap cases were raticed to the corresponding AT for the reference case. This dimensionless f actor ranged from 0.981 to 1.022 for the hottest f ace and f rom 0.978 to 1.017 for the coldest f ace, with an average of 1.01 and 0.99, respectively. These were the values adopted as positive and negative uncertainty.

O 5737A-650A-(51421) 32

-[\ 3/ Subchannel Flow Area This subf actor accounts for variations in side channels temperature due to geometrical tolerances and bundle ' displacement. An analysis was conducted, using a COBRA-IV across' assembly strip model for a typical fuel assembly, under nominal conditions as well as the following perturbed condition:

assembly of minimum diameter rods, compacted to the minimum pitch in a duct of maximum across flats dimensions, with the bundle moved to contact the duct at tne flat side towards the core centerline. A ratio of the side average coolant AT under perturbed conditions to the nominal case will yield the positive uncertainty when the f ace towards the core centerline is considered (flow channels area reduced to a minimtsn on the hot side) and the negative uncertainty when the 180U opposite face is considered (flow channels area increased to the maximum on the cold side). As in the case of the wire wrap orientation f actor, several axial locations were examined and the dimensionless f actors varied slightly with axial position. Values adopted were average in the top of the core region which is the area of most interest for the- core restraint design.

U. Flow Distribution Calculational Uncertainty (Calibration)

This subf actor accounts for deviations observed between experimental data and the corresponding predictions by the subchannel analysis code COTEC, which is the basic building block of TRITON. As mentioned in Section 3.1.2, calibration and comparison of COTEC against experimental data has been an ongoing task in the past few years, as more data became available. The value (1.08) selected for duct temperature calculations is somewhat larger than the one used in rod tenperature calculations (1.054 - Table V-A). In fact, most of. the data available (and all the theoretical elaborations of data), are for inboard channels, while side channels are of primary importance in duct temperature calculations, and therefore, the statistical uncertainty would be higher for side than for inboard channels due to the paucity of pertinent data and absence of analyses. As more detailed analyses, as

/^%

b v

5737A-650A-(S1421) 33

discussed in Section 3.1.2, are performed, the flow distribution calculational uncertainty (both the direct simulation bias and the statistical calibration components) will be reassessed.

When going from positive to negative uncertainties, this being a statistical calibration factor, it is quite legitimate to assume that the negative uncertainty is symmetric to the positive value, i.e., 0.92. However, caution should be exercised regarding how the negative uncertainty is used; once a poisitive or negative uncertainty is selected, that value applies to all the fuel assemblies. On the other hand, nothing in principle would preclude to having, e.g., a positive uncertainty in the fuel assemblies and a negative uncertainty in the blanket or vice versa. The point is that if it is decided to maximize differential gradients, then opposite uncertainty signs can be taken when different types of assemblies are involved, but such a procedure is illogic when considering the same type of assembly.

Coolant Properties This factor accounts for uncertainties in sodium coolant properties (density, enthalpy, conductivity). The value recommended in Reference 3 is adopted.

3.2.2.4 Film This uncertainty comprises the various subf actors which affect the value of the film heat transfer coefficient between side channels and duct wall. These factors are: a) variations in 1.he channel coolant flow rate due to pertinent uncertainties (inlet flow maldistribution, flow distribution calculation uncertainties) cause a variation in the Reynolds number, hence in the Nusselt number, hence in the convection heat transfer coefficient; and b) discrepancy between the film heat transfer coefficient correlation adopted in TRITON and experimental data.

The film coefficient correlation adopted in TRITON is:

Nu = 5.8 + 0.02 Pe 0.8 I

O 5737A-650A-(S1421) 34

which applies to sodium flow between parallel plates. The side channels Q geometry is not exactly representable by parallel plates; additionally,

, experiments are conducted for either constant temperature or constant heat flux, neither of which represents the side channels. Finally, no sodium heat transfer data exist yet for side channels. Since the contribution of the film resistance to the overall resistance between the side channels-of two adjacent assemblies- is negligible when compared with the ducts and gap resistance,

~

during the TRIT 9N development it was concluded that an extensive effort in securing a more accurate representation of the film heat transfer was not warranted. From the above premises, it follows that a sophisticated analysis of the film heat transfer coefficient uncertainty is completely meaningless.

It was, therefore, decided that in TRITON uncertainty analyses the valuc or the film coefficient will be bracketed by the two extremes: for positive uncertiinties (maximum duct temperature ) h = k/6 where k is the sodium conductivity and 6 is the distance between the side channel centroid and the duct wall; for negative uncertainties (minimum duct temperature) h = =, i.e.,

) the duct ID temperature is taken equal to the side channel bulk coolant G temperature.

t

.J 3.P.2.5 Duct

. Two subf actors directly affect the tempe. ature drop through the duct:

geometrical- deviations and thermal conductivity of the duct. They are treated differently in TRIT 9N as discussed in the following.

J Geometrical Deviations Variations in the assembly pitch and duct thickness are accounted for in TRIT 9N by physically changing the value of these dimensions during the calcul ations . This approach is conservative, since it actually treats these uncertainties as direct, rather than statistical. TRIT 9N .has the capability of varying pitch and duct thickness independently for each pair of interf acing assemblies (i.e., in a cluster of seven assemblies, the user can specify seven different duct thicknesses and six different assembly-to-assembly pitches).

~

5737A-650A-(S1491) 35

When considering positive uncertainties, dimensions leading to the highest duct temperature are considered, i.e., maximum duct thickness and inter-assembly pitch; for negative uncertainties, minimum duct hiickness and pitch will yield the lowest duct temperature.

Duct Thermal Conductivity The nominal duct (316SS) thermal conductivity is divided in TRITON by the corresponding uncertainty f actor. This f actor accounts for deviations in experimental data and irradiation effects. A 10% value was adopted, consistently with the value reported in Table V-A, where out of the 1.12 factor for cladding thickness and conductivity, variations in cladding thickness account for the balance.

3.2.2.6 Interstitial Gap Subf actors affecting the temperature drop through the interstitial gap between ad,)acent assemblies are geometrical variations and either uncertainties on the film heat transfer coefficient (if flowing interstitial sodium is considered) or uncertainties on the sodium thermal conductivity (if stagnant sodium is assumed). They are discussed in the following.

Geometrical Deviations Variations in the interstitial gap due to tolerances are treated in TRITON I

similarly to variations in duct dimensions (see Section 3.2.2.5). Once the assembly pitch and the cucts thickness are specified, the gap thickness is also necessarily fixed; thus, the previous discussion applies.

Film Heat Transfer Coefficient TRITON has the option of conside 'ing either flowing or stagnant sodium in the gap. However, a numerical instability developed for sodium flows of the order of 200 lb/hr, an instability which has not yet been corrected. All TRIT 9N analyses are therefore currently performed under the conservative assumption O

5737A-650A-(51421) 36

p .of stagnant sodium. Thus, 'this uncertainty does not apply at present. When V

the option of flowing sodium is fully operational in TRITON, the uncertainty factor affecting the heat transfer coefficient in the gap will-be evaluated.

Coolant Properties This uncertainty applies only when the sodium in the gap is assumed as stagnant. TRITON divides the sodium conductivity by the corresponding uncertainty f actor. A value of 1.017 was assumed as discussed in Section 3.2.2.3.

3.2.3 Inner and Radial Blanket Assemblies The same philosophy as discussed in detail in Section 3.2.2 for the fuel assemblies applies to the blanket assemblies. Therefore, only the areas in which the blanket assemblies differ from the fuel assemblies evaluation will be discussed, f) .3.2.3.1 Heat Generation o

The major difference in nuclear uncertainties evaluation with respect to the fuel assemblies is that rod power rather than assembly power uncertainties are adopted. The reason is that modeling uncertainties in power predictions of blanket assem5 lies show a much more pronounced local dependence, and therefore, adoption of assembly power uncertainties would not have been sufficiently conservative.

Different nuclear related uncertainties are calculated for inner blanket, radial blanket first row, radial blanket second row at beginning and end-of-life. All uncertainty factors are treated as direct. Priitive and negative uncertainties are calculated directly from Reference 11. The

" control rod banking" and " criticality" uncertainties do not change from positive to negative, as discussed in Section 3.2.2.1. The various subfactors comprising the heat generation nuclear uncertainties are reported in Table  ;

XIV. In addition, the power level measurement and control system dead band l g_ factor must be censidered as discussed in Section 3.2.2.1.

Lt i

]

I v

5737A-650A-(S1421) 37 l

l l

t

3.2.3.2 Assembly Flow A 7% variation in assembly flow due to inlet flow maldistribution is adopted, O

as per Table VIII-A.

3.2.3.3 Coolant Enthalpy Rise (AH)

The same discussion as in Section 3.2.2.3 applies for the flow distribution calculational uncertainty (simulation bias), reactor AT variation, and coolant properties subf actors.

Wire Wrap Orientation The same analysis as for the fuel assemblies was performed for the blanket assemblies. Slight deviations between the reference and the rotated cases were found at an elevation of 50", with a positive and negative uncertainty f actor of 0.98 and 0.96, respectively. However, at elevations of 64" and 112" there was no difference in the side temperatures predicted for different rotation of the assemblies. Thus, it was concluded that this effect is negligible in blanket assemblies and no uncertainty f actor was considered.

Subchannel Flow Area An analysis similar to the one perfonned for the fuel assemblies resulted in a positive f actor of 1.15 and a negative f actor of 0.75. The negative factor is somewhat conservative since it is slightly lower than values calculated at 50" and 64".

Flow Distribution Calculational Uncertainty (Calibration)

For added conservatism, the same value (rounded to 1.2) as for rod temperature calculations in inner blanket assemblies is adopted here for both inner and radial blanket assemblies. Regarding the negative value, the same precautions as previously discussed apply, i.e., once a sositive or negative uncertainty l

is chosen, it applies to all blanket assemb,ies.

O 5737A-650A-(S1421) 38

{J '3.2.3.4 Film See corresponding Se: tion 3.2.2.4.

3.2.3.5 Ouct See corresponding Section 3.2.2.5.

3.2.3.6 Interstitial G @

See corresponding Section 3.2.2.6.

3.2.4 Primary Control Assemblies Uncertainty analyses for primary control assemblies rod temperatures calculations have not been conducted to the same level of depth and detail as for the fuel and blanket assemblies (in fact, control assembly uncertainty factors reported in Reference 3 were obtained in the only analysis performed

, so far, which referred to the homogeneous core). Thus, treatment of the

~

control assemblies uncertainties is generally less detailed and sophisticated than for the fuel and blanket assemblies; see, for example, nuclear and flow distribution calculational uncertainties. The same situation of course applies to the duct temperature uncertainties analysis, since it relies on analyses previously conducted for rod temperature uncertainties.

As was the case for the blanket assemblies, only those uncertainties where a significant difference exists in respect to fuel assemblies analysis will be discussed in the following. The most significant feature characteristic only of the control assemblies is the need for additional uncertainties on the bundle / bypass flow split and on the bypass enthalpy rise. Regarding the latter, it is obvious that both inner and outer duct temperatures depend on the amount of flow and enthalpy rise in the bypass; on the other hand, the enthalpy rise in the bundle and the bypass are affected by different types of uncertainties, thus, the need to handle'them separately.

! O U/

'5737A-650A-(S1421) 39

(

- -- , - , e ., - - - - e , - - . , , . - 7

3.2.4.1 Heat Generation Nuclear uncertainties are as recommended in Reference 3. The uncertainty on O

power level measurement and control system dead band is obviously the same as for the fuel and blanket assemblies.

3.2.4.2 Assembly Flow The inlet flow maldistribution uncertainty is from Reference 3.

3.2.4.3 Bundle / Bypass Flow Split An additional uncertainty has to be considered for the control assemblies, i.e., the one affecting the flow split between absorber bundle and bypass.

The flow split predicted by the CRAB code (19) is input to TRITON; a comparison between CRAB predictions and HEDL experimental data for the FFTF 61-pin absorber assembly showed a maximum over-prediction of the bundle flow fraction by CRAB of 6%. To accour,t for additional uncertainties for a 37-pin bundle, a 10% positive uncertainty is considered. Since the CRAB code over-estimates the amount of flow through the bundle, the negative uncertainty is taken as unity.

TRITON calculates the nominal bundle flow as the total assembly flow multiplied by the nominal flow split. If uncertainties are accounte6 for, then the following calculations are performed: a) the nominal flow split is divided by the flow split uncertainty; b) the total assembly flow is divided by the assembly flow uncertainty; c) the bundle flow is equal to the total flow (with uncertainties) from (b) multiplied by the flow fraction (with uncertainties) from (a); and d) the bypass flow is equal to the total assembly flow from (b) multiplied by the difference to unity of the flow fraction from (a).

O 5737A-650A-(S1421) 40

. [g 3.2.'4.4. ' Bundle Enthalpy Rise

\_)

Since no calibration / verification of subchannel analysis codes against experimental data has yet been done, a direct f actor of 8% is estimated (3) for the flow distribution calculational uncertainty. The reactor AT variation and coolant properties uncertainties are obviously the same as for the fuel and blanket assemblies. Analyses were performed to assess the effect of wire wrap variations and it was found that the average side channel temperatures were insensitive to such variations, thus yielding a hot channel f actor of. unity. Finally, an analysis similar to the fuel assemblies, to assess the effect of subchannel flow area variations, showed on the average a positive uncertainty of ~ 1.16 and a negative uncertainty of 0.87.

3.2.4.5 Bypass Enthalpy Rise The subf actors affecting the bypass enthalpy rise uncertainty, which still need to be considered, are: reactor AT variation, subchannel flow area ar.d coolant properties. In f act, the effect on the bypass flow due to the O uncertainty on the bundle / bypass flow split has already been accounted for as discussed in Section 3.2.4.3. While the reactor AT variation and the coolant properties uncertainties are necessarily the same as discussed for the other camponents, ad hoc analysis was needed to assess the effect of subchannel flow area variations since the same rod bundle displacenent affects differently the temperature in the bundle side channels and in the bypass. In f act, positive and negative uncertainties of 1.23 and 0.89, respectively, were calculated for the bypass, which are different from the corresponding values for the rod bundle.

3.2.4.6 Film 1

See corresponding Section 3.2.2.4.

3.2.4.7 Duct

-See corresponding Section 3.2.2.5.

.. r._J 5737A-650A-(S1421) 4l

I 3.2.4.8 Interstitial Gap See corresponding Section 3.2.2.6.

l l

O O

573/A-650A-(S1421) 42

~

K (f 4. STEADY STATE THERMAL, PERFORMANCE 4.1 Plant Conditions Following the practice established during previous analyses II) , two sets of plant conditions are used in the themal-hydraulic design, i.e., plant thermal-hydraulic design value (THDV) conditions and plant expected operating 0

conditions (PE0C). The THDV conditions (730 F in13t/995 F outlet 0

temperature; total reactor flow 41.446 x 10 lb/hr) are the Clinch River rated plant r )nditions and are used in: a) analyzing pemanent components which have the same 30-year lifetime as the plant; b) transient and safety 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 uncertainties, fouling and plugging of heat exchangers, etc.

i Expected operating values for the primary heat transport system principal Q parameters (inlet, outlet temperature and AT) are thus evaluated, together with the associated uncertainties. The results of this study for the heterogeneous core, which comprised a Monte Carlo type analysis, are reported in Table XVI. The major differences in respect to an analogous study previously perfomed for the homogeneous core are: 1) the consideration of the progressive fouling of the heat exchangers during the plant 30-year lifetime, which affects the predicted values of the plant operating conditions

. (in the previous studies, end-of-life fouling, i.e., af ter thirty years operation, was conservatively assumed in evaluating plant expected operating conditions); and 2) a more comprehensive accounting of all uncertainties af fecting plant operation. Plant expected operating conditions are adopted in core thermofluids analyses of replaceable conponents, such as the core assemblies, chiefly. in determining the fuel rod parameters (cladding tenperature, fission gas pressure) which are the basis for evaluating the structural behavior and for assessing whether lifetime /burnup objectives are actually met.

A

(_).

5737A-650A-(S1421) 43

As mentioned in Section 1, a critical reevaluation of the flow impedance in the entire primary system led to a significant increase in the value of the expected reactor flow. A new consistent set of plant expected operating conditions to be used in thermal calculations has not yet been developed, thus, the decision of utilizing in the interim the more conservative conditions already determined, which will yield higher temperatures in the core. Updated values of expected reactor flow were used for pressure drop calculations, as reported in Sect _n 6.3.

Plant expected operating conditions and associated uncertainties adopted in the thermal performance analyses are reported in Table XVII. Following is a brief discussion of the rationale in detennining the values reported in Table XVII from the ones in Table XVI.

First, the mean values of Table XVI are chosen as the nominal values of Table XVII, thus, conservatively including the bias f actor directly into the nominal values. Since the most critical time for core assemblies is at the end-of-life, when the cladding strain and damage function are maximum, second core values have been selected as corresponding to four-year fouling conditions. Due to the f act that four-year fouling conditions were not evaluated, it was assumed that the same difference in plant parameters between year two and year zero repeats between year four and year two. Again, the selected approach is conservative for two reasons: 1) plant conditions have been considered constant over the two-year span and equal to the worst end-of-span conditions, thus reeglecting the more favorable conditions which exist throughout the cure lifetime; and 2) the effect of fouling is not linear with time, but it is rather pronounced at the beginning and then tapers off during the plant lifetime, as can be seen by comparing plant parameters in Table XVI for 0, 2 and 30 years. Thus, the assumption that the same deterioration of plant conditions which occurs in the first two years (first core) also occurs during the third and fourth year (second core) is conservative.

While the mean values of plant parameters are consistent (i.e., outlet temperature equals inlet temperature plus AT), the same is not true when uncertainties are included. In f act, uncertainties quoted in Table XVI are O

5737A-650A-(S1421) 44

q for each parameter independently; thus, if the inlet temperature and the AT at the 97.7 confidence level (e.g., for the two year fouling) are added, the outlet temperature is equal to 999 F, significantly higher than the 976 0F reported. Actually, 976 0F represents the 20 outlet temperature, while 999 F is, roughly, a 40 value.

Because the inlet temperature and AT are defined, while the outlet temperature is derived, the following procedure is used:

e The uncertainty on the AT is calculated as a dimensionless factor and is combined statistically with other engineering and nuclear uncertainties.

e The uncertainty on the inlet temperature is combined statistically with the loop-to-loop imbalance effect and the combined uncertainty is directly added to the nominal value.

This approach is conservative. The loop-to-loop imbalance effect is much smaller than the inlet temperatum uncertainty. If there were no other-uncertainties, the outlet temperature would be at approximately the 40 level, as previously mentioned. However, other uncertainties, engineering and nuclear, do affect the reactor AT, and therefore, when combined statistically with the plant conditions uncertainty on AT, will actually decrease its value.

Finally, with regard to the uncertainties on plant operating conditions reported in Table XVII, the following must be noted:

e Uncertainties on rppctor AT are calculated and applied as in previous analysesU/;

e Uncertainties on inlet temperature are treated differently than for the homogeneous core studies. Previously, an uncertainty on reactor inlet temperature, as evaluated through Monte Carlo type analyses, was translated into a reactor AT uncertainty and combined statistically with other subfactors affecting the reactor AT. In addition, a 160F uncertainty was superimposed directly on the nominal inlet temperature value to account for loop-to-loop imbalance and primary loop temperature control uncertainties. The improved procedure adopted in these analyses was, as previously (t

Q,/

5737A-650A-(S1421) 45 i l

l

mentioned, to combine statistically the loop-to-loop taperature imbalance effect with the uncertainty on inlet temperature (which is due to all effects related to plant operating conditions) and to add the cmbined uncertainty to the inlet temperature. The loop-to-loop imbalance effect was evaluated from experimental data obtained in the CRBR inlet plenum feature test conducted at HEDL where different values were obtained for each inlet module ranging fra practically zero near the reactor center to a maximum value of 4.60F at the core periphery. Because such variation was minimal, compared with the much greater plant uncertainty with which the imbalance effect is cabined, for simplicity the maximum value is conservatively used for all assemblies.

e In previous analyses (l), the power level measurement / control dead band uncertair.ty was considered separately in the hot spot / channel tables as a direct f actor. However, in the present analyses, it is already included in the Monte Carlo evaluation of plant operating conditions uncertainties.

Therefore, the power level measurement / control dead band uncertainties are not considered separately in themal-hydraulic analyses conducted for plant expected operating conditions, but are still considered separately in those analyses perfomed for plant thermal-hydraulic design conditions.

e Even though the plant cperating conditions uncertainties adopted in these studies represent a sensible improvement over the analogous analyses perfomed for the homogeneous core, they are still quite conservative, as previously mentioned.

In f act, a statistical analysis perfomed in Reference 3 showed that the value assigned here to the uncertainty f actor on AT is actually inclusive of the inlet temperature variation effect as well. The loop-to-loop imbalance uncertainty recommended in Reference 3 is 7.40F, higher than the 4.60F used here, but this is more than offset by the f act that a 2a inlet temperature uncertainty of 330F (first core) or 360F (second core) is adopted here (see Table XVII), while in Reference 3, the inlet temperature uncertainty is included in the AT f actor (1.14), as mentioned above.

4.2 Linear Power 0

Linear power ratings over a 60 core sector (fuel, inner and outer blanket assemblies) are reported in Figures 11 through 20.

Average, peak, 30 and 30 plus overpower linear ratings are reported, To clarify the adopted nmenclature, " average" represents an arithmetic average over the 217 (61) rods of the fuel (inner / outer blanket) assembly. Theref ore, O

5737A-650A-(S1421) 46

( it generally represents a fictitious rod not exactly corresponding to any physical rod in -the assembly. " Peak" refers to the rod in the assembly having the highest power; i.e., no uncertainty f actors are applied iii the evaluation of the peak power rating. "30"' power rating refers to the value resulting from applying to the peak rod both the uncertainties on the nuclear peaking factors (radial and axial) and the engineering uncertainty f actors, both at the 3a level of confidence. The "3a plus overpower" values are derived from the 3a linear power ratings by applying an additional 15% over the CRBR rated nominal full (975 PWt) power.

The maximum (at 3a plus 15% overpower conditions) fuel assembly total linear power rating occurs in assemblies 101 and 68 (15.8 kw/f t at beginning-of-cycle 1; 15.1 kw/ft at beginning-of-cycle 3, Figures 11 and 15, respectively). The maximum inner blanket assembly linear power rating occurs in assembly 99 at end-of-cycle 4 (20.6 kw/ft, see Figure 18). During the first core, the highest inner blanket power rating is in assembly 67 at end-of-cycle 2 (19.2 kw/ft, Figure 14). Finally, the maximum linear power rating in radial blanket

, assemblies occurs in assembly 201, and the symetrical 301, at end-of-cycle 3 (16.4 kw/ft, Figure 18), in spite of the f act that the assembly power is highest at end-of-cycle 4 (compare peak power ratings at EOC3,12.2 kw/ft, Figure 16 and at E0C4, 12.4 kw/ft, Figure 18). This is due to the time-dependence of the nuclear uncertainties (see Table VIII-B) which are maximurn at BOL and decrease to a minimum value at E0L.

r 4.3 Assemblies Mixed Mean Temperatures Assemblies mixed mean temperatures are presented in Figures 21 through 30 for beginning and end of each of the first four (five for second row radial blanket assemblies) cycles. Since calculated values of the mixed mean temperatures, specifically maximum temperatures and temperature g, adients between adjacent assemblies, are a critical input to the upper internals structure design, plant THDV conditions were adopted in these calculations.

In f act, the UIS is a pennanent, 30 year lifetime, component.

A F i L/

r 5737A-650A-(S1421) '47

Nominal (no uncertainty f actors applied), Oo (only direct uncertainty f actors are applied), 20 and 3a (direct plus statistical uncertainty f actors at the 2a and 3a level of confidence are applied) are reported in the figures.

First core conditions are the worst for the UlS from the point of view of both maximtsn temperatures and temperature gradients. The maximum temperature I*)

is 1123 F in assembly 45 at BOC1, with a maximum gradient (273 F) between assmbly 52 and 302 (see Figure 21). Mixed mean temperatures follcw the same lifetime pattern as the power generation, thus inner blanket assemblies which start very cold at beginning-of-life (450 F at B0C1) attain temperatures caparable with those of the fuel assemblies at E0C2. The same pattern repeats in tm second core, cycles 3 and 4. The radial blanket assemblies start at approximately the same temperature as the inner blanket at B0C1, but it is not until cycle 4 or 5, which is the end of their life, that their tmperatures are cmparable with those of the other assemblies. The maximum mixed mean temperature for the second core occurs at B0C3 in assembly 45 (1115 F). While first and second cores are quite similar in tenns of maximum mixed mean temperature, they show a markedly different behavior as f ar as maximtsn gradients are concerned. During the first core, the maximum tmperature difference between adjacent assemblies occurs at the fuel / radial blanket interf ace: in cycle 1 between assemblies 52 and 302 (273 F at BOC1, 225 F at E0C1), in cycle 2 between assemblies 24 and 202 (227 UF at BOC2, 163 F at E0C2). In the second core, the maximum gradient position moves at the fuel / inner blanket interf ace: betwe.n assemblies 37 and 99 at BOC3 (239U F), assemblies 2 and 128 at EOC3 (136U F), assemblies 4 and 62 at B0C4 (149 F). At the end of cycle 4, the mixed mean temperatures of fuel and inner blanket assmblies are quite close, so that the maximtsn gradient occurs between two radial blanket assemblies, i.e., assemblies 206 and 213 (96U F).

Mixed mean taperatures reported in this section are calculated assuming adiabatic boundaries at the assemblies interf ace. Thus, the beneficial effect of inter-assmbly heat transfer in flattening the high temperature gradients is not taken into account.

(*)All temperature values reported in this discussion are nominal.

O 5737A-650A-(S1421) 48

[^] Mixed mean temperatures accounting for inter-assembly heat transfer are calculated by TRITON. Core-wide TRITON calculations were performed for 80C1 and -E0C4, as reported in Section 4.5. A comparison of mixed mean temperatures under adiabatic conditions and more realistically accounting for inter-assembly heat transfer (Figures 66 and 67) is discussed in that section.

4.4 Rod Lifetime Cladding Temperature / Pressure Histories All fuel, inner blanket and outer blanket assemblies in a 60 core synnetry sector were followed during their lifetime (i.e., first and second core for fuel and inner blanket assemblies; row 1 outer blanket assemblies over cycles 1 through 4 and row 2 radial blanket assemblies over cycles 1 through 5). The maximum cladding temperature and fission gas pressure in the hot rod at the 2a level of confidence for plant expected operating conditions were predicted for each assembly. Lifetime profiles are reported here (see Figures 31 through 51) only for selected assemblies, due to obvious space

. limitations. The limiting, for both first and second core, fuel and inner blanket assembly in each orificing zone, are provided. The limiting assembly j is the one requiring, in a given zone and for a given core, the highest amount of flow (see Figures 9 and 10). Note that the driver assembly of each orificing zone as defined in Section 2.4, is the assembly requiring the highest amount of flow in either core. Thus, the limiting assemblies in zone 1 (#10) and zone 4 (#52) in the first core are the drivers of the respective zones, while the limiting assemblies in the second core for zones 2 (#37),

zone 3 (#2) zone 7 (#99) and zone 8 (#46) are the drivers of the respective zones. The same assembly (#49) in zone 5 is limiting for both the first and second core and is obviously the driver. The radial blanket assemblies remain in the reactor through the first and second core, thus, no distinction is necessary: the drivers of zones 9 (#201),10 (#203) and 11 (#206) are first row assemblies and have a 4 cycle lifetime; the drivre of zone 12 (#212), a second row assembly, has a 5 cycle lifetime. Regarding orificing zone 6, where fuel and inner blanket assemblies are alternating, typical examples of the various combinations are shown here: inner blanket assembly remaining in the same location in the first two cycles (#98, Figure 41), inner blanket assenbly in the odd cycle followed by a fuel assembly in the even cycle (#62, 7 cycles 1 and 2, Figure 42; #62, cycles 3 and 4, Figure 43). As a general (d

\

5737A-650A-(S1421)' 49

trend, fuel assemblies in the inner region of the core (see, e.g., Figure 36),

experience a cladding temperature jump from the odd to f ie even cycle, while the fuel assemblies in the outer region (see, e.g., Figure 37), experience a drop. The cladding temperature generally decreases during a given cycle. The cladding temperature, however, increases during the even cycle (see Figures 31, 33, 34) in those assemblies adjacent to the Row 7 corner cortrol assemblies. Blanket assemblies have obviously a continuously increasing temperature during their lifetime, a direct consequence of the increase in power. The lifetime behavior of a given assembly during first and second core is quite similar, both qualitatively and quantitatively. The only major difference is that a higher fission gas pressure is attained in the second core, due to the longer residence time and burnup.

As an overall summary, the highest cladding ID temperature attained in each assembly in the first and second core is reported in Figures 52 and 53, respectively, together with the time of occurence. As shown in the figures, the maximum cladding temperatures vary significantly from assembly to assembly, a direct consequence of the orificing philosophy, where assemblies were orificed to satisfy burnup/ lifetime goals and transient limitations, rather than equalizing cladding temperature. It can be noted for example that fuel assemblies in the inner core region, which are CDF limited, require lower temperatures than assemblies in the outer region, which are transient limited. Blanket assemblies, when starting from the same steady state temperature, attain a higher transient temperature than fuel assemblies.

Thus, the steady state maximum cladding temperature in blanket assemblies is lower than for the transient limited fuel assemblies, a direct reflection of the adoption of the same transient limit for all core assemblies.

Structural analyses performed utilizing the core assemblies thermal performance data reported here verified that indeed the burnup/ lifetime goals are satisfied in all assemblies during the first and second cores. Similarly, transient analyses were performed for some worst assemblies, starting from the steady state conditions reported here, and it was verified that the transient limitations are indeed met.

O 5737A-650A-(S1421) 50

i g 4.5 Duct Temperature and Related Analyses Detailed three-dimensional duct temperature distributions (axially, radially and circumferentially) were predicted in support of the core restraint design and related.(duct bowing, dilation, reactivity coefficient) analyses. A 60 0 core symetry sector was analyzed at plant THDV conditions at BOC 1 and E0C4, thus bracketing the entire lifetime considered.

The current version of the TRITON code which models a cluster of seven adjacent assemblies was used. The outer boundaries of the cluster are assumed to be adiabatic, while heat is transferred across the internal interf aces 0

(i.e., ducts and interstitial sodium flow gap). The CRBRP core 60 symmetry sector, plus one row of assemblies at each of the boundaries to provide the necessary boundary conditions to the assemblies within the sector, was analyzed in groups of seven assemblies at one time and changing each time to a different central assembly, which is " dumped" to output. The TRITON model, based on the subchannel analysis code COTEC, explicitly solves the thermal-hydraulics of wire wrapped assemblies (by considering turbulent mixing, sweeping, pumping and swirl flow). In addition, it accounts for the

("%

'N exchange of heat between adjacent assemblies. The code is able to model all types of core assemblies, including the radial shield, which were in fact analyzed to provide the proper boundary to the second row radial blanket assemblies.

Gama-heating in the ducts is considered in TRITON by including it in the total assembly power. For the control assemblies case, the gamma-heating is split between the absorber bundle and the bypass. Therefore, gamma-heating is accounted for in a global fashion when calculating duct temperatures, rather than as a localized effect.

Duct. temperatures were calculated under nominal conditions and accounting for uncertainties (at the 2a level of confidence). The uncertainties runs were as follows:

(.

Ev r

5737A-650A-(S1421) 51

e 80C1 - Fuel assemblies have positive uncertainties, blanket and control assemblies have negative uncertainties, uncertainties on power and flow in the radial shield were such to minimize their temperature; e E0C4 - All fuel, blanket and control assemblies have positive uncertainties, radial shield power and flow uncertainties are such to maximize their temperature.

The rationale behind the choice of tF" : combinations of uncertainties was that at BOC1 fuel assemblies have their maximum power, thus temperature, while blanket assemblies are at the coldest time in their life. Therefore, choosing the uncertainties in order to increase the fuel Lssemblies temperature and at the same time to decrease even further the blanket and control assemblies temperature, i.e., "make the hot hotter and the cold colder", will yield the worst possible cross-duct gradients in the CRBRP first two cores. At end-of-cycle 4, blanket assemblies are at their maximum temperature, thus adoption of positive uncertainties in all core assemblies will yield the highest level of duct temperature across the core. It is apparent that the selected combination of uncertainties and time in life is the one giving the highest temperatures when it is considered that selection of positive uncertainties will tend to increase the temperature in the affected and neighboring assemblies, while EOC4 is the only time when all the blanket assemblies (except the radial blanket second row) are at end-of-life (maximum power) conditions.

Uncertainty values used in these analyses were reported in Section 3.2.

Geometrical variations were selected to minimize inter-assembly heat transfer  !

(thus maximize thermal gradients) at 80C1, hence the maximum values of assembly pitch and duct thickness were used for all assemblies. At EOC4, ]

wnere the purpose is to maximize temperatures, the duct thickness and assembly l pitch was maximum for the fuel assemblies, but was minimum for the other I assemblies and at the interfaces of different types of assemblies. In fact, increasing the heat transfer between fuel and blanket assemblies has a noticeable impact on increasing the blanket temperature, but it is not very significant in decreasing the fuel assembly temperatures l l

52 9

5737A-650A-(S1421)

It should also be noted that even though no formal uncertainties were

[\ discussed in Section 3.2 for the radial shield assemblies, their power and flow were adjusted in the TRITON runs, in order to provide a " cold" and " hot" boundary to the radial blanket at BOC1 and E0C4, respectively.

Typical results of duct temperatures calculated by TRITON are reported in Figures 54 through 65. Midwall duct temperatures reported for each face are the average over the face of the detailed temperature profile calculated by l

TRITON (TRITON calculates local duct temperatures circumferentially along the f ace at each peripheral .subchannel). In addition, the direction (with arrow) and magnitude of the maximum cross-duct (midwall) temperature gradient in each -

assembly is shown. Generally, the largest gradients occur at the fuel / radial blanket interface, which is therefore a very critical region for the core restraint design. The typical mappings shown here are for three axial elevations; i.e., 32" (middle of the core), 60" (approximately the above-core-load-pad location) and 112" (top of the rod bundle). TRITON calculates core-wide duct temperatures at 0.5" intervals and all these data were transmitted, via computer tapes, to the core restraint designers.

U In examining the results reported in Figures 54 through 65 it is evident that the worst gradients, by far, are at BOC1 when positive uncertainties in the fuel and negative uncertaint;es in the blanket are considered. While this result was anticipated, some interesting observations can be made. Blanket assemblies temperatures accounting for uncertainties are generally lower than the corresponding nominal values, which means that the effect of negative uncertainties is greater than that of inter-assembly heat transfer, even though the latter was enhanced by considering opposite uncertainties.

Cross-duct gradients in individual assemblies increase significantly, in some instances they almost double. Even more importantly from the core restraint standpoint, is the increase in the temperature difference between adjacent faces of fuel and blanket assemblies. For example, at the top of the core (see Figure.s 56 and 57) the adjacent faces temperature difference is of the order of 50-80GF under nominal conditions, but of the order of 100-140 0F if the effect of uncertainties is considered. At E0C4, the duct temperatures increase quite uniformly throughout the core if the uncertainties are p

5737A-650A-(S1421) 53 i

j

considered, with increases in excess of 100 F. The crnss-duct gradients and temperature differences between adjacent f aces also increase, but not nearly as much as for the B0Cl case, the increase being of the order of 20 F.

It is obvious that many cmbinations of uncertainties exist; the two cabinations adopted in these analyses are to be considered only as representatives of typical extreme conditions. Calculations of duct temperatures provide the input necessary to structural analyses. Only after these are completed and the designers examine both the themal and structural behavior, will it be possible to define more precisely which distribution of uncertainties throughout the core should be used in the design. Such distribution must be tailored to the particular analysis being conducted, since a certain cambination of uncertainties could be the critical one for the core restraint design, a different one for duct dilation analyses, and so on.

Calculations of assemblies mixed mean temperatures have also been perfomed by TRITON; the results are reported in Figures 66 and 67 for BOC1 and E0C4, respectively, which cmpare TRITON calculated temperatures with the corresponding temperatures obtained for adiabatic conditions in Section 4.3 (and reported in Figures 21 and 28), thus showing the effect of inter-assembly heat transfer. As expected, at B0C1 where the fuel assemblies have the maximum power in life and the blanket assemblies the minimum, heat is transferred frm the fuel to the blanket assemblies and mixed mean ta peratures are higher in the blanket and lower in the fuel than otherwise calculated under adiabatic conditions. The maximum temperature difference between adjacent assemblies, which occured between assemblies 52 and 302 was 273U F under adiabatic conditions and is reduced to 261 F when considering inter-assembly heat transf er, a reduction of 4%. Reductions in adjacent assemblies gradients are greatest at the core center (where the power production in the f uel is maximum); for example, between assemblies 34 and 59, the gradient of 227 F calculated under adiabatic conditions, is reduced by 10% to 205 F when accounting for inter-assembly heat transfer. It should be also pointed out that the coolant does not exit fra the assembly with a uniform temperature equal to the mixed mean, 5 a radial temperature gradients exists within the assembly exit. f0N divides the assembly exit O

5737A-650A-(S1421) 54

l

. /7 area into 6 s'ectors and calculates the local temperature in each of these 5

sectors; thus, this detailed information-is available to the structural designers for refined analyses..

Finally, as input to duct dilation and bundle / duct interaction, selected assemblies (see Figure 68) were followed through their entire lifetime with TRITON. Parameters calculated and transmitted to the assembly structural designers were duct midwall temperatures over each assembly face and rod cladding midwall temperatures averaged over the rod circumference for two transverses; i.e., from corner to corner and flat to flat of the assembly (see Figure 69). 11 temperatures were calculated as a function of axial position.

The capability to calculate rod cladding temperature was therefore added to TRITON to perform these calculations. Obviously, calculating both duct and cladding temperatures in the same TRITON runs was much more efficient than calculating duct temperatures with TRITON and rod cladding temperatures via COTEC and NICER. More importantly, the COTECfNICER route considers adiabatic assembly boundaries, thus duct temperatures would have been calculated

/~} accounting for inter-assembly heat transfer, while calculated cladding temperatures would have neglected this effect, which is significant in the blanket assemblies. Thus, the temperatures calculated here and provided to the mechanical designers are correctly self-consistent. Typical examples of

, these calculations are reported in Figures 70 and 71.

It should be pointed out that TRITON now has the capability of calculating not only average cladding temperatures as in this case, but also local cladding temperatures accounting for hot channel / hot spot factors and in the same detail as presently calculated by COTEC and NICER. As mentioned in Section 1, this significant upgrading of analytical methods capability now opens the possibility of predicting fuel and blanket rods thermal performance rigorously accounting for the effect of inter-assembly heat transfer.

P t'

'-l L v) '

l 5737A-650A-(S1421) 55 L

i n, , - - - - ,em ,y g- -r-we---- . ---,,..- -- . - , --- <.,-r-- -ee -r

^'T

/ 5. POWER-TO-MELT ANALYSES LI 5.1' Introduction Analyses for the worst fuel and inner blanket assemblies were perfomed to investigate whether the criterion of no incipient melting at 115% rated power and accounting.for uncertainties at the 3a level of confidence is actually met. The fuel assablies analyses focused on the peak power pins at beginning-of-life, which are the most critical from the point of view of power-to-melt, and showed that adoption of a progranmed startup will guarantee satisf action of the no-melting criterion. Analyses of the inner blanket assablies (which envelope the radial blanket assemblies, since the latter attain much lower powers at E0L) investigated both the hot (maximum cladding temperature) pin and the peak (maximum power) pin. In fact, the cladding temperature has a very marked effect on cladding swelling and ultimately, through gap size, gap conductance and fuel centerline temperatures, on the power-to-melt. Analyses indicated that the no fuel melting criterion in the inner blanket is satisfied, in spite of the adopted conservatism, which is I

o i believed to be excessive.

V 5.2 Fuel Assemblies Power-to-Melt Analyses A detailed power-to-melt of the peak power cins in fuel assemblies 101 and 14 were perfomed, using the LIFE-3 code (20) Gemetry and operating conditions of these pins (calculated by NICER for plant THDV conditions and Oc hot spot factors) are reported in Table XVIII. Assembly 101 was selected as the one with the maximum power rating at 3a overpower conditions, while assembly #14 has the highest peak power pin (see Figure 11). LIFE-3 input was obtained from appropriate NICER runs. The startup ramp used in this study is shown in Figure 72. No effort was devoted in this study to the optimization of the startup procedure, which will be performed for final, as-built design.

An experimental program to investigate the startup procedure and to guide in its optimization has been proposed. Since BOL is the time when the power is maximum and the power decreases with life due to the depletion effect, power-to-melt analyses for.the fuel assembies were not conducted beyond the p progranmed startup.

' N,1 5737A-650A-(51421) .56

The cladding swelling used in this study was Revision 5 of the first core steel swellingIIO) Note, however, that since early-in-life cladding swelling is negligible, the choice of cladding swelling equations is actually unimportant.

The nomina' powers-to-melt computed using LIFE-3 are tabulated in Table XIX.

The axial location (X/L) values of Table XVIII were chosen since they are the positions of maximum power. Due to the axially increasing temperature in the cladding, melting will first occur slightly above the position of maximum power, but this effect is neglected here, since it can be easily accommodated in the optimization of the programed startup. Note that the programed startup can not affect the power-to-melt in blanket, which occurs at end-of-life, thus, the actual temperature gradient was taken into account for the analyses reported in Section 5.3.

To detemine whether the pin designs and power histories are able to withstand a 15% overpower transient without melting within a 3o level of confidence, relative statistical power-to-melt uncertainties must be applied. The same uncertainty values and procedure reported in Reference 3 were adopted in this analysis and are briefly sumarized in the following and in Section 5.2.1 The primary data used for this work were the results of the short time (low burnup) HEDL P-19(21) and P-20(22) tests which were designed to provide themal perfomance infomation.

The LIFE-3 code used in this study is a detailed model for describing the themal and mechanical behavior of f ast reactor oxide fuel pins and has been calibrated and verified with the HEDL P-19 and P-20 tests in addition to a number of intemediate and high burnup pins. Figure 73 and Table XX show how well the code represents the P-19 and P-20 data. The code was used to calculate the power-to-melt of the CRBRP fuel and blanket rods, and the sensitivities to variations in fuel pin parameters used for uncertainty analysis.

O 5737A-650A-(S1421) 57

[* For fuel applications, the uncertainties in the code predictions of the

] power-to-melt arising from data scatter and the over all accuracy of measurements in the EBR-II reactor are first analyzed. The uncertainties which occur when applying the code to the CRBRP conditions are then evaluated. The f actors considered here include the tolerances on f abrication parameters and reactor instrumentation, nuclear, themal-hydraulic and systenatic uncertainties. The individual uncertainties were determined by calculating the effect of a variation in each parameter on the _ nominal power-to-melt.

All the individual uncertainties are then statistically combined and the probability distribution for the reactor power-to-melt determined. The design criterion can then be measured by detennining if the 115% of nominal power is at least three standard deviations below the power-to-melt.

5.2.1 EBR-II Uncertainties The f actors associated with EBR-II experiments which contribute to uncertainties in power-to-melt measurements are listed and defined below (where o is the standard deviation):

1. otime is due to uncertainty on overall power level due to variations in EBR-II instrumentation and the uncertainty in the neutronics calculation for a given core loading. This uncertainty causes random fluctuations in quoted power level that vary with time.

2. sys is due to a difference between actual and quoted overall EBR-II power that doesn't change with time. It is known that a systematic shift in EBR-II power level exists and a correction is made by experimenters.

An estimate is required of the uncertainty on this correction for this analysis which is identified here as sys. This uncertainty does not show up as scatter in the data. Such a systenatic uncertainty would show up in the scatter of data comparing different reactors.

!. /^\

V ,

l 58 5737A-650A-(S1421)

+

3. space is 60e to uncertainties in the spatial dependence of neutronics calculations and local inhomogentities in the EBR-II core.

4. o fab is due to variation in fuel p!n fabrication parameters from their nominal values.

5. is due to uncertainty in post-irradiation examination measurements.

PIE The evaluation of each of these uncertainties will now be discussed.

U time The major contribution to this uncertainty results from fluctuations in the primary and secondary EBR-Il coolant loop heat balance. This yields a 2% standard deviation (23) . Neutronic calculations of y precursors in the P-19 test introduce a further uncertainty of less than 1%(24) and a value of 0.8% was used. Accordingly, this factor was evaluated as at s,, f .4%2 o,gx2 = 2.2%

This uncertainty would not show up as fluctuations in results from a single subassembly since all pins would have been subjected to the same errors in overall power determination during the same time.

"sys A detailed analysis of the P-19 test indicated that a correction factor of 0.94 must be applied to the calculated EBR-Il power level (24) This factor has been used in all reported analyses of these results and those of the P-20 test and was used for the LIFE-3 calibration. Subsequent work suggested a factor of 0.91(23) , while a recent analysis by the EBR-II Project indicates a value of 0.96(25) As an interim position, the 0.94 factor is being retained. An evaluation of the recent analysis 5737A-650A-(51421)

O 59

(3 by the EBR-II Project will be perfonned when an official documentation is V issued. In addition, burnup analysis data will be evaluated with respect to this f actor. The EBR-II Project also . estimated that further systematic bias should be not more than +2%(25) _

. An uncertainty, c

3yy, of X is used here which covers both the .91 and .96 power f actors and is more conservative than the 2% estimate by EBR-II.

" space The spatial uncertainties in pin powers arise from uncertainties in neutron transport calculations. Calculations have been checked by conparison of measurments of control rod worth (24) . Table XXI and Figure 74 show the resulting percentage difference between measured and calculated flux, which has a standard deviation, Space,of1.7%. In addition, local flux peaks produce an estimated uncertainty, ospace' UI about 0.5%. Finally in the highly enriched fuel of P-19, differences in the y absorption can produce a +0.7% to -0.7% variation in pin power 1 going frm the center to the outside of the assembly. This introduces a (G standard deviation of 0.7/,/I f 0.4%. Combining:

" space 1.72 + 0.52 + 0.42 = 1.8%

, " fab The scatter in burnup measurements as compared to values calculated from EBR-II powers has been analyzed and found to be BU T 1.1%. This analysis is from individual subassemblies, and since pins in a subassembly are in the reactor for the same time, they would not reflect the uncertainties time and a sys.

The scatter in the burnup data includes a combination of EBR-II power uncertainties and uncertainty in the burnup measurement technique. Thus:

+ (0.5%)2 "BU T .1%

1 i[o' space + " fab ($)

5737A-650A-(S1421) 60

,,e . ,, - - - - . --e -

.ne - , - - - - - < s

where the 0.5% is the estimated burnup measurement accuracy of the mass spectronetry. The spatial uncertainty o' space 4.5% is that for an individual subassenbly and not the entire core.

The burnup uncertainty has also been independently calculated by HEDL to be 3% (la). In this work, a single power f actor for all suassemblies was used. The data scatter due to power fluctuation with time, time' and core-wide spatial uncertainty' s ace, w uld therefore apply and the 3% standard deviation obtained can be attributed to 3% i[ time + " space * " fab +(0.5M (ii)

Using the previously detemined values for o time' U space and a space in equations (i) and (ii) two values for o are obtained, fab 0.85% and 0.82%, which are therefore in excellent agreenent. For simplicity, a rounded-off value of 0.8 was used in these studies.

PIE O

The scatter of the P-19 and P-20 experiments relative to the LIFE calibration has been calculated (see Table XX and Figure 73). The standard deviation is about 1.3%. It is interesting to note that only one pin in this group, P-19-30, was significantly outside this standard deviation. If P-19-30 had been excluded, the standard deviation would have been 0.8%.

It might be more reasonable to use 0.8%, but to be conservative,1.3%

will be used. The scatter in P-19 and P-20 data reflect uncertainties in fabrication, local spatial fluctuations and post-irradiation examination measurement uncertainty. Since P-20 powers were nonnalized to P-19 results, P-20 is not an independent experiment at a separate time and o time does not apply. So O

5737A-650A-(S1421) 61

ym, .

1. I*

• U fab *

"P-19/P-20 + " space PIE and PIC

=OM Values of the above uncertainties are sunmarized in Table XXII.

Total Uncertainty The total uncertainty is a combination of all the components

+ +

tot

• i fab PIE + " space time + "sys = 4.3%

Resolution of the EBR-II power f actor is expected to reduce the systenatic uncertainty to 2%, and a will & op to 3.7%.

tot O atot represents the scatter and uncertainty in the power-to-melt data. The uncertainty in the average of these data is given by the standard deviation of the mean. The standard deviation of the mean takes into account the number of data points N, and is given by o/MT. In the various components of atot, a separate measurement toward detennining o would require use of a gy3 different. reactor; a separate measurement for o would require an time experiment done at a different time and a separate measurenent for Space would require another experiment done in a different position in EBR-II. As explained above, P-19 and P-20 cannot be counted as being done at different times or different positions. So for asys'"Npace and o the it is N = 1. For the other canponents N is equal.to 10, the number of rods used. Thus,

+ + +

"mean " sys + " time 5 pace *I fab PIE space)/10 = 4.1%

b o)'

5737A-650A-( S1421) . 62 s.wm-- m.- - ,

is dominated by o sys time with the other components making a o and mean small contribution. The mean of the power-to-malt dita is represented by the LIFE-3 calibration and the uncertainty on thic iiisan is given by mean*

5.2.2 Results To evaluate the probability of melting in CRBRP, the L2FE-3 calibration will be extrapolated from P-19/P-20 to CRBRP conditions. Since only f resh and very low burnup fuel is considered and since this is the burnup range covered by P-19/P-20, the extrapolation is small and it is assumed that no biases are introduced. An additional uncertainty comes frcm the use of fuel with 33% Pu enrichment instead of 25% enrichment used in the calibration. An estimate of the values and uncertainties of power-to-melt in CRBR with reference fuel can be made. Then the uncertainties in the CRBR power-to-melt are statistically added to the P-19/P-20 omean uncertainty and to the extrapolation uncertainty to define a probability of melting.

The uncertainties in a CRBRP power-to-melt analysis due to the f actors listed in Table XXIII are considered. Fabrication and irradiation uncertainties arise from design tolerances and uncertainties in neutron physics and themal-hydraulic calculations. These uncertainties can be estimated from the design tolerances by computing their effect on power-to-melt using LIFE-3.

The design tolerances will be for pellet density, cladding ID, enrichment, instrumentation and pellet diameter. There is also a tolerance on fuel weight per length which prevents certain combinations of density and diameter. To simplify analysis and add the conservatism of not taking credit for the weight per length restriction, this effect will be neglected. Tolerances lead to fabrication of parts whose means may lie with approximately uniform probability anywhere within the f abrication tolerance bounds. The bounds of such a rectangular distribution correspond to + /35. The tolerances are listed in Table XXIII. Actual distributions of dimensions and fuel density

are expected to be available for use in the FSAR hot channel factor analyses.

O 5737A-650A-(S1421) 63

' N LIFE-3 runs were made to a..aiyze the effect of variations in the parameters of k

Table XXIII on the melting of the pins investigated. A programed reactor startup is specified by giving the steady state reactor power as a function of time, REPOW (t). This is the total reactor power as determined by the reactor control settings, while Q(t) is the corresponding linear power, including direct f actors, of the peak pin near its axial midplane at the location of peak centerline fuel temperature. The reactor power units are nomalized to a value of 1.0 at nominal full power, i.e., REP 0W(t) is 1.0 when the nominal power of the highest peak power pin (assembly #14) at X/L = 0.45, Q(t), is 12.73 kw/ft, which is the nominal power used for the uncertainty analyses.

The power-to-melt is defined in LIFE-3 by ramping the reactor power up until melting starts. This is done at vaaious times during the program startup.

The programed startup assumed for this study is illustrated in Figure 72.

Using the nominal conditions of Table XXIII and the power history of Figure 72, the mean reactor power-to-melt for the nominal peak pin is defined as:

RTF0V g (nominal, t) = CP M I, nominal, t) g whe e D g is the peak pin linear power at X/L = .45 when the LIFE-3 centerline temierature reaches the fuel melting point and REFoV g is the reactor power at that time. CP g= 12.73 kw/f t because of the normalization of units chosen for reactor power. REPW g is defined as the mean reactor power-to-melt (signified by the bar) since the mean of P-19/P-20 has been used to calibrate LIFE-3 and the P-19/P-20 power uncertainty will be subsequently added.

The effect of small perturbations of the design parameters was analyzed to determine the sensitivity of power-to-melt to variations in each parameter.

The sensitivities are shown in Table XXIV and the corresponding standard deviations in power-to-melt obtained from the standard deviations in the parameters from Table XXIII are shown in Table XXV.

i v!

I 5737A-650A-(S1421) 64

.m

It is assumed that reactor power-to-melt can be approximated by a linear function of each parameter about the nominal point. This assumption was checked for pellet diameter variations for which the greatest non-linearities would be expected. Figure 75 shows the actual variations of power-to-melt with pellet diameter and the assumed linearizations which were selected to give a close fit over the range where power-to-melt was below nominal.

The melting temperatura of (0,25%Pu)0 2 is 2760 C (5000 F) witi a 3a uncertainty of 160 C (108 F). This value was based on an ext ensive review of all published data and a thorough analysis of the techniques and data can be found in the Nuclear Systems Material Handbook.(IO) It is not necessary to use an uncertainty on melting temperature when the LIFE-3 code is employed to analyze (U,25%Pu)0 2 since the code was calibrated with fuel with this plutonium content. The nominal melting point value was used for calibration.

Since the current CRBRP fuel contains 33% Pu0 , additional uncertainties due 2

to increased Pu content must be considered. Based on the calibrated melting point of 2760 C for 25% Pu0 2 fuel, and considering the +3a uncertainty of 160 C for the entire range of (U,Pu)0 2 solid solutions, the uncertainty in extrapolating from 25% to 33% Pu0 2 was detemined by estimating the U

uncertainty of the slopes. The 3a uncertainty in extrapolation is 1.2 C per wt% Pu0 2 or 9.6U C for the 33% Pu02 f uel (i.e., melting point uncertainty U

oTM = 5.8 F). The sensitivity of power-to-melt changes in melting point O  !

REPOW MP M

is calculated from the nminal LIFE-3 run, and this is used to convert the uncertainty on melting temperature to uncertainty on power-to-melt, i.e.,

"extrap " M !M MP U TM l

l 4

O\

5737A-650A-(S1421) 65 1

1

p To obtain the probability distribution for the reactor power-to-melt, the uncertainty in the P-19/P-20 experiments and the extrapolation uncertainties are added to the CRBRP design uncertainties, resulting in the probability distribution for reactor power-to-melt with standard deviation REPOW (t) = [og (t)2 + Umean +U extrap and a mean of RTF0V g (nominal,t).

The design criterion is that,- throughout life, the fuel linear power rating is at least three standard deviations below melting power, when the reactor power is 15% above nominal conditions, i.e.:

'K

• REPOW(t) < REPOWg (nominal,t) - 3 REPOW (t) g

.p Note that the uncertainty (1.03) on power level (see Table V-A) is already accounted for in the Oc operating conditions (Table XVIII) of the pins, which represented the starting point for this analysis. Thus, it must be extracted from the 15% overoower f actor, to avoid being considered twice.

Consequently, the value of K in the present study is 1.15/1.03; obviously, if nominal conditions were used as the operating conditions for power-to-melt analyses, K would have been equal to 1.15.

The results of the analyses outlined above are sunnarized in Tables XXIV through XXVII. In Table XXIV, the derivatives of the power-to-melt with respect to the different parameters are calculated. In Table XXV, these are used to convert design uncertainties into uncertainties on power-to-melt which

'are then statisti ally combined.

The uncertainty on neiting point due to extrapolation to a Pu enrichment different from P-19/P-23 is given in Table XXVI. The nominal LIFE-3 run is used to find the effect cf a change in melting point on power-to-melt. This is used to convert the uncertainty on melting point into an uncertainty on

-power-to-melt.

v 5737A-650A-(S1421) 66

In Table XXVII, the uncertainties from the measured power-to-melt (converted to reactor power units), the extrapolation to CRBRP conditions, and the CRBRP design uncertainties are statistically combined, and the design criterion for melting is examined to see if it is fulfilled for the programed reactor startup chosen.

Table XXVIII gives a comparison of the 30 minimum powers-to-melt with the power at 15% overpower for the two pins examined in this study. The 3a uncertainties in power-to-melt (3 REPOW ) from Table XXVII are tabulated in column 5 of Table XXVIII (expressed as 5 fraction of nominal full power). It can be seen in the last column of Table XXVIII that there exists an adequate margin against melting for both pins when the present calibration of LIFE-3 is used; i.e., the no-melting criterion is satisfied at the 30 level of confidence with additional margin in power remaining.

The results clearly demonstrate that a programmed startup for the fuel rods can be identified which ensures that the no-melting criterion can be satisfied. The programmed startup used was noi an optimum one and further work will be perfonned in this area. As previously mentioned, an experimental program has been identified to characterize the startup procedure, nanely duration of initial period during which power is held below the rated level and magnitude of the hold power as percentage of the rated full power, as well as to determine the enrresponding power-to-melt. This program will help in defining the optimum startup procedure to be recommended for CRBRP.

5.3 Inter Blanket Assemblies Power-to-Melt Analyses Inner blanket assembly #99 was investigated in these analyses, 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-cycle 4 (see Figure 18). Inner blanket assemblies are enveloping as regards power-to-mnlt conditions, since their much higher power rating overshadows, by f ar, the additional decrease in fuel melting temperature occurring in the longer residence time radial blanket assemblies. As in the fuel assemblies case, thermal parameters (circumferentially averaged cladding and surf ace 5737A-650A-(S1421) 67 O

'g - temperatures, rod ' linear power and burnup) predicted by NICER at THDV O( . conditions and Oo level of confidence were used as input to LIFE-3 analyses.

- Additionally, thennal parameters at the 30 level were calculated by NICER for use, together with the Oo values, in the uncertainties calculations. Cl adding properties were those for first core steel with Rev. 5.swe11ing(18) and cladding swelling uncertainties (T = 5.0, R

• 1.30) were fram the same source. Table XXIX sumarizes the LIFE-3 runs.

For the first seven LIFE-3 runs, the neutron flux uncertainty was assumed to be the same as the power uncertainty; for the final run in which the axial variation of power-to-melt was investigated at two-inch intervals, the neutron flux uncertainty was conservatively estimated as 1.15 throughout life, and the power uncertainty was fitted to an inverted chopped cosine distribution between the . values given at core midplane and fuel / blanket interf aces (see discussion in Section 3.1.4). The power-to-melt uncertainties other than those for cladding swelling and surf ace temperature are not expected to be significantly different from those given in Reference 3, thus the Reference 3 values were used.

U The power-to-melt in reactor power units (nominal full power = 1.0) is shown

in Table XXIX for each run.

The margin-to-melt analyses were performed as in Reference 3. Two cases were considered:

1

1. Reactor operating at Oo power including the 1.03 hot channel f actor for power level measurement and control system dead band, with an overpower margin of 1.15/1.03 required at E0L.

2. Reactor operating at Oo power not including the 1.03 hot channel f actor, with an overpower margin of 1.15 required at E0L.

.v 5737A-650A-(S1421) 68

Re-evaluated values for the uncertainty analysis are shown in Table XXX, with the remaining uncertainties the same as those in Tables 4.4 through 4.7 of Reference 3. From Table XXX, it is seen that case (2) has 0.004 less margin-to-melt than case (1), and that the peak pin has 0.004 less margin-to-melt than the hot pin.

From Run 8, the elevation with the minimum margin-to-melt was found to be 2 inches above the core midplane. Applying the margin-to-melt analysis at this elevation to the peak pin, case (2) resulted in zero margin-to-melt. Thus, the design goal of no incipient center melting at 30115% power conditions in satisfied, although with no margin.

It should be noted, however, that in reality additional margin exists, because of the conservative manner in which power uncertainty f actors have been applied. Each blanket power uncertainty (see Section 3) is expressed as a i uncertainty, plus (in some cases) a bias. For the analyses to-date, all of these f actors have been combined directly with the other uncertainties in the blanket assemblies, which is a conservative approach. However, the overall f power uncertainty could be combined statistically with the remaining uncertainties, since they are independent. Adopting this approach would substantially improve the power-to-melt margin.

Future analyses could also investigate other pins, located between the peak and hot pin. In f act, their behavior in tenns of margin-to-melt is so similar that it cannot be excluded "a priori" that an intermediate pin could become limiting, due to a worst combination of linear power rating and cladding temperature / swelling.

l l

O 5737A-650A-(S1421) 69

6. CORE ASSEMBLIES PRESSURE DROPS p

-t i b' 6.1 Introduction The area of pressure drop calculations has gone through many improvements and reevaluations throughout the CRBRP core design. Following is a brief discussion of the reasons for this continuous reassessment:

e More experimental data from prototypic hydraulic experiments of various reactor and assembly components have become available, thus replacing analytical / engineering estimates with experimentally derived correlations.

e Uncertainties on predicted pressure drops of several components have changed from engineering estimates (a typical example being a 20% uncertainty on form losses) adopted in previous analyses to actual calculations based on regression analysis of experimental data, e Removal of unnecessary, parasitic pressure drops in the assemblies orificing has led to an increase in available actual reactor flow, by moving to the right along the pump head / flow characteristics curve.

m ,' e Several modifications have been factored into the lower inlet modules (LIM) design. The requirement of a constant pressure drop across the various LIMs, and consequent orificing, has been removed. Thus, assemblies belonging to the same orificing zone but to different LIMs will have slightly different flows, depending on the total pressure drop across the respective LIMs. As mentioned in Section 1, the capability for radial blanket assemblies shuffling has been retained in the heterogeneous design. All the radial blanket assemblies are therefore identical and the radial blanket orificing is apportioned between the assembly and the LIM.

Thus, in zone 9 (highest flow orificing zone in the radial blanket) all'the orificing is in the assembly, while in zones 10, 11 and 12, a progressively increasing pressure drop is incorporated in the LIM orificing to balance the decrease in assembly pressure drop going to lower flow orificing zones.

The amount of orificing in the LIMs is thus designed to permit shuffling of zone 9 assemblies in zones 10,11 and 12, zone 10 assemblie; in zones 11 and 12 and zone 11 assemblies in zone 12.

.I

/^N\

5698A-643A-(S1421) 70

e A new computer code, CATFISH, whose features have been briefly surmiarized in Section 1, has been developed to conventiently handle all the modifications and improvements discussed above. Specifically, CATFISH is necessary to calculate the reactor flow conditions corresponding to the given pump head and reactor hydraulic impedance and to evaluate the coolant flow in each individual core assembly following removal of the pressure drop equalization in the LIMs.

e Finally, development of CATFISH and correlating the reactor impedance with the pump characteristic curve has very clearly pointed out a rather ovbious fact which was however overlooked in previous analyses. Reactor pressure drops had been previously calculated for a given reactor flow under nominal conditions and accounting for uncertainties (either positive-maximum AP or negative-minimum AP); thus, more than one pressure drop was calculated as corresponding t< .he same fl ow. It is clear that this is physically impo ible, since if, e.g., the reactor impedance is greater than predicted, the reactor flow has to be less than the nominal flow corresponding to the predicted hydraulic resistances, in order not to violate the pump characteristics. Therefore, in this study only mutual pairs of reactor flow /AP are reported.

Fcr example, Section 6.3 reports calculations performed by CATFISH for two such pairs: a) when the reactor flow is at its THDV value and the hydraulic resistance is maximum (which are the worst possible conditions for the thermal performance of the core assemblies) and b) when the reactor hydraulic resistance is nominal with a corresponding reactor flow equal to the maximum flow compatable with the pump operating range and flow measurement / controller uncertainties.

The following sections discuss the correlations used in calculating the pressure dr9s across the assemblies, their bases and some typical results.

6.2 Pressure Drop Correlations Preliminary hydraulic characteristics of the CRBRP fuel and blanket assemblies are reported here. Since several test programs are still in progress, the best available information has been included in this analysis which will be reviewed as new test data are obtained.

Correlations reported here are valid over the entire range of operation, either turbulent, transition or laminac. However, in this study only full O

5698A-643A-(S1421) 71

flow conditions are of interest, therefore, the flow in all components is in f_\ .

-- () the turbulent regime. This distinction is important when considering the uncertainties associated with the proposed correlations, as discussed later.

The rod bundle frictional pressure drop is defined by:

AP = (fL/D) pV2 /2 where -

f is the friction factor; L is the rod length; D is the rod bundle hydraulic diameter calculated from the rod bundle total flow area, A, and wetted perimeter, P,:

D = 4A/P,;

V is the rod bundle average velocity; and p is the fluid density.

The component hydraulic characteristic, K, is defined for all other components 1

() by the form losses equation:

K = 2AP/pV r

Wherever possible, the irreversible pressure loss, AP, was determined from test data

/2 - pV /2 AP = Ps1 - Ps2 + PY1 2 1

where is the differential pressure measured between upstream Psl - P s2 (1) and downstream (2) static pressure taps; Vy and V2are the average velocities over cross-sections 1 and 2; p is the test fluid density.

5698A-643A-(S1421) 72

Thus K has the form K = 2 (Psl - Ps2)/PYr +(A 2/A 1 2 -Ar /A2) where V is the ref erence velocity, i .e., the average velocity r

through some convenient reference cross-section Ap ;

A and A are the areas at cross-sections 1 and 2; t 7 Ar is the reference cross-section area.

The first component of the form loss equation represents the measured static pressure difference between points 1 and 2 and the second component represents the fluid acceleration between cross-sectional areas 1 and 2 and has a constant value. The form loss coefficient, K, represents the non-recoverable pressure drop between points 1 and 2 and can generally be correlated with the Reynolds number by a f unction of the f orm CRe-". The hydraulic characteristics thus defined combine both the form and friction losses of the component into one expression.

The data included in the characterizations developed herein are from tests of FFTF assemblies and components (26-28) , and tests of CRBRP components (29-31) .

Linear regression analyses were performed on the fuel assembly component pressure drop data. The coefficients of the regression functions were calculated for the fuel assembly components and are presented in Table XXXI along with the relevant statistical parameters. Rigorous statistical analyses were not performed for the blanket assemblies, due to the incompleteness of data.

The data in Table XXXI may be used to calculate the confidence band on K at any values of the independent variables from the equation ( ):

O 5648A-643A-(514?l) 73

M AY = to i(V) Y (1+1/Nj=1+ E (X$ - XM)2/No .2)1/2 where t is Student's T statistic; o is the " standard error of estimate" from Table XXXI; y

N is the number of data points; M is the number of independent variables; Xj is the experimental mean of the 1-th independent variable from Table XXXI; X$ is the value of the i-th independent variable at which o is being calculated; and o xj is the standard deviation about X from Table XXXI.

Note that t is evaluated at some desired confidence level and hence, the resulting value of AY has associated the same confidence limit.

p The rod bundle is the only assembly component which can operate in all flow regimes. Since, however, only full turbulent flow is of interest in these studies, the uncertainty on the pressure drop correlation at 100% flow is lower than over the entire flow range, as indicated in Table XXXI. In fact, the data scatter in three separate flow tests (26-28) was greater in the transition region than at higher Reynolds numbers.

The recommended hydraulic correlations for fuel assembly components are presented in Table XXXII. All fuel assemblies are hydraulically identical except for the orifice stack which is unique to each flow orificing zone.

Selction of the proper orifice design from the five orifice correlations requires selection of both the number of plates and the hole diameter to achieve the zone flow rate. Selection should be done to minimize the number of orifice plates while avoiding cavitation which may damage hardware or

, affect pressure drop. While /inal selection must await the results of final i

l

{ \

d 5698A-343A-(S1421) 74

cavitation tests, in the present analyses o design criterionP) of 40 f t/sec maximum coolant velocity in the orifice holes was selected as a first approximation and the maximum achievable pressure drop for each orifice configuration was determined and is presented in Figure 77 as a function of assembly flow rate. Given an assembly flow rate, the pressure drop resulting from all components except the inlet nozzle / orifice / shield can be determined from the correlations in Table XXXII. This result subtracted from the total assembly pressure drop is the required inlet nozzle / orifice / shield pressure drop. The number of orifice plates for this pressure drop and flow rate can be determined f rom Figure 77 and the corresponding Reynolds number-dependent hydraulic correlation determined by solving for the orifice hole diameter.

The correlation presented for the rod bundle outlet results from a calculation of expansion form losses (33) , since no relevant data are available.

Because the associated pressure drop is of the order of 1 psi or less, the inherently large uncertainty of such a calculation, is acceptable for preliminary calculations. Data from the CRBR fuel assembly flow and vibration test will be used when available.

The correlation presented for the rod bundle inlet results from a calculation of expansion and contraction form losses (33) Although this component's pressure drop was measured (28) , the data are ambiguous since the

(*) Analyses were in progress at tne time of writing, to better define the cavitation margin criterion. It appears that the 40 f t/sec limit will be replaced by a maximum allowable cavitation number to be cubstantiated by testing. Figure 76 shows typical results being obtained in the cavitation tests of CRBRP fuel assembly orifice stacks. It should be noted, however, that the amount of pressure drop allocated in the orificing is practically independent of the particular no-cavitation criterion selected, which will only determine the geometry necessary to satisfy the no-cavitation limit.

Therefore, all the data reported in Section 6.3 for the various components pressure drop will vary only marginally following redefinition of the no-cavitation criterion.

O 5698A-643A-(S1421) 75

l C) - -experimentally determined form loss was negative. This could be due to either t

j V poor data or to the fact that the pressure taps were located in regions of not

' fully developed flow distribution. While the data scatter was excessive, it is, however, suspected that the latter was primarily responsible. Jets were observ,ed to issue from the top shield block and to impinge upon the support

' 2 grid; the dynamic head (pV /2) of the jet was 2 to 3 times the estimated grid pressure drop. For best accuracy, the rod bundle inlet pressure drop should be combined with the upstream or downstream components to avoid applying the standard pressure drop definition to regions where fully developed flow does not exist.

The recommended hydraulic correlations for the inner blanket assembly components are presented in Table XXXIII. The inner blanket assemblies utilize a fuel assembly type inlet nozzle, radial blanket type shield, rod bundle and outlet nozzle, and a unique orifice stack within the assembly.

The inlet nozzle geometry and hydraulic correlation are identical to the fuel assembly. The inlet nozzle-orifice-shield has not been tested, so the weak O

V Reynolds number dependence shown in Table XXXIII is recommended, based on the Reynolds number dependence of the four and five plate fuel assembly orifices (29) and the dependence of sample cases of the radial blanket inlet module orificing (30) The constant, C, represents the capability to design an orifice stack for any required pressure drop. The shield, rod bundle inlet and rod bundle outlet form losses were estimated based on Reference 33. The rod bundle friction was measured (31) over a sufficiently wide range of flow rates to cover the full range of inner blanket flow rates. The inner blanket assembly outlet nozzle form loss assumes sufficient similarity exists to apply the fuel assembly outlet nozzle test data (29) with the smaller blanket assembly reference area.

Except for the inlet nozzle and rod bundle friction, no data are available on the inner blanket assembly component pressure drops. However, because there is no a priori reason to anticipate worse uncertainties on inner blanket

i. assembly components than were found on fuel assembly components once tests are completed, it is recommended to apply the same uncertainties as for the (3._

(./

L

! Sf;98A-643A-(S1421) 76 L

corresponding fuel assembly components f ) (Table XXXI). In the case of the blanket rod friction, where data exist , the standard error on the recorrmended correlation is 5% over the ene. ire flow range and 3.3% over the full flow design range.

The recommended hydraulic correlations for the radial blanket assembly components are presented in Table XXXIV. The radial blanket assemblies have

' low rod bundle pressure drops and are dominated by orificing located both within the assembly and in the lower inlet module. The orificing located in the lower inlet module was characterized 0) over a range of flow rates near design conditions. A preliminarv examination of the test data shows the losses can be modeled proportional to Re -0.05 as was typically found for the four and five plate f uel assembl_y orifices (29) ,

The remainder 3f the radial blanket assembly inlet hardware has not been tested, so it is recommended that the combined inlet nozzle-orifice-shield and lower inlet module orificing, identified as " inlet region" in Table XXXIV, be modeled proportional to Re-0.05 The rod bundle inlet and outlet form losses were estimated based on the methods of Reference 33. The rod bundle friction factor was experimentally determined (31) over the full range of radial blanket flow rates and the inlet and outlet nozzle characteristics are based on a similarity extrapolation from the fuel assembly tests ( .

Following the previous logic, the inner blanket assembly component uncertainties aro recommended for use in radial blanket analyses. Final characterization of all inner and radial blanket assembly components over the necessary range of Reynolds numbers will result from the CRBRP blanket assembly flow and vibration test.

(*) It should also be kept in mind that the limiting pressure drop is represented by orificing zone 1, thus the blanket orificing can be easily adjusted to accomendate possible discrepancies between predictions and actual test data.

5608A-643A-(514?l) 77 O

\

r-jx -6.3 Results V

Pressure' drops were calculated for all reactor components by the newly developed CATFISH code. As previously mentioned, CATFISH models all the primary system resistances including the fuel, inner and radial blanket, radial shield and control assemblies, vessel cooling path and leakage as parallel flows. Below and above this parallel flow network are the inlet and outlet plena plus the primary loop resistances (check valve, piping, IHX).

The pump characteristics curve ties, and provides the boundary conditions to, the hydraulic network. The preceding Section 6.2 discussed modeling of the fuel and blanket assemblies resistances. Modeling of all the other resistances and of the pump characteristics are not discussed here for brevity since they are not, strictly speaking, directly related to the fuel and blanket assemblies. They will be discussed in detail in the CATFISH user's

~

manual to be issued in the future. CATFISH also has the capability of accounting for the effect of uncertainties associated with prediction of the various resistances; thus, for each calculation it shall be clearly stated p whether nominal resistances are considered or whether uncertainties are b accounted for and at what level of confidence. In addition to the hydraulic resistance, an uncertainty also exists on the pump head / flow characteristics where a 5% spread was specified between minimum and maximum pump head capability for a given flow.

Before presenting some typical results obtained with CATFISH, it is necessary to discuss how orificing requirements are considered in the overall picture of the-entire primary system hydraulics. A required flow allocation to the various orificing zones has been determined in Section 2 and has been used in calculating the thermal performance in Sections 4 and 5. The flow orificing can quantitatively be reduced to two requirements, i.e., fixed flow ratios between zone 1 and all the other flow orificing zones and amount of flow in zone 1.

NlQ.

LJ '

l' l

5698A-643A-(S1421) 78

Other defined constraints are: a) pump h acteristics curve; b) whether uncertainties are considered; and c) resistances in all flow paths. The designer has one degree of freedom, i.e., the resistance in the core assemblies orifices, or, more precisely, the resistance in orificing zone 1, since, once this is defined, the orifice resistance in all the other zones is automatically set to provide the required flow ratio among orificing zones.

It follows that, since all the other variables are defined by their geometry, operational characteristics (pump) or flow requirements, one-to-one correspondence between total reactor flow and zone 1 orificing resistance exists, which provides a unique solution to the primary system flow network.

The designer can thorefore specify the total reactor flow and obtain orificing resistances or specify an orificing resistance in zone 1 and obtain the total reactor flow. Examples will be shown .i the fol13 wing.

The first case considered was plant THDV conditions, where the flow is 6

specified (41.446 x 10 lb/hr). Since these are the worst possible conditions from the thermal performance standpoint, i.e., minimum reactor flow, it is consistent to assume positive uncertainties for the various hydraulic resistances and the minimum pump head curve. As the most conservative assumption, uncortainties on all resistances were taken at their maximum value, either 30 where a statistical basis existed, or at their boundary value (generally 1.2) where only an engineering estimate was available. It is rather obvious that assuming that all the resistances in all components are simultaneously at their maximum value, is extremely pessimistic, in fact, it is extremely improbable. Calculations were however performed for this set of conditions to provide an absolute minimum, worst case and the results are reported in Ta'ble XXXV. The zone flow reported in the table is an average for the zone (remember per previous discussion, that due to the fact that assemblies in the same orificing zone are in different LIMs, their flow will be slightly different). The components pressure drop correspond to the average assembly flow, thus again slightly different pressure drops are attained in the individual assemblies. Flow variations from assembiv-to-assembly in the same orificing zone are generally less than 11%, with maximum variations not exceeding 12%. All pressure drops are reported for cycle 4 conditions, which were the ones used in determining the orificing (see Sec'. ion 2). The only obvious exception is zone 6 inner O

5698A-643A-(S1421) 79

A

( -blanket, since only fuel assemblies are in zone 6 during cycle 4. Flow and _

(pressure' drops.for zone 6 IB are at cycle 2, conditions adopted for orificing of _ assembly #98, the only inner blanket remaining for two years in. zone 6.

The distinction between zones 10C and 10P is due to the 'f act that zone 10 assemblies are the only assemblies which belong to the same orificing zone, but. to two different types of modules. The modules at the core periphery (see Figure 78) have a dif, ent geometry from the other modules (called central) and therefore,'a different resistance and pressure drop. Assemblies 10C are the zone-10. assemblies located in the central modules, while assemblies 10P are located _in the peripheral modules. The lower inlet modules have been orificed for zones 10, 11 and 12 to provide the capability of shuffling radial blanket assemblies if so desired, as discussed in Section 6.1. Thus, the pressure drop reported under " LIM upper portion and orifice" for zones 1 through 9 is due to the hardware of the upper portion of the LIM, i.e., the region where the orificing plates are inserted for zones 10 through 12; the plates pressure _ drops are added for zones 10,11 and 12. Whether or not

- orificing plates are physically in the LIM is quite obvious from the relative

-h%~/

magnitude of the reported pressure drops. The " LIM, lower portion" AP is attributable to the resistance in the. remainder of the LIM. While the pressure drop in the upper portion depends on the individual assembly flow, the pressure drop in the remainder of the LIM depends on the total flow through the LIM, i.e., on the sum of the flows in the seven-assemblies fed by

,the LIM; It is not possible, therefore, to attribute the LIM, lower portion

. pressure drop to a given orificing zone, since the same orifice zone assemblies can be fed by different LIMs and the same LIM can feed different orifice zone assemblies. Pressure drops reported in Table XXXV under " LIM, lower portion" are therfore only indicative of the order of magnitude.

The average zone flows thus calculated for THDV conditions by CATFISH and reported _in Table XXXV agree (within 0.05%) with the flows reported in Table

-IV, which is quite a strong indication of the soundness of -the CATFISH model.

The: vessel nozzle-to-nozzle pressure drop was calculated as 126.4 psi, which is_ actually consistent with the 123 psi value commonly adopted, when one considers that ~ 126.4 psi -is- the vessel nozzle-to-nozzle irreversible pressure b_

-V loss, while 123 psi is the vessel nozzle-to-nozzle static pressure difference minus the nozzle-to-nozzle velocity head. The difference of 3.4 psi is 5698A-643A-(S1421) 80

therefore the difference in velocity head Y_ ween the reactor vessel inlet nozzle and outlet nozzle. Similarly, the value of the pump head, 163.5 psi, reported in Table XXXV, is the total developed head, while the commonly used value of 160.3 psi is the static pump head.

The most important point to be noted in examining the results in Table XXXV is the value of the orifice pressure drop in zone 1, i.e., 38.7 psi. As mentioned in Section 1, the zone 1 orifice pressure drop is by and large parasitic, since only a few psi are needed for final flow trimming; the pressure drop dissipated in the case just considered is a staggering 23.6% or nearly one-fourth of the entire pump head capability.

The next step is therefore rather obvious; i.e., decrease the parasitic pressure drop, thus increasing the total reactor flow and the flow through each assembly. The substantial payoff in thermal performance and economics has been outlined in Section 1. However, a limit exists en the present plant and pump specifications regarding the maximum amount of reactor flow, which was set at 115% of THDV conditions, or 47.663 x 106 lb/hr. It should be pointed out this is not a hard "no go" physical limit, but a design specification and its removal will require a lengthy reevaluation of the various impacts, which was deemed impractical at the present time. It was therefore decided that the orificing resistance will be reduced to the amount necessary to stay within the present pump " operating window", namely a flow 6

range of 41.446 to 47.663 x 10 lb/hr. Uncertainties of 1% in flow controller and 2% in flow measurements (calorimetric) were taken into account and conservatively were used in summation rather than in r.m.s. statistical combination. Thus, the nominal plant expected flow for the CRBRP was set at 112% of THDV conditions or 46.420 x 106 lb/hr. Consistent with the definition of nominal flow, nominal flow resistances and nominal pump head curve were adopted in the CATFISH analyses. The results are reported in Table XXXVI. One additional difference exists between conditions considered in Tables XXXV and XXXVI. The LIM orificing pressure drop reported in Table XXXV for zones 10,11 and 12 was the exact amount necessary to guarantee the same flow ratio among the various orificing zones as reported in Section 2. For a given orifice plate geometry, where the number of holes and hole diameters are defined and only the number of plates is variable, it is obvious that only a O

5698A-643A-(S1421) 81

^ 1 i , , , ,

.-'ractional.-number.of-plates'will f yield the exact value specified for the

- orifice pressure -drop. Thus, for the actual design' case, which is the one reported in. Table XXXVI', a finite number of LIM orifice plates had to be

'specified; the chosen number was the closest, by defect, integer to the fractional number which would have exactly satisfied tne orificing requirements . It was decided to r'ound the number of plates to the next lower rather than the next . higher integer because by so doing, the blanket would have been overcooled rather than undercooled. Since the fuel assemblies flows are~much higher than-the blanket assemblies flows, the reduction in fuel assemblies flow corresponding to a blanket overcooling is quite minor; thus, the overall flow allocation among the core assemblies is comparatively better than if the other alternative were chosen.

The pressure drop in orificing zone 1 is reduced to 17.6 psi, which is still 11.2% of the total head. These new plant conditions are, by their very definition, plant expected operating conditions and they represent a

.significant improvement over the previous conditions used in the thermal performance analyses reported in Section 4, with an increase in asse.nbly flows y of-the order 5-6%. An evaluation of system temperatures corresponding to these new flow conditions will soon be initiated. Analyses were conducted to evaluate +2a limits on the nominal expected plant flow. The hydraulic resistance in each component, e.g., zone (i) orificing, fuel bundle frictional resistance,' vessel, etc. was varied individually and the corresponding effect on the total reactor flow calculated. In total, the resistance in 83 components was varied. A root mean square of all the variations in reactor flow 'due to' the individual variation in hydraulic resistance (as well as pump head curve) at the 20 confidence level, resulted in a maximum 2a expected 6

plant flow'of 47.159 x 10 lb/hr, i.e., 1.6% higher than the nominal Lexpected flow. Thus, the total reactor flow was within the specified window

(*) Number of LIM orifice plates are as follows: 10 plates both in zone IOC

- and 10P (required fractional number is 9.8 in 10C and 10.3 in 10P); 7 plates in zone 11'(required number 7.6); 6 plates in zone 12 (required number 6.6).

Note that different hole diameters and number of holes are used in the various zones, which explains the apparent discrepancy that a larger number of orificing plates are used in the higher flow' zone and vice versa.

- 5698 A-643A-(S1421) 82.

even at the 02c leve l . Analyses will be updated as new evaluations of components pressure drop and relative uncertainties becme available f rom ongoing and plannea experimental tests.

The f act that over 10% of the pump head is still parasitic losses points out that ample room is lef t for improvement. To provide a quantitative evaluation, a CATFISH run was performed for nminal conditions and specifying no resistance in orificing zone 1. This resulted in a total reactor flow 6

equal to 48.445 x 10 lb/hr or #117% of THDV conditions, an additional 5%

flow increase over the updated nominal plant expected operating conditions with #18 psi parasitic AP. Whether this improvement can ever be utilized is speculative at present, however, it is important to know that it exists and is avai l ab l e.

O O

5698 A-643A-(51421 ) 83

f 7. ' CONCLUSIONS

\g% l Specific conclusions have been provided throughout th ss report. Rather than repeating previous statements, a very brief overview of the core T&H i analytical process is given here. As indicated in Sections 2 and 4, the current core T&H design, which will provide the basis for fuel and blanket

' assembly final usign, was successful since all the constraints were met.

This occurred since the core was orificed 'and flows allocated a priori, to guarantee satisf action of all criteria based on the adopted orificing philosophy which consisted of identifying all constraints, quantifying them in tems of comparable parameters and meeting these constraints. Obviously enough flow must be available to the core to utilize this orificing philosophy to its full advantage. This was the case for this analysis.

At the same time, a method to significantly increase, without associated penalty, the total plant flow was identified and pursued. Reduction (roughly halving) of parasitic pressure drops in the orifices would allow a flow increase of the order of 5%, with a comparable amount still available if the remaining parasitic pressure drop were eliminated altogether. The actual

%J thermal performance of the CRBRP core would therefore be better than reported here and significant margin exists to accomodate more restrictive constraints and more severe flow requirements which may be incurred in the future by the core thermofluids designers.

Improvements to existing analytical tools, development of new computer codes, f actoring of available experimental data in the analyses, implementation of new design approaches and concepts, extension of in-depth analyses, critical reevaluation and elaboration of design uncertainties have all contributed to a more realistic assessment of the core themofluids design and predicted

' performance.

It could be stated therefore, and this is perhaps the most important conclusion of this work, that the T&H design has satisf actorily fulfilled its scope, i.e., to provide a smooth translation of the core power distribution

'and nuclear characteristics into the core structural and mechanical behavior, ab 5698A-643A-(S1421) 84

This means more than a good thermal-hydraulic performance; it means that the T&H design, by integrating the nuclear design on one side and the structural design on the other, can highlight the strength and weaknesses implicit in both, point out potential problem areas and lead to an overall optimization of the integral core design.

Continuous improvements in core T&H design and analyses will contribute to minimize uncertainties germaine to the T&H functional link in the overall core design. To this end, several suggestions have been made throughout this report and they will be implemented to the maximum extent possible in the next design round for final, as-built conditions.

i e

l O

5698A-643A-(S1421) 85

REFERENCES 7 ~~

.4 l1. - M. D. Carelli, C. W. Bach, F. C.' Engel and D. R. Spencer, " Clinch River Breeder Reactor Plant; Predicted Thermal-Hydraulic Performance of CRBRP Fuel and. Blanket Assemblies", CRBRP-ARD-0054, August 1976 (Availability

US/ DOE Technical Information Center).

, 12. M. D. Carelli and R. A. Markley, " Preliminary Thermal-Hydraulic Design and Predicted Performance of the Clinch River Breeder Reactor Core", ASME paper _ 75-HT-71.

3 .' A. J. Friedland, "CRBRP Assemblies Hot Channel Factors Preliminary Analysis", WARD-D-0050, Rev. 3, September 1979.

4. _W. C. Chiang, "9CT9 PUS: A Flow Orificing Optimization Code for Fast Breeder Reactor Cores. Preliminary User's Manual", to be published as topical report.

5. M. D. Carelli and C. W. Bach, "LMFBR Core _ Thermal-Hydraulic Analysis Accounting for Inter-Assembly Heat Transfer", Trans. Amer. Nucl. Soc.,

2_8, pp. 560-562 (1978).

! 6.: M. D. Carelli and J. M. Willis, "An Analytical Method to Accurately Predict LMFBR Core Flow Distribution", Trans. Amer. Nucl. Soc., 32, pp. -

575-576 (1979).

7. M. D. Carelli, A. J. Friedland, C. W. Bach and R. A. Markley, "An-O- Optimized Method for Orificing LMFBR Cores", Trans. Amer. Nucl. Soc., 26, pp. 437-438 (1977).

8. E. H. Novendstern, " Mixing Model for Wire Wrap Fuel Assemblies", Trans.

Amer. Nucl. Soc., 15_, pp. 866-867 (1972).

9. M. D. Carelli, C. W. Bach-and R. A. Markley, " Analytical Techniques for Thermal-Hydraulic Design of LMFBR Assemblies", Trans. Amer. Nucl. Soc.,

1_7,, pp. 423-424 (1973).

10. J..V. Miller and R. D. Coffield, " FORE-2M: A modified Version of the

. FORE-II Computer Program for the Analysis of LMFBR Transients",

4 CRBRP-ARD-0142, May 1976. (Availability: US/ DOE Technical Information

. Center).

11. R. A. Doncals, et 'al., "CRBR Nuclear Design Data", to be published as topical report.

12. M. C. Chuang, M. D. Carelli, C. W. Bach and J. S. Killimayer, "Three-Dimensional Thermal-Hydraulic Analysis of Wire Wrapped Rods in

, Liquid Metal Fast Breeder Rea:: tor Core Assemblies",'Nucl. Sci. Eng., 64, pp. 244-257 (1977).

O s

5698A-643A-(S1421)- 86

13. F. C. Engel, R. A. Markley and B. Minushkin, " Heat Transfer Test Data of a 61-Rod Electrically Heated LMFBR Blanket Assembly Mockup and Their Use for Subchannel Code Calibration", in Fluid Flow and Heat Transfer Over Rod or Tube Bundles, pp. 223-229, ASME, December 1979.

14. A. S. Hanson and N. E. Todreas, " Fluid Mixing Studies in a Hexagonal 61-Pin Wire-Wrapped Rod Bundle", C00-2245-51TR, August 1977.

15. H. Hof fmann and E. Baumgartner, " Experimental Investigation of the Thermodynamic Behaviour of Fast Breeder Reactor Fuel Elements with Different Spacer Types", in Fuel and Fuel Element for Fast Reactors, Vol. 1, pp. 351-368, International Atomic Energy Agency, Vienna,1974.

16. K. Takahashi, E. Ishibashi, " Experimental Study on Coolant Sodium Mixing Effect in J0Y0 Fuel Assembly with Spiral Wire Spacer. Experimental Results on Core Fuel Subassembly", SN 941-75-77, July 1975.

17. L. Patch, R. M. Roidt, M. D. Carelli and R. A. Markely, " Experimental Studies of Flow Distribution in a Wire Wrapped LMFBR Blanket Assembly",

in Fluid Flow and Heat Transfer Over Rod or Tube Bundles, pp. 55-65, ASME, December 1979.

18. " Nuclear Systems Materials Handbook", TID-26666. (Availability: Hanford Engineering and Development Laboratory, Richland, WA).

19. M. D. Carelli, C. W. Bach and R. A. Markley, " Hydraulic and Scram Dynamics Analysis of LMFBR Control Rod Assemblies", Trans. Amer. Nucl.

Soc., 16, pp. 218-219 (1973).

20. M. C. Billone, et al., " LIFE-III Fuel Element Perfonnance Code",

ERDA-77-56, July 1977. (Availability: US/D0E Technical Information Center).

21. R. D. Leggett, E. O. Ballard, R. B. Baker, G. R. Horn and D. S. Dutt,

" Linear Heat Rating for Incipient Fuel Melting in UO2 - Pu02 Fuel",

Trans. Amer. Nucl. Soc., 1,5, pp. 752-753 (1972).

22. R. D. Leggett, R. B. Baker, D. S. Dutt and S. A. Chastain, " Influence of Burnup on Heat Rating-to-Melting for U02 - Pu02 Fuel", Trans. Amer.

Nucl. Soc.,19, pp.136-137 (1974).

23. G. H. Golden, A. Gopalakrishnan and R. A. Laskiewicz, " Correlation and Interpretation of Data Relative to EBR-II Power Level", in Irradiation Experimentation in Fast Reactors, pp. 314-332, American Nuclear Society, Hinsdale, Illinois.

24. L. B. Miller, G. H. Golden and R. E. Jabkac, " Characterization of the Power in an Experimental Irradiation Subassembly of Mixed-0xide Fuel in EBR-II", Ahl/EBR-047, September 1971.

25. G. H. Golden (EBR-II), private comunication with A. Biancheria (ARD),

1977.

9 5698A-643A-(S1421) 87

p. 26. " Covered Pressure Drop Flow Test / Cross Flow Mixing Test", HEDL-TI-76049, Q November 1976. (Availability: US/ DOE Technical Information Center).

27. W. L. Thorne, " Pressure Drop Measurements in FFTF Fuel Vibration Tests",

HEDL-TC-812, April 1977. (Availability: US/00E Technical Information Center).

28. W. L. Thorne, " Pressure Drop Measurements from Fuel Assembly Vibration Test II", HEDL-TC-824, April 1977. (Availability: US/D0E Technical Information Center).

29. P. M. McConnell, " Clinch River Breeder Reactor Fuel Assembly Inlet / Outlet.

Nozzle Flow Tests", HEDL-TME-77-8, February 1977. (Availability: US/ DOE TechnicalInformationCenter).

30. 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/D0E Technical Information Center).

31. F. C. Engel, R. A. Markley and A. A. Bishop, " Laminar, Transition and i Turbulent Parallel Flow Pressure Drop Across Wire Wrap Spaced Rod Bundles", Nucl. Sci, Eng., 69, pp. 290-296 (1979).

32. M. R. Spiegel, Schaum's Outline of Theory and Problems of Statistics,

p. 247, Schaum Publishing Co., New York, 1961.

33. I. E. Idel'chik, " Handbook of Hydraulic Resistance. Coefficients of Local Resistance and of Friction", AEC-TR-6630, 1960.

v 34 L. F. Moody, " Friction Factors for Pipe Flow", Trans. Amer. Soc. Mech.

Eng., 6_6_, pp. 671-684 (1944).

C) v 5698A-643A-(S1421) 88

- '-5 TABLE I CRBRP CORE LOADING DURING CYCLES 1 THROUGH 5 CYCLE TOTAL NUMBER NUMBER ASSEMBLY TYPE OF ASSEMBLIES REMARKS Fuel 156 All Fresh Assemblies 1 Inner Blanket 82 Radial Blanket 132 Fuel 159 3 Fresh Fuel Assemblies 2 Inner Blanket 79 Radial Blanket 132 Fuel 156 All Fresh Fuel and Inner 3 Inner Blanket 82 Blanket Assemblies Radial Blanket 132 e Fuel 162 6 Fresh Fuel Assemblies 4 Inner Blanket 76 (sd

"'$ Radial Blanket 132 Fuel 156 All Fresh fuel and Inner 5 Inner Blanket 82 Blanket Assemblies plus 60 Radial Blanket 132 Fresh Row 1 Radial Blanket Assemblies l

l l

D

)

6001A-661A 89 r

. . _ . . .- - _= ,

TABLE II SUfHARY OF STRAIN EQUIVALENT LIMITING TEMPERATURE CALCULATIONS FOR SECOND CORE FUEL ASSEMBLIES l ASSEMBLY E0L STRAIN (%) E0L PRESSURE (psi) SELT (UF) SET ( F) 2 .003 816 1264 1180 3 .007 845 1260 1192 13 .068 875 1258 1232 15 .031 891 1254 1211 24 .037 781 1270 1227 25 .039 861 1258 1220 26 .025 725 1281 1235 27 .057 7 48 1277 1247 28 .13 771 1272 1264 29 .003 815 1265 1181 30 .004 828 1262 1184 i 33 .001 855 1259 1194

34 .0086 862 1258 1193 37 .096 921 1249 1233 43 .015 892 1254 1197 44 .18 930 1248 1247 45 .034 876 1256 1217 47 .13 914 1251 1241 48 .12 773 1272 1261 49 .06 750 1276 1249 50 .023 726 1281 1232 51 .04 862 1258 1221 10 .19 920 1250 1250 11 .34 941 1247 1259 14 .36 928 1249 1263 36 .33 943 1247 1259 68 1.37 966 1243 #1295 101 1.43 970 1242 e1295 62 .029 529 not a limit 98 .018 488 not a limit l

9 6001A-661A 90

TABLE III COOLANT LIMITING TEMPERATURES FOR TELT CALCULATIONS (TEMPERATURES IN F)

. STEADY STATE TEMP. STEADY STATE TEMP.

AFMS MAXIMLN CORRESPONDING TO AFMS CORRESPONDING TO TRANSIENT TEMP. MAXIMUM TRANSIENT 15500F MAXIMUM ASSEMBLY TYPE TM (FORE-2M CALCULATED) TEMP. (FORE-2M) TRANSIENT TEMP.

Fuel Assembly 1571 1331 1316 1252 First Core (F/A #46 @ BOC1) 1261 Second Core Inner Blanket Assembly 1498 1247 1282 1198 First Core (IB/A #100 0 EOC4) 1207 Second Core I

Radial Blanket Assembly 1580 1331 1310 1232 (RB/A #212 @ E0CS) e G y - V Temperatures at THDV, 30, 7500F Inlet Temperatures for PE0V, 2a 6

6001A-661A 4

1

TABLE IV CORE ORIFICING ZONES FLOW ALLOCATION FLOW (lb/hr)

NO. ASSYS/ CYCLES CYCLE CYCLES ZONE TYPE ZONE 1,3,5,... 2 4,6,8,...

1 Fuel 39 189,990 (201,900) 188,520 (200,340) 187,050 (198,780) 2 Fuel 54 176,790 (187,870) 175,420 (186,420) 174,060 (184,970) 3 Fuel 21 166,900 (177,360) 165,610 (175,990) 164,320 (174,620) 4 Fuel 18 153,400 (163,020) 152,220 (161,760) 151,030 (160,500) 5 Fuel 24 149,480 (159,850) 148,330 (157,630) 147,170 (156,400)

Fuel 0,3 or 6 178,590 (189,780) 177,190 (188,300) 6 68,790 (73,100) 69,330 (73,680)

Inner Blanket 6,3 or 0 7 Inner Blanket 57 88,790 (94,360) 88,110 (93,630) 87,420 (92,900) 8 Inner Blanket 19 78,030 (82,920) 77,420 (82,270) 76,810 (81,620) 9 Radial Blanket 12 62,300 (66,210) 61,820 (65,700) 61,340 (65,190) 10 Radial Blanket 36 48,306 (51,330) 47,930 (50,930) 47,550 (50,530) 11 Radial Blanket 48 35,090 (37,290) 34,820 (37,000) 34,540 (36,710) 12 Radial Blanket 36 25,740 (27,350) 25,540 (27,140) 25,330 (26,920)

NOTE: Flows are for THDV (PE0C) conditions.

N CORE REGION FLOW FRACTIONS CYCLES CYCLE CYCLES REGION 1,3,5... 2 4,6,8...

Fuel 0.65 0.66 0.66 Inner Blanket 0.17 0.16 0.16 Radial Blanket 0.12 0.12 0.12 To tal 0.94 0.94 0.94 6001A-661A O O O

. _ . _ . . . _ - . . , . . . . .- _ . . . . . _ . . . _ . _ _ . . - _ , -,_ . - - . - . - _,, . . ,e .. .m.,.- _.

~

^

TABLE V-A CRBR FUEL ASSEMBLIES R00 TEMPERATURE ENGINEERING UNCERTAINTY FACTORS . '

. COOLANT - FILM CLADDING GAP FUEL ~ HEAT FLUX' DIRECT (o)J '

Power Level Measurement and Control System Dead Band .1.03(1' .0) . 03 Inlet Flow Ma1 distribution 1.05 .

2 Flow Distribution Calculational  ;}--------1.022

. Uncertainty (Simulation'81as) 1.03 .'

Cladding Circumferential Temperature Variation 1.0(9) ..1.0(9)  !

1.7 I.) '

STATISTICAL (3o)(0)--

Reactor at Variation' 1.0(1.144)

Wire Wrap Orientation 1.01

.Subchannel Flow Area 1.028 1.0 .

• Film Heat Transf er Coefficient 1.12 .

Pellet-Cladding. Eccentric ity 1.15 1.15

. Cladding Thickness and Conductivity 1.12. i Gap Conductance 1.48(4) 4 Fuel Conductivity 1.10' j e - , Coolant Properties 1.01

Flow Distribution Calculational Uncertainty (Calibration) 1.054----------1.015

(*) For cladding midwall temperature calculations. Applies to the nominal temperature drop between cladding midwall and culk coolant.

l (t) For fuel temperature calculations.

1 ($) Applies to BOL conditions.

i; (o) Nuclear Uncertainty Factors are given on Table V-B.

NOTE: Same values of subfactors apply to both Plant T&H and Expected Operating conditions except'when two values are given; in this case, the parenthesized values apply to Plant Expected Operating conditions while the non-parenthesized values apply to THDV conditions.

I l'

1

~

i l

i 6001A-661A h

TABLE V-B CRBR FUEL ASSEMBLIES ROD TEMPERATURE NUCLEAR UNCERTAINTY FACTORS WITH AND WITHOUT CONTROL ASSEMBLY INFLUENCE Coolant Heat Flux Direct (a)

Physics Modeling 1.02(1)

Control Rod Banking 1.09(2) (*)1.02(((.10)(1) 1,og )i ZPPR-7 Flux Tilt 1.0I4) 1.0l4 Statistical (3c)(a)

Nuclear Data 1.07 1.07 Criticality 1.01(3) 1.01(3)

Fissile Fuel Maldistribution 1.03 1.03 If assembly is influenced by adjacent control rod, replace with:

Coolant Heat Flux

" Peak Power Position" " Top of Core" BOL E0L BOL EOL BOL EOL Adjacent 1.04 1.02 1.03 1.02 1.15 1.15

1) Physics Modeling Far Side 1.01 1.02 0.95 1.02 1.30 1.15

. Adj acent 1.04 1.02 1.04 1.02 1.01 1.02 a 2) Control Rod Banking Far Side 1.02 1.02 1.02 1.02 1.01 1.02 Adj acent 1.04 1.04 1.C4 1.04 1.0 1.01

3) Criticality Far Side 1.01 1.01 1.C1 1.01 1.03 1.01

4) ZPPR-7 Flux Tilt - Assemblies 9, 10, 13, 14, 15,16,17, 23, 25, 37, 33, 41, 42, 43, 44, 45, 51, 53 # 0.97 9 BOL, 1.0 9 EOL). Assemblies 8,11,19, 36, 39, 47, 65, 68,101,104 (0.99 9 BOL,1.0 0 EOL). Assemblies 62, 98 (0.99 9 BOL,1.0 9 E0t ).

I Non-parenthesized value applies at the peak power position (i.e., core midplane). Parenthesized vah.e applies at the core lower / upper axial blanket interf ace except as superseded ::y note (1).

a) Engineering Uncertainty Factors are given on Table V-A EZ = Seginning of life EOL = End of life 6001A-661A O O O

. -_ . __ ._ ._. . - . . ~.

-TABLE VI-A-CRBR FUEL ASSEMBLIES MIXED MEAN EXIT TEMPERATURE ENGINEERING UNCERTAINTY FACTORS 4

ASSEMBLY EXIT Power. Level Measurement and Control System Dead Band 1.03(1.0)

Inlet Flow Maldistribution 1.05' STATISTICAL (30)(0)

Reactor AT Variation 1.0(1.144)

Coolant Properties 1.01

-(0) Nuclear Uncertainty Factors are given on Table VI-B.

NOTE: Same values of subf actors apply to both plant T&H and expected

operating conditions except when two values are given; in this case, the parenthesized values apply to plant expected operating' conditions while the non-parenthesized values apply to THDV conditions.

i

+

4 J

f

, O 6001A-661A 95

TABLE VI-B CRBRP FUEL ASSEM8 LIES MIXED MEAN EXIT TEMPERATURE NUCLEAR UNCERTAINTY FACTORS ASSEMBLY EXIT Physics Modeling (*)1.01(1.02 @ BOL, 1.01 @ E0L)

Control Rod Banking (*)1.02(1.03 @ BOL, 1.02 0 E0L)

ZPPR-7 Flux Tilt 1.0(4)

STATISTICAL (30)(a)

Experimental (Nuclear) 1.07 Criticality (*)1.01 (1.02)

Fissile Fuel Maldistribution 1.052 (4) ZPPR-7 Flux Tilt - Assy's. 9, 10, 13, 14, 15, 16, 17, 23, 25, 37, 38, 41, 42, 43, 44, 45, 51, 53 (0.97 @ BOL, 1.0 @ 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).

(*) Non-parenthesized values are applied for assemblies not adjacent to control assemblies. Parenthesized values are applied for the control assembly effect for assemblies adjacent to control assemblies.

(a) Engineering uncertainty f actors are given on Table VI-A.

O 6001A-661A 96

e s

~

TABLE VII-A

- CRBR FUEL ASSEMBLIES PLENLM PRESSURE i: ENGINEERING UNCERTAINTY FACTORS t

PLENUM TEMPERATURE BURNUP

! ., Power. Level Measurement and Control

. System Dead Band 1.02(1.0)' 1.02

' Inlet' Flow Maldistribution 1.05 Flow Distribution Calculational Uncertainty.~(Simulation Bias) 1.03

' STATISTICAL ~(30)(0)

Reactor AT Variation- 1.0(1.144)

{l Wire Wrap' Orientation

[ 1.01 ,

! Coolant Properties 1.01

Flow

Distribution Calculational 3 Uncertainty (Calibration) 1.085 1

-(0) Nuclear. Uncertainty Factors are given:on Table VII-B NOTE: -Same values of subfactors apply to both plant T&H and expected operating conditions except when two' values are~given; in this case, ,

l- the parenthesized values; apply to plant expected operating conditions 1 l while the non-parenthesized values apply to THDV conditions.

I i- 1 F

i 4

i-LO 6001A-661A 97

~. .,

...--..:-..--.-..-...---.-.-... --.-,-..-,-.,w.-.,,,

TABLE VII-B CRBR FUEL ASSEMBLIES PLENUM PRESSURE NUCLEAR UNCERTAINTY FACTORS WITH AND WITHOUT CONTROL ASSEMBLY INFLUENCE PLENUM TEMPERATURE BURNUP Physics Modeling 1.02(I) 1.02(1)

Control Rod Banking 1.02(2) 1.02(2)

ZPPR-7 Flux Tilt 1.0(4) 1.0(4)

STATISTICAL (3o1(0)

Nuclear Data 1.07 1.07 Criticality 1.01(3) 1.01(3)

Fissile Fuel Maldistribution 1.03 1.03 If assembly is influenced by adjacent control rod, replace with:

BOL E0L Adj acent 1.04 1.02

1) Physics Modeling Far Side 1.01 1.02 Adj acent 1.04 1.02

2) Control Rod Banking Far Side 1.02 1.02 Adjacent 1.04 1.04

3) Criticality far Side 1.01 1.01

4) ZPPR-7 Flux Tilt - Assemblies 9, 10, 13, 14, 15, 16, 17, 23, 25, 37, 38, 41, 42, 43, 44, 45, 51, 53 (0.97 @ BOL, 1.0 0 E0L). Assemblies 8, 11, 19, 36, 39, 47, 65, 68, 101, 104 (0.99 @ BOL,1.0 0 EOL ). Assemblies 62, 98 (0.99 @ BOL, 1.0 0 E0L).

(0) Engineering Uncertainty Factors are given on Table VII-A.

O 6001A-661A 98

- v- V v TABLE VIII-A "R8R INNER / RADIAL.8LANKET ASSEMBLIES R00 TEMPERATURES ENGINEERING UNCERTAINTY FACTORS ,

COOLANT FILM. CLADD1hG E FUEL HEAT FLUX

' DIRECT (0)

Power Level Measurement and Control System Dead Band 1.03(1,0)- 1.03

'!nlet Flow Maldistribution ~ 1.07 '

Flow Distribution Calculational }---------1.03 Unce'rtaint) (Simulation Bias) 1.03 Clacaing Circumferential Temperature Variation 1.0(t) (,) 1.0(t)

STATISTICAL (3o)(0)

Reactor AT Variation 1.0(1.144)

Wire Wrap Crientation - 1.01 Subchannel F'ow Area 1.035 1.0 Film Heat. Transfer Coefficient 1.21 Pellet-Cladding Eccentricity 1.15 1.15 Cladding Thic< ness and Conductivity 1.12 Gap Conductance 1.48(,)

Fuel Conductivity 1.10 o

Coolant Properties 1.01.

Flow Distribution Calculational Uncertainty (Calibration) 1.199/1.1(+)- 1.056.

(*) For cladding midwall temperature calculations. _ Applies to the nominal temperature drop between cladding midwall and bulk coolant.

t) For fuel temperature calculations.

$) Applies to BOL conditions.

o) Nuclear Uncertainty Factors are given on Table VIII-8.

(+) Inner / radial blanket.

NOTE: Same values of subfactors apply to both Plant TI,H and Expected Operating conditions except when two values are given; in this case, the parenthesized values apply to Plant Expected Operating conditions while the non-parenthesized values apply to THDV conditions.

6001A-661A I

i Tr.3LE VIII-B CRBR INNER / RADIAL BLANKET ASSEMBLIES ROD TEMPERATURE NUCLEAR UNCERTAINTY FACTORS INNER BLANKET RADIAL BLANKET Row 1 Row 2 COOLANT HEAT FLUX COOLANT HEAT FLUX COOLANT HEAT FLUX DIRECT (80L)I )

Physics Modeling 1.06 (*)1.07(1.11) 1.07 I*)1.07(1.02) 1.03 (*)1.07(0.99)

Control Rod Banking 1.02 1.02 1.02 1.02 1.02 1.02 Nuclear Data 1.12 (*)1.10(1.17) 1.13 (*)1.11(1.18) 1.27 I

(*I.24(1.32)

Criticality 1.02 1.02 ---- ---- ---- ----

Heavy Meta 1 1.01 '.01 1.01 1.01 1.01 1.01 U-235 1.01 1.01 1.01 1.01 1.01 1.01 8

DIRECT (E0L)IO)

Physics Modeling 1.02 (*)1.04(1.12) 1.01 I*)1.07(1.05) 1.01 I*)1.07(1.02)

Control Rod Banking 1.02 1.02 1.02 1.02 1.02 1.02 Nuclear Data 1.03 (*)1.00(1.07) 1.05 (*)1.03(1.10) 1.15 (*)1.13(1.21)

Criticality 1.01 1.01 ---- ---- ---- ----

Heavy Metal 1.01 1.01 1.01 1.01 1.01 1.01 U-235 1.0 ---- ---- ---- ---- ----

(*) Non-parenthesized values apply at the peak power position (i.e., near core midplance). Parenthesized values apply at core upper / lower axial extension interf ace.

(o) Engineering Uncertainty Factors are given on Table VIII-A.

6001A-661A O O O

TABLE IX-A

(-')

CRBR INNER /RADI AL BLANKET ASSEMBLIES MIXED MEAN EXIT TEMPERATURE ENGINEERING UhtERTAINTY FACTORS DIRECT (0)

Power Level Measurement and Control. System Dead Band 1.03(1.0) 1 Inlet Flow Maldistribution 1.07 STATISTICAL (3o)(0)

Reactor AT Variation 1.0(1.144)

Coolant Properties 1,01 (0) Nuclear Uncertainty Factors are given on Table IX-B.

NOTE: Same values of subf actors apply to both plant T&H and expected operating conditions except when two values are given; in this case, the parenthesized values apply to plant expected operating conditions while the non-parenthesized values apply to THDV conditions.

e i

I e

O 6001A-661A:

101

TABLE IX-B CRBR INNER / RADIAL BLANKET ASSEMBLIES MIXED MEAN' EXIT TEMPERATtRE NUCLEAR UNCERTAINTY FACTORS INNER BLANKET RADIAL BLANKET Row 1 Row 2 DIRECT (BOL)IO)

Physics Modeling 1.01 1.02 1.02 Control Rod Banking 1.02 1.02 1.02 Nuclear Data 1.12 1.13 1.27 Criticality 1.02 ---- ----

Heavy Metal 1.01 1.01 1.01 U-235 1.01 1.01 1.01 DIRECT (E0L)IO)

Physics Modeling 1.01 1.02 1.02 Control Rod Banking 1.02 1.02 1.02 Nuclear Data 1.03 1.05 1.15 Criticality 1.01 ---- ----

Heavy Metal 1.01 1.01 1.01

.-235 ---- ---- ----

I ) Engineering uncertainty factors are given on Table IX-A.

O 6001A-661A 102

TABLE X-A

'i CR8". INNER / RADIAL BLANKET ASSEMBLIES PLENUM PRESSURE Ll ENGINEERING UNCERTAINTY FACTORS PLENUM TEMPERATURE BURNUP Power' Level Measurement ~and Control System Dead Band 1.03(1.0) 1.02 Inlet Flow Maldistribution 1.07 Flow Distribution Calculational

- Uncertainty (Simulation Bias) 1.03 5

STATISTICAL (3a)(0)

Reactor AT-Variation 1.0(1.144)

Wire Wrap Orientation ' 1.01

- Coolant Properties 1.01 Flow Distribution Calculational Uncertainty (Calibration) 1.299/1.1(+)

(0) Nuclear Uncertainty Factors are given on Table X-B

(+) Inner / radial blanket NOTE: Same values of subf actors apply to both plant T&H and expected operating conditions except when two values are given; in this case, the parenthesized values apply to plant expected operating conditions

while the non-parenthesized values apply to THOV conditions.

1 4

O 6001A-661A-103 4

- .;-m - - - , . , . . . . .,,ip- - --w -e .-- - - . , - , , - - - , - , , - - ,

TABLE X-B CRBR INNER / RADIAL BLANKET ASSEMBLIES PLENUM PRESSURE NUCLEAR UNCERTAINTY FACTORS INNER BLANKET RADIAL BLANKET Row 1 Row 2 F' ENUM TEMPERATURE BURNUP PLENUM TEMPERATURE BURNUP PLENUM TEMPERATURE BURNUP DIRECT (BOL)(0)

Physics Modeling 1.06 1.06 1.03 1.03 1.03 1.03 Control Rod Banking 1.02 1.02 1.02 1.02 1.02 1.02 Nuclear Data 1.12 1.12 1.13 1.13 1.27 1.27 Criticality 1.02 1.02 ---- ---- ---- ----

Heavy Metal 1.01 1.01 1.01 1.01 1.01 1.01 U-235 1.01 1.01 1.01 1.01 1.01 1.01 8

DIRECT (EOL)(0)

Physics Modeling 1.02 1.02 1.01 1.01 1.01 1.01 Control Rod Banking 1.02 1.02 1.02 1.02 1.02 1.02 Nuclear Data 1.03 1.03 1.05 1.05 1.15 1.15 Criticality 1.01 1.01 ---- ---- ---- ----

Heavy Metal 1.01 1.01 1.01 1.01 1.01 1.01 U-235 ---- ---- ----

(0) Engineering uncertainty factors are given on Table X-A.

l

6001A-661A O -

O O

)

% f w s ' TABLE XI.

1 FUEL ASSEPELIES UNCERTAINTY. FACTORS FOR DUCT TEMPERATURE CALCULATIONS Direct. ' ' eat Generation .. Assembly F'ow g Film Duct. Interstitial Gap

' Power Level Measurement and 1.03(1.03)

Control System Dead Band Inlet Flow Ma1 distribution 1.05(0.95? (1)-

' Flow Distribution Calculational. 1.03(1.03) (1)

Uncertainty (Simulation Bias)L ,,'

Nuclear See Table XII

. Statistical-(3o)' ,

Reactor AT variation 1.0/1.144(2)

(1.0/1.144)

Nuclear See Table XII Wire Wrap Orientation. 1.01(0.99)

, Subchannel Flow Area 1.071(0.9)

O ' Flow Distribution Calculational 1.08(0.92)(3)

Uncertainty ' Calibration)

• Geometrical Deviations- (4) (4)

Duct Conductivity 1.1(1.1)

Film Heat Transfer Coefficient (1) 1.0 Coolant Properties 1.017(1.017) 1.017(1.017)

-NOTES

Numbers outside parentheses 'are positive uncertainties, in parentheses are negative uncertainties.

, (1) Bracketed by: KNa/6(=) - see Section 3.2.2.4 for explanation.

(2) Plant. thermal-hydraulic design value/ expected operating conditions. I (3) Negative value to be used against a different type of assembly, e.g., fuel adacent to blanket, not fuel to fuel.

- unless uniformly applied to all assemblies of the same type. See Section 3.2.2.3 tur further explanation.

(4) Variations of dimensions due to tolerances, etc., are explicitly considered in TRIT 9N by changing value of inter-assemblies pitch.and ducts thickness.

6001A-661A I

TABLE XII FUEL ASSEMBLIES HEAT GENERATION NUCLEAR UNCERi AINTIES Group I - Assemblies #10, 14, 15, 37, 43, 44 POSITIVE NEGATIVE Direct BOL E0L BOL E0L Physics Modeling 1.02 1.01 0.98 0.99 Control Rod Banking 1.03 1.02 1.03 1.02 ZPPR-7 Flux Tilt 0.97 1.0 0.97 1.0 Statistical (3a)

Experimental 1.07 1.07 0.93 0.93 Criticality 1.02 1.02 1.02 1.02 Fissile Maldistribution 1.03 1.03 0.97 0.97 Total Oo 1.019 1.030 0.979 1.010 20 1.073 1.084 0.931 0.961 30 1.10 1.111 0.908 0.936 Group II - Assemblies #62, 98 Direct Physics Modeling 1.02 1.01 0.98 0.99 Control Rod Banking 1.03 1.02 1.03 1.02 ZPPR-7 Flux Tilt 0.99 1.0 0.99 1.0 Statistical Same as Group I Total Oo 1.040 1.030 0.999 1.010 20 1.095 1.084 0.950 0.961 30 1.122 1.111 0.926 0.936 i

l O

6001A-661A 106

TABLE XII (CONT'D)

G_roup III - Assemblies #9, 13, 16, 17, 23, 25, 38, 41, 42, 45, 51, 53 POSITIVE NEGATIVE Direct BOL E0L BOL EOL Physics Modeling 1.01 1.01 0.99 0.99 Control Rod Banking 1.02 1.02 1.02 1.02 ZPPR-7 Flux Tilt - 0.97 1.0 0.97 1.0 Statistical (3a)

Experimental 1.07 1.07 0.93 0.93 Criticality 1.01 1.01 1.01 1.01 Fissile Ma1 distribution 1.03 1.03 0.97 0.97 Total Oc 0.999 1.030 0.980 1.010 20 1.050 1.083 0.931 0.959 3a ' 1.076 1.109 0.907 0.934 Group IV - Assemblies #8, 11, 19, 36,'39, 47, 65, o8, 101, 104 Direct O Physics Modeling Control Rod Banking 1.01 1.01 0.99 0.99 1.02 1.02 1.02 1.02 ZPPR-7 Flux Tilt 0.99 1.0 0.99 1.0 ,

Statistical I i

Same as Group III  !

Total Oo 1.0 20 1.030 1.0 1.010 2o 1.072 1.083 0.95 0.959 30 1.099 1.109 0.925 0.934 )

l l

y V

6001A-661A 107

TABLE XII (CONT'D)

Group V - Assemblies #2, 3, 26, 27, 28, 29, 30, 33, 34, 48, 49, 50 POSITIVE NEGATIVE Direct BOL & E0L BOL & E0L Physics Modeling 1.01 0.99 Control Rod Banking 1.02 1.02 Statistical Same as Group III Total Oo 1.030 1.010 20 1.083 0.959 3o 1.109 0.934 O

O 6001A-661A 108 I

[

-t/( a v

-TA8LE XIII BLANKET ASSEMBLIES UNCERTAINTY FACTORS FOR-DUCT TEMPERATURE CALCULATIONS Direct Heat Generation Assembly Flow g Film Duct Interstitial Gap Power Level Measurement and- 1.03(1.03)

? Control System Dead Band ;

Inlet Flow Maldistribution 1.07(0.93) (1)

Flow Distribution Calculational 1.03(1.03) (1)

Uncertainty (Simulation Bias)

Nuclear See Table XIV.

+

1

-Statistical (3a) {

Reactor AT Variation 1.0/1.144(2)

(1.0/1.144) <

Wire Wrap Orientation 1.0 Subchannel Flow Area 1.15(0.75)

Flow Distribution Calculational 1.2(0.8)(3)

_. Uncertainty (Calibration) o Geometrical Deviations' (4) (4)

Duct Conductivity 1.1(1.1)

Film Heat Transfer Coefficient (1) 1.0 Coolant Properties 1.017(1.017) 1.017(1.017)

NOTES:

Numbers outside parentheses are positive uncertainties, in parentheses are negative uncertainties.

(1) Bracketed by: KNa/6(=) - see Section 3.2.2.4 for explanation.

(2) Plant thermal-hydraulic design value/ expected operating conditions.

(3) Negative value' to be used against a different type of assembly, e.g., fuel adjacent to blacket, not fuel to fuel, unless uniformly applied to all assemblies of the same type. See Section 3.2.2.3 for further explanation.

(4) Variations of dimensions due to tolerances, etc., are explicitly considered in TRIT 9N by changing value of inter-assemblies pitch and ducts thickness. L

.T 6001A-661A

TABLE XIV INNER AND RADIAL BLANKET ASSEMBLIES HEAT GENERATION NUCLEAR UNCERTAINTIES Factor Inner Blanket Radial Blanket First Row Radial Blanket Second Row (all factors are direct) Positive Negative Positive Negative Positive Neaative BOL EOL BOL lE0L BOL EOL BOL E0L BOL EOL BOL EOL Experimen ta1 1.12 1.03 0.92 0.93 1.13 1.05 0.95 0.91 1.27 1.15 0.99 0.99 Heavy Metal Content 1.01 1.01 0.99 0.99 1.01 1.01 0.99 0.99 1.01 1.01 0.99 0.99 U-235 Content 1.01 1.0 0.99 1.0 1.01 1.0 0.99 1.0 1.01 1.0 0.99 1.0 Modeling 1.06 1.02 0.94 0.98 1.07 1.01 0.97 0.95 1.03 1.01 0.93 0.95 Criticality 1.02 1.01 1.02 1.01 1.0 1.0 1.0 1.0 1.0 1.0 1.0 1.0 Control Rod Banking 1.02 1.02 1.02 1.02 1.02 1.02 1.02 1.02 1.02 1.02 1.02 1.02 TOTAL 1.26 1.09 0.88 0.93 1.26 1.09 0.92 0.87 1.36 1.20 0.92 0.95 o

6001A-661A O O O

I

[ ("% . f)

( d TABLE XV PRIMARY CONTROL ASSEMBLIES UNCERTAINTY FACTORS FOR DUCT TEMPERATURE CALCULATIONS Direct Heat Generation Assembly Flow Flow Split 0" bundle A" bypass Film puct Interstitial Gap Power Level Measurement' and 1.03(1,03)

Control System Dead Band Inlet Flow Ma1 distribution 1.08(0.92) (1)

Flow Distr'ibution Calculation 1.08(0.92)(3) (1}

Uncertainty Bundle / Bypass Flow Split 1.l(1.0) (1)

Nuclear' 1.15(0.85)

Statistical (3a)

Reactor aT Variation 1.0/1.144(2) 1.0/1.144(2)

(1.0/1.144) (1.0/1.144)

Absorber Maldistribution 1.03(0.97)

] Wire Wrap Orientation 1.0 Subchannel Flow Area 1.16(0.87) 1.23(0.89)

Geanetrical Deviations (4) (4)

Duct Conductivity 1.1(1.1) ,

Film Heat Transfer Coefficient .(1) 1.0 Coolant Properties 1.017(1.017) 1.017(1.017) 1.017(1.017)

NOTES:

Numbers outside parentheses are positive uncertainties, in parentheses are negative uncertainties.

(1) Bracketed by: KNa/6(=) - see Section 3.2.2.4 for explanation.

(2) Plant themal-hydraulic design value/ expected operating conditions.

(3) Negative value to be used against a different type of assembly, e.g., fuel adjacent to blanket, not fuel to fuel, unless uniformly applied to all assemblies of the same type. See Section 3.2.2.3 for further explanation.

(4) Variations of dimensions due to tolerances, etc., are explicitly considered in TRIT 9N by changing value of inter-assertblies pitch and ducts thickness.

6001A-661A i

TABLE XVI CRBR EXPECTED OPERATING CONDITIONS DURING PLANT LIFETIME Clean & Unplugged Heat Fouled & Plugged Heat Exchangers Exchangers (New Plant) Estimated (2 Year Fouling) (30 Year Fouling?

Parameter Nominal Mean o T Nominal Mean o T Nominal Mean a T 97.7 97.7 97,7 Primary Hot Leg 943 946 13 968 950 954 13 976 960 964 13 987 Tre:perature (OF)

Primary Cold Leg 698 697 13 722 705 704 11 725 714 714 12 736 Temperature (OF)

Primary AT (OF) 245 249 12 273 245 250 12 274 246 250 12 275 Power (M4t) 975 975 1004 975 975 1004 975 975 1004 NOTE: Design and control uncertainties are included.

}

6001A-661A l

O O O

l

)- TABLE XVII PLANT EXPECTED CONDITIONS AND ASSOCIATED UNCERTAINTIES

, CONSIDERED ON CRBRP CORE THERM 0 FLUIDS ANALYSES s

First Core Second Core Nominal Inlet' Temperature (OF) 704 711 Nominal Reactor AT (DF) 250 250

, Reactor AT Uncertainty Factor (3a)(*) 1.144 1.144 Inlet Temperature Uncertainty (30) (OF) 33 36 Loop-to-Loop Imbalance (30) (OF) 4.6 4.6 Combined Uncertainty on Inlet Temperature

, (20) (OF) - (Additive to Nominal) 22.2 24.6

(*) To be combined with other uncertainties on FAH

, NOTE: Power Level Measurement / Control Dead Band Uncertainty is already

included in above uncertainties.

')

?

O 6001A-661A 113

TABLE XVIII GEOMETRIES AND OPERATING CONDITIONS (Oo LEVEL) CHARACTERIZING FUEL PINS USED IN POWER-TO-MELT UNCERTAINTY ANALYSIS F/A 101 F/A 14 PARAMETER (X/L = 0.409) (X/L = 0.318)

Cladding Length, in. 77.19 77.19 Fuel Length, in. 36.24 36.24 Cladding Inner Radius, in. 0.100 0.100 Cladding Thickness, in. 0.015 0.0 %

Fuel-Cladding Radial Gap, mils 3.25 3.25 Reactor Inlet Temperature, OF 735 735 Initial Gas Pressure, atm. 1.0 1.0 Initial Xenon Fraction 0.094 0.094 Fu'l e Density, % T.D. 91.3 91.3 Initial Pu02 , w/o 32.9 32.9 BOC1 Power, kw/ft 13.23 12.77 B0C1 Flux (E > 0.1) (x 1015) 3.18 3.29 BOC1 C1 adding 00 Temp., OF 960 875 B0C1 Plenum Temp., OF 1171 1082 9

6001A-661A 114 l

l l

. _ _ _ _ _ _ _ _ _ - - - - _ - - _ _ _ _ _ _ _ _ _ _ _ _ . - --___ . . . _ _ _ . - _ _ =_ _ . . - - . . __. _ . . . . _ - . - . _ _ . _ . _ .

L O O O '

TABLE XIX.

POWER-TO-MELT RESULTS -

l MAXIMlM LINEAR POWER SELT OVERPOWER

~ ASSEMBLY (PEAK PIN) RATING, Oo (kw/ft) HISTORY (kw/ft) (% NFP)  !

F/A 101 13.23 8 hrs. to 0.7 NFP(*) 14.70- 11.1 I

!' 13.23 above + 50 hrs. to 0.9 NFP 17.41 -31.5 13.23 above + 50 hrs, to NFP 18.39 39.0 i

i F/A 14 12.77 8 hrs, to 0.7 NFP 14.97 ,

17.2 -

12.77 above + 50 hrs. to 0.9 NFP 17.24 35.0 12.77 above + 50 hrs, to NFP 18.42 44.2 f

i

(*) Nominal Full Power i

i I

f i

i i

f

6001A-661A

TABLE XX CENTER TEMPERATURES AND POWERS-TO-MELT OF INCIPIENT HELT SECTIONS AS COMPUTED BY LIFE-3 Calculated Observed Calculated Tmax Power-to-Mel t Power-to-Melt Pin No. OF (OC) kW/ft (kw/m) kW/ft (kw/m)

P19-35 5009 (2765) 16.62 (54.5) 16.69 (54.6)

P19-02 4978 (2748) 16.04 (52.6) 15.89 (52.1)

P19-24R 5024 (2773) 15.44 (50.6) 15.59 (51.1)

P19-27R 4990 (2754) 19 .23 (63.1) 19.16 (62.8)

P19-28 4966 (2741) 19.02 (62.4) 18.82 (61.8)

P19-30 5070 (2799) 17.22 (56.5 ) 17.79 (58.3)

P19-08 5006 (2763) 15.36 (50.4) 15.40 (50.5)

P20-07 5004 (2762) 15.72 (51.7) 15.75 (51.7)

P20-30 4985 (2752) 16.82 (55.2) 16.70 (54.8)

P20-33 4982 (2750) 19.07 (62.5) 18.98 (62.2)

Average T = 5001 1 28 F r.m.s. error (2761 1 16 C) = 0.21 kw/ft (.69 kw/m) std. dev.*

= 0.24 kw/f t (.77 kw/m)

OTwo degrees of freedom lost in calibrating LIFE-3 parameters.

6001A-661A O

116 i

i TABLE XXI SPATIAL DEPENDENCE.0F POWER UNCERTAINTY CALCULATED FROM

COMPARISON OF MEASlRED AND CALCULATED CONTROL R00 WORTHS

!' Control

- Calculated Minus Rod Number Measured Power (%)

i 1 -0.67 ,

F 2 -1.75

j. 3 -1.8

4 -2.35 5

0.

6  !

1.85 7

2.25 8 1.5 900 0.45 4

10 1.2 11 0.7 '

i 12 2.25

O 1 Average A Power = 0.3%

Standard Deviation = 1.7%

l 1

i t

1 r

1-LO 6001A-661A 117 I' ,

,,---,y.- --- , w,,*, - - - ...-,r- ,m.. , --.4.-r.----,%-,4---w-.--e--.--.~,,,..-.,,----.-vc . .. , ,, . . .- . - - - ., -,...,.,-e .. w ,.,,,m _.,-rs,*m e-m--.

TABLE XXII T

ADOPTED VALUES FOR EBR-II UNCERTAINTIES O

atime 2.2

! asys 3.0 o space 1.8 Ofab 0.8 OPIE 0.9 Otot 4.3 i

Umean 4.1 l

1 l

9 l

l l

l 6001A-661A l 118 l

._ = -_ . - - - _ - - , -

TABLE XXIII CRBRP FUEL R00 TOLERANCES AND UNCERTAINTIES Nominal /Joj Fuel Pellet Diameter 0.1935 in. 1 0.0015 in.

Cladding Inside Diameter -'

O.200 in. 1 0.0005 in.

Fuel Pellet Density 91.3% TD 1 1.6% TD Reactor Power

• 1 5.07%

C1 adding 0.D. Temperature Near Midplane ** -+ 15.60F

• Statistical combination of nuclear data, criticality, and fissile fuel maldistribution uncertainties. Value reported was calculated on the basis of preliminary nuclear uncertainties. A recalculation was not deemed O necessary at this time since updated nuclear uncertainties are not too b dissimilar, considerable margin-to-melting exists and programed startup parameters are not yet optimized.

• • Due to thermal-hydraulic engineering uncertainties (not including uncertainties already accounted for in reactor power).

i il J

6001A-661A 119 i

+ ---- , e - - - w -

, - , . - - . - - . - - - , , , , ---n-y n.---.w---- e ,-

TABLE XXIV SENSITIVITY OF POWER-T0-MELT TO DESIGN UNCERTAINTIES CHANGE IN POWER-T0-MELT FOR ONE PERCENT CHANGE IN DESIGN UNCERTAINTY 3REPOW g /a (%X$ )

U ME , t.

P O O T hrs REPOW i pel pel clad Power clad 8 0.7 .0040 .112 .098 .0107 .00092 58 0.9 .0017 .043 .062 .0027 .00069 108 1.0 .0051 .051 .060 .0093 .00081 158 1.0 .0051 .050 .060 .010 .00081 i

l O

f

• Derivative taken from Figure 76.

l O

6001A-661A 120

i

4. '[

f TABLE XXV STATISTICAL COMBINATION OF POWER-TO-MELT UNCERTAINTIES DUE TO DESIGN UNCERTAINTIES 4

ONE SIGMA POWER UNCERTAINTY  !

RESULTING FROM DESIGN UNCERTAINTY

! STATISTICAL

"ia EPOWg /B X i COMBINATION TIME, t hrs. P pel D D T U

[

'REPOW pel clad Power clad _E i

f 8 0.7 .0041 .050 . .014 .0337 .0036 .062 l 58 0.9- .0017 .019 .0089 .0085 .0027 .023

. 108 1.0 .0052 .023 .0087 .0294 .0032 .039 158 1.0 .0052 .022 .0087 .0315 .0032 .040 a.

i i

1 4

i

< O

! 6001A-661A 121 I.

!1_.

L .. . .- . _ - . _ . _ . _ _ . . _ . - - - - _ . - - - _ _ _ . _ . _ _ - - - - -

e TABLE XXVI EFFECT OF MELTING POINT UNCERTAINTY OF 0

TM = 5.8 F ON UNCERTAINTY OF POWER-TO-MELT 0 M/8T M U extrap U

t(hrs) kw/ft/ F (kw/ft) (REPOW Units) 8 .0095 .055 .0043 58 .0069 .040 .0031 108 .0061 .035 .0027 158 .0058 .034 .0026

l O

O 6001A-661A -

122

L i

O O O TABLE XXt'II .

TOTAL UNCERTAINTY ON CRBRP REACTOR POWER-TO-MELT  ;

IN REACTOR POWER UNITS WHERE NOMINAL FULL POWER IS 1.0

i REPOW - 3aREPO -

REP 6W l TOTAL o M M REP 0W

! t (hrs) -REPOW R mean "extrap M (nominal, t) -1.15/1.03 REPOW l'

8 0.7 .062 .047 .0043 .078 1.152 .136 i

58 0.9 .023 .055- .0031 .060 1.335 .150 108 1.0 .039 .058 .0027 .070 1.415 .088 158 1.0 .040 .059 .0026 .071 1.437 .107

! g -

l i

i t

1 i

t a  :

I, i 1 6001A-661A i 1

4 I

,,, .y ,- . . - , - , - -

TABLE XXVIII COMPARISON OF 3-0 MINIMJM POWER-TO-MELT TO POWER AT 15% OVERPOWER Q1* Q2

• 3 Power at 3a Margin =

Omelt REPOW g Q4=

Start of Jump 1.15 Q l Q4 -1.15 Q I ASSEMBLY NFP* T.0J NFP* (Table XXVII) 02-03 M F/A 101 0.70 0.782 1.111 0.234 0.877 0.095 F/A 101 0.90 1.005 1.315 0.180 1.135 0.130 F/A 101 1.00 1.117 1.390 0.210 1.180 0.063 F/A 14 0.70 0.782 1.172 0.234 0.938 0.156 F/A 14 0.90 1.005 1.350 0.180 1.170 0.165 F/A 14 1.00 1.117 1.442 0.210 1.232 0.115 C

a

• NFP = Nominal Full Power l

6001A-661A O O O

. . . . . . - . . _ . _ . _ _ . . . _ . .. _ ..... . . ._ __. _ _ . .. _ .~ ~ . _ . . _ .

'e O O O TABLE XXIX SUP91ARY OF LIFE-3 RUNS FOR INNER BLANKET POWER-TO-MELT ANALYSES NEUTRON ELEVATIONS, IN. POWER HCF FLUX HCF POWER LEVEL POWER-TO-MEL' RUN PIN - (WORST ELEV. UNDERLINED) CLADDING CLADDING AT WORST AT WORST (1.03) *

1. (IB/A 99) (Midplane = 32 in.) TEMP. HCF SWELLING ELEVATION ELEVATION HCF (NFP = 1.0) 1 Hot 18, 24, 32_, 40, 46 00' Nominal Midplane Same as Yes 1.360 Value Power 2 Hot 18, 24, 32, 40, 46 -~

30 Nominal Midplane Same as Yes 1.301 Value Power 3 Hot 18, 24, 32, 40, 46 ~~

Oo Maximum Midplane Same as Yes 1.333 Value Power g 4 Peak 23, 30, 32 Oo Nominal Midpane Same as Yes 1.366 m Value Power 5 Peak 23, 30, 32 3g Nominal Midplane Same a., Yes 1.280 Value Power 6 Peak 28, 30, E Og Nominal Midplane Same a:; No 1.398 Value Power 7 ' Peak 28, 30, 322 Og Maximum Midplane Same as No 1.359 Value Power 8 Peak 30, 32, H , 36, 38 Og No.ni nal Chopped 1.15 Yes 1.338 Cosine

(*)NFP = Nominal Full Power 6001A-661A

TABLE XXX MARGIP-TO-MELT ANALYSIS (REACTOR POWER UNITS - Oo POWER = 1.0)

PEAK PIN HOT PIN WORST ELEVATION MIDPLANE MIDPLANE CASE 1 CASE 2 CASE 1 CASE 2 CASE 1 A. Reduction in power-to-melt because of uncertainties in cladding swelling parameters (30) 0.039 0.027

8. Overall extrapolation uncertainty, a extrap (Item A combined with remaining uncertainties) 0.0377 0.0365 C. Reduction in power-to-melt because of 0.086 0.059 uncertainties in cladding temperature (30)

0. Overall design uncertainty, R (Item C

- combined with remaining uncertainties) 0.0289 0.0200 E. Power-to-melt for base case, REPTRM l 1.338 1.366 >

1.398 1.360 F. EBR-II experimental uncertainties, g , = 0.041

• REPOWM 0.0549 0.0560 0.0573 0.056 G. Trui "REPOWM =

p / ., g,-ran2+oextrap2)1/2 0.0726 0.0734 0.0744 0.070 H.

Marg h do-melt = REPOWM - 3aREPOWM~

Cl *REPOW; where C1 = 1.15/1.03, Case 1 0.004 0.000 0.0 29 0.025 0.033 C1 = 1.15, Case 2 6001A-661A O O O

-O.

O O%.) .' %..

[]N

' TABLE'XXXI

' FUEL ASSEMBLY COMPONENT PRESSURE DROP. DATA LINEAR RECRESSION ANALYSIS STANDARD STANDARD STANDARD' NO. OF 'MEAN' DEVIATION MEAN DEVIATION ERROR DATA 0F ABOUT MEAN- 0F - ABOUT MEAN OF

' COMPONENT LINEAR REGRESSION FUNCTION POINTS In(Re) 0F In(Re) in(D) 0F In(D) EST' MATE' Inlet Nozzle in(K)=0.9177 .05289 in(Re) 222 '13.64 0.3560 0.0841 Inlet Nozzle-Orifice Shield:

-- 1 Plate in(K)=2.352 .092111n(Re)-1.4521n(D) 41 13.95 0.3763 .1845 0.1256 0.0170

-- 2 Plates in(K)=1.708 .050221n(Re)-3.2931n(D) 73 13.73 0.3684 .3528 0.1057 0.0472

-- 3 Plates In(K)=2.240 .082261n(Rei-3.8911n(D) 60 13.61 0.3560 .4064 0.1165 0.0207

,-- 4 P1ates in(K)=2.293 .071411n(Re)-4.0321n(D) 43 13.52 0.2982 .4454 0.1040 0.0136 y -- 5 Plates in(K)=2.225 .030721 n(Re)-3.6511n(D) 42 13.45 0.2589 .4484 0.0997- 0.0165 Shield in(K)=0.3988 .038791n(Re) 17 13.82 0.3768 0.0966 Rod Bundle:

-- Inlet K = 0.370 --

0.2(*).

-- Rod Friction see Table XXXII 161 entire range 0.0524 46 full flow +0.03I2,-0.0262

-- Outlet K = 0.178 --

0.2(*)

Outlet Nozzle in(K)=.00495 .049021n(Re) 16 13.67 0.7483 0.0450

(*) A 20% uncertain'ty was selected as a bounding value (not standard error), since no data are available.

i 6001A-661A

TABLE XXXII FUEL ASSEMBLY COMPONENT HYDRAULIC CORRELATIONS REFERENCE REFERENCE HYDRAULIC COMPONENT CORRELATION 2 AREA (IN ) DIAMETER (IN) REFERENCES Inlet Nozzle K=2.504 Re-0.0529 3.976 2.250 29 Inlet Nozzle-Orifice-Shield:

-- 1 Plate K'10.50 Re -0.0921 D-1.452 3.976 2.250 29

-- 2 Plates -5.519 Re-0.0502 D-3.293

-- 3 Plates K=9.396 Re-0.0823 D-3.891

-- 4 Plates K=9.909 Re-0.0714 D-4.032

-- 5 Plates K=9.253 Re-0.0307 D-3.651 Shield K=1.490 Re-0.0388 3.976 2.250 29 q Rod Bundle:

-- Inlet K=0.370 6.724 0.1281 33

-- Rod Friction f=84/Re for Re 5 1000 6.724 0.1281 26-28 f= (1.080+0.0927 (1000/Re)2+0.1694*

(1000/Re) ) f c where fc(*}= 4 Log 10(2.51/(Re Q))

-- Outlet K=0.178 6.724 0.1281 33 Outlet Nozzle K=1.005 Re-0.0490 5.899 2.116 29

(*) f is c the Colebrook friction factor correlation (34)for a smooth tube.

6001A-661A O O @

4 TABLE XXXIII' INNER BLANKET. ASSEMBLY COMPONENT HYDRAULIC CORRELATIONS ,

REFERENCE R EFERENCE HYDRAULIC COMPONENT CORRELATION -AREA (IN ) 2 REFERENCES'

' DIAMETER (IN)

. Inlet Nozzle K=2.504 Re-0.0529 3.976 2.250 29 Inlet Nozzle Orifice Shield K=C Re

-0.05 3.976 2.250 29,30 Shield K=2.0 2.405 1.750 33 Rod Bundle:

-- Inlet' K=0.4 27 3.956 0.1338 33

-- Rod Friction f=110/Re for Re < 400 3.956 0.1338 31 0

f=(110/Re) @ + (0.55/Re .25) 4 I $ where 4 = (Re-400)/4600 for 400 < Re < 5000 0

f=0.55/Re .25for Re > 5000

! -- Outlet K=0.290 3.956 0.1338 33 Outlet Nozzle K=1.005 Re-0.0490 3.976 2.250 29

}

TABLE XXXIV RADIAL BLANKET ASSEMBLY COMPONENT HYDRAULIC CORRELATIONS REFERENCE REFERENCE HYDRAULIC COMPONENT CORRELATION 2 AREA (IN ) DIAMETER (IN) REFERENCES Inlet Nozzle K=2.504 Re-0.0529 1.767 1.500 29 Inlet Region K=C Re-0.05 1.767 1.500 29,30 Shield K=2.0 2.40s 1.750 33 Rod Bundle:

-- Inlet K=0.427 3.956 0.1338 33

-- Rod Friction f=110/Re for Re < 400 3.956 0.1338 31 f=(110/Re) /1-$ + (0.55/Re0 .25)g y for 400 < Re < 5000 where $ = (Re-400)/4600 0

f=0.55/Re .25 for Re > 5000

-- Outlet K=0.290 3.956 0.1338 33 Outlet Nozzle K=1.005 Re-0.0490 3.976 2.250 29 6001A-661A e 9 9

_. . . . . .- ._ . _ , _ _ m. , _ _ . , - ._ . , . ____ _., , . ._ _ , s. - ,

(~

(:f L)q v TABLE XXXb DETAILED PRESSURE DROP BREAKOOWN FOR PLANT THOV CONDITIONS (41.446 x.106 lb/hr) AND MAXIWM UNCERTAINTIES Orificing Zone 1 2 3. 4 5- 6 F/A 6 IS/A 7: '8 9 IOC. 10P 11 '12-Average Zone Flow 187131 174126 164395 151088 147235 177270' 69072 87447 76853 61360 47569 47569 34563 25356 Component AP (psi)'

. -Inlet Nozzle - 8.7 7.6 6.8 5.8 5.5 7.9 1.3 '2.0 1.6 4.9 3.0 .3.0 1.5- 0.9 Assembly. Orifice f40.0 48.6 55.3 63.5 66.9 46.5 69.9 44.9 60.5 -69.6 42.2 42.2 22.5 -12.2 Shield 6.4 5.5 5.0 4.2 4.0 5.7 5.6 8.9- 6.9 4.4 ' 2.6 c 2.6 : 1.4 0.7 R'od Bundle Inlet 1.1 1.0 0.9 0.7 0.7 1.0 3.5 0.8 0.6 0.4 0.2 .0.2. 0.1 0.1 Rod Bundle '51.5 45.3 40.8 35.1 33.6 46.7 35.9 54.3 43.3 29.3 18.7 '18.7 10.7 6.2 Rod Bundle Outlet - 0.6 0.5 0.4 0.4 0.3 0.5 0.4 0.6 0.4 0.3- 0.2 0.2 0.1 0.1 Outlet Nozzle 1.6 1.4 1.3 1.1 1.0 1.5 0.5 0.8 0.6 0.4 0.2 0.2 0.1 - 0.

Tota 1' Assembly .109.9 109.9 110.5 110.8 112.0 109.8 114.1 .112.3 113.9 109.3 67.1 67.1 36.5 20.3 LIM, Upper Portion - .

and Orifice 2.9 '2.5 2.3 1.9 1.8 2.6 - 3.4 0.6 0.5 5.7 .47.3 48.8 79.6 95.5 LIM, Lower Portion 5.6 6.0 5.6 5.7 4.6 6.0 4.8 5.5 4.0 3.4 4.0 . 2.5 ' 2.3 2.6 Inlet Plenum 5.8 4

y Outlet Plenum 2.2 Vessel Nozzie-to-Nozzle 126.4 Primary Loop 37.1 Pump Head 163.5 NOTES:

-- All flows for cycle 4, except zone 6 IB/A which is for cycle 2;

-- All flow resistance uncertainties at +3a or maximum bounding value, whichever appropriate;

-- Minimum pump head curve; I

-- Zone I through 8 and 10C are fed by central modules, zones 9,10P,11 and 12 by peripheral modules.

6001A-661A i.

TABLE XXXVI DETAILEO PRESSURE OROP CORRELATIONS FOR UPDATED PLANT EXPECTE0 OPERATING CONDITIONS (46.420 x 106 lb/hr)

Orificing Zone 1 2 3 4 5 6 F/A 6 18/A 7 8 9 IOC 10P 11 12 Average Ione Flow 210118 195515 184588 169646 165320 199045 77613 98189 86292 68897 53317 53715 39909 29518 Cumponent aP (psi)

Inlet Nozzle 10.4 9.1 8.1 6.9 6.5 9.4 1.5 2.4 1.8 5.9 3.6 3.6 2.0 1.1 Assembly Orifice 17.5 27.8 35.4 44.8 48.5 24.2 52.5 24.6 42.1 52.3 31.6 32.1 17.9 9.9 Shield 7.6 6.6 5.9 5.0 4.8 6.8 6.7 10.7 8.3 5.3 3.2 3.2 1.8 1.0 Rod Bundle Inlet 1.? O.9 0.8 0.7 1.0 0.5 0.8 0.6 0.4 0.3 0.3 0.1 0.1 Rod Bundle 52 .4 t. ' '

46.4 39.9 38.1 53.1 40.0 60.4 48.2 32.6 20.8 21.1 12.5 7.4 Rod Bundle Outlet 0.6 0. 0.5 0.4 0.4 0.5 0.4 0.6 0.5 0.3 0.2 0.2 0.1 0.1 Outlet Nozzle 1.8 1.6 1.4 1.2 1.1 1.6 0.6 0.9 0.7 0.4 0.3 0.3 0.1 0.1 Total Assembly 97.5 98.0 ;3.6 99.0 100.1 97.6 102.2 100.4 102.2 97.2 60.0 60.8 34.5 19.7 LIM, Upper Portion and Orifice 3.1 2.7 2.4 2.0 1.9 2.8 0.4 0.7 0.5 6.0 43.0 43.6 69.9 84.4 LIM, Lower Portion 6.3 6.2 5.9 5.9 4.9 6.5 5.0 5.8 4.2 3.7 3.9 2.5 2.5 2.8 Inlet Pier.ua 6.0 Outlet Plenum 2.4 k Vessel Nozzle-to-hozzle 115.3 Primary Loop 41.8 Pump Head 157.1 NOTES:

-- All flows for cycle 4, er:ept zone 6 IB/A which is for cycle 2;

-- All ficw resistances are rominal;

-- Nominal pump head curve;

-- Zone 1 through 8 and 10C s e fed by central modules, zones 9,10P,11 and 12 by peripheral md sles.

6001A-661A O O O

t INNER BLANKET: HEAVY LINED ASSEMBLIES

\ / 313 \ RADIAL BLANKET: 200 SERIES (300 SERIES

( 308 SYMMETRICAL) 302 306 311 301 305 310 )

/ / \ / \ .

( /\ / \ " / ' ' X " / '" \ I

(,,\\"\ / ,,

/ ,,

,, \

,,\

/ 20, \

/ 20

/,,,\

\ /,,,\

\ ,/ \ /

102 47 27 206 209 (C/A 37 101 s 12

/ 26

\ / 207

\ 1 l

• 98 36

/ 68 13

\ / 201

\ / 208 67 / 25 \ / 202

\

32 99 100 (C/A 24 O

11 14 33

/ 60 10

/ 15

\ /

C/A 34 61 C/A 94 3

• 62 30 59 29 128

' FUEL / INNER BLANKET ALTERNATING POSITIONS 57 69 Figure I CRHRP lieterogeneous Core (Rev. 4) 60" Symmetry Sector and Assemblies Numbering Scheme 166848 l

4 133

t 101 9 -

8 -

y -

• 6 5

4 F/A 15 F/A 25 3

F/A 13 F/A 24 2

Si u

Y 5

100 _

g -

$ 8 6 7 s.

t; 6 - 1 I

Q l l 4 -

ll l l ll lDELT l 3 -

ll l l1 l

II l ll 2

ll,i  ;

II;I ll l I I l I I 10-1 I 1200 1220 1240 1260 1280 1300 1320 1340 EOL CLADDING 10 TEMPERATURE (OF) i 17igure 2 Typical 1)ELT Determination for Iirst Core Iruel Assemblies 166847 O

134 l

l

4 100 8  :

I 6 cap tig,y 4 -

2 -

IB/A 46 l

10-1 _

j 8 1 6 - IB/A 57

4 -

2 1 IB/A 32 10-2 _

u. 8 o -

u 6 -

i w -

) Q 4 i M -

l $

2 -

w

- 5 10-3 _

1 4 -

2 10-4 _

, 8 6 -

l 4 -

g gg ELT g gg 2

10-5 I I I I i 9

'ni 105G 1100 1150 1200 1250 1300 1350 1400 EOL CLAD 0 LNG ID TEMPERATURE (DF) 1 Figure 3 Typical DELT Determination for First Core Inner Blanket Assemblies 1668-66 1

135 l - --- - - - - - -

1 10 9 -

8 7

6 5

4 -

F/A 48 3

F/A 51 2 -

i o

Y 100 _

>- g -

0 -

CDF LIMIT 6 -

l 5 - II Il 4 -

ll l l DELT 3 -

l ll Ill lI l

2 l! l 1I II II j l 1 I I f111 1 1110 1130 1150 1170 1190 1210 1230 1250 EOL CLADDING ID TEMPERATURE (OF)

Figure 4 Typical DFLT Determination for Second Core Fuel Aweml> lies 1668-65 136

i 6 - CDF LIMIT /

I' 4 -

~

IB/A 67 2

10-1 _.

8  :

, 6  :

4 -

2 -

IB/A 93 10-2 _

g 8  :

o 6 2

1. .c 4 -

4

$ - IB/A 69

$ 2 -

5 5 10-3 _

8  :

/ 6

}

4 -

2

! 10 4 -

8  :

6 I

1. ~

O r. 3 DELT 5 55 5 2 - E 22 2 jg-5 I I I I I "l 4 1050 1100 1150 1200 1250 1300 1350 1400 EOL CLADDING 10 TEMPERATURE (OF) j Figure 5 Typical DELT Determination for Second Core inner Illanket Ass mblies 1668-64 O .

137 i

(

..,..,,-,--.c ,, . , , -

l i

)

I O

RB/A 201 0

10 -

9 -

8

, 7 CDF LIMIT g 5

5 m 4 -

$ l

, 5 l g 3 -

lDELT 2

l l

l

' ' ' I ' '

10-1 1120 1140 1160 1180 1200 1220 1240 1260 EOL CLADDING ID TEMPERATURE (OF)

Figure 6 Typical DELT Determination for Radial Blanket Assemblics 1668-63 O

138 l

l

v y ASSEMBLY NUMBER 308 as

\ /

[ 313 \ nana - TELI 307 312 mana OELT 302 21 306 x 7 3,, x 7 2., x 7 3,. x

,,,'\ / "\ / '"\ / '"\

"" / ,,,,"\ "" / ,," \ / '"\ / \

,,,:'\"" / ,,0\ "" / ,,,,"\ / '"\ /

44 1246 46 1259 28 20 21 1225 / /

/ 1242 1229 1227 \ / \

M 102 1318 47 1259

[ 27\ [ 206 [ 209

( 1232 1237 / 1239 \ / i 37 13gg 101 1235 12 1262 26 207

,,; " / ,,56 "

\ "" / ,,',' \" / "'\ / '"

,, 'l = / c,, \\ "" / ,,4,"\ = /,,,"\ / '"\

o ,,,:'

1341 33

,l:\ / u.."\ "" / ,,;; \ "" / ,,5"\

1335 ""' 1260

/

1300 60 1230 1307 10 1250[1290 i

15 J

f 1256 1253 34 1297 1343 1244 61 1256 I

[

C/A (C/A )M g4 3 1358 62 1299 1309 30 134! Sg 1264 4 131, 1306 1222 1236 '

LIMITING CONSTRAINT UNDERLINED 1265 29 1344 128 1364 TEMPERATURES IN 'F 1308 1219 93 1269 2 1357 l'I' 1308 1310 57 1219 1269

/

/

,,,:'\"" /

1320 Figure 7 Imthittual Assemblics Limitir meratures at First Core 166M-62 0

139

O r ASSEMBLY NUMBER oll'\ / ,l." \  ::iih::

1320

/ 307 \ \1336/ / 312\

/ 1250 1243 \

,,,a I

j SELT (ONLY FOR FUEL ASSEM8Lt!Si

> >\"a / ,'" \ ""/ m;"\

/,,,',"\ "" / ,,'"\"" /,,4"\

52 1222 1293 210 1111 50 \ 1260/ [1266304 1188 \ 1246 1202 49 1293 1270 1211 1215/ / 1181, \\1260 1281 1267 0 02 43

" 45

' 48 k 1 46. / 2041301 \g[ 212 1232 1210 1185 \ / 1266 \ / 1243 1208 1199 1257 46 1206 / 28 \ 1260 / 205 \ I3JJ f 21 \

1240 1256 12?2 / 1186 \ / 1264 \ / 1240 \

\'!:: ,;;; g / ,,9, ' :::: ,,8 '\ '"/ ,lll\ "" / J

,,,," ,,;,'\;;;' ,,;>  : ;; / ,,8!\ "" / ,,'"\ ""

" e ,,:: = / ,,,," " ,,,;'\ :::: / ,,','O'\ " / ,lll

'!' ,,,,"  :::: \+;n ,,ll  :::: / ,,,'!\ "" / ,,:l'\ ""

" * ,,.;'\ n/ ,,9!\ "" / '

,,48

": ,,,l>

• ,,::  :::: , ,i,'  !!n / ,,1t\ :::/ O 1219 34 1334 ' /

i{fg 61 1250 C/A

' 1254/

94 1217 3 1240 1258 1295 1336 62 1306

"[ '

30 1342 59 1221 4 1315 1299 1739 1260 1251 29 LIMITING CONSTRAINT UNDERLINED 1346 128 1342 TEMPERATURES IN 'F 1239 1229 2 93 1348 g,g 1246 1303 1347 57 1229_

1246 1350 1111 1374 Figure 8 Indi.idual Awemblies Limiting Temperatures at Second Core 1668 59 9

140

(V\

r ASSEM8LY NUMBER

/ taa\ / 313 \ ^^

-'\

23186

/

21574

\ as.. r RST CORE 3 mann SECOND CORE 30, x - 7 306 x > = > f >;; x 7 30,x>->f 305 x, / ; x 5,x 7,,,,0 x f - x,, ,/ ~ x a ,

5, y ,.8,,, 7 153972 x 47 170738 49 7 203 x > = >31422 f 211 x 1723 156707 2

4 ,,0,,0 46

7 .55599 7 28 x 4> ,, f 205 x 26 / mx

\ ;;;;; .0, ':::: 4, '::::: ,, -

15929 181810 180321 153273 x- f 36553 x 2""f 142 83 3, 90150 101 12 152447 26 207 193005 70498 149765

' 34286 gg 36 191632 68 79260 148890 201 6 13 208 69163 189053 196027 180677 64213 23186 169218 gg 189802 190893 67 170337 25 202 3

g 44 100 11 '3 30 74654 196745 197499 159744

':" ~ >8 'a>> ' '

3>

e0

,,,,;, ' a> f,,,6 :\"'*/

C/A

,804,,

94

,4 176053 181573 9,,,,

3 6,

82069 92037

,84,99 62 CIA xm-/

70664 176103 188241 30 88663 59 179073 4 169519 166014 I 175426 29 87256 128 W >-

164007 7 27 2 80226 ~ FLOWS AT PL ANT EXPECTED OPERATING CON 0lil0NS 93 173518 --FIRST CORE FLOWS AT CYCLE 2

'65049 165314 --SECOND CORE FLOWS AT CYCLE 4 (EXCEPT SECOND 80613 57 173786 ROW RADIAL BLANKET AT CYCLE 5) 65082 ONSTRAINT UNDERUNED 80492

~~

$' 8 69

$$127 I.' . '

l'igure 9 Initisiilual Assemlilies \lininutm I-low Rates at Iirst anil Seconil ( ore Nemsary to Satisfy liie Constraints 166& 60 m

141

O ASSEMBLY NUMBER ZONE NUMBER l

11

\ / 3"\ \ /

306 311 II 301 305 310 58 304 210

' 10 51 49 211

, 203

' 10 212 43 45 48 204 j F2 F 3 4 10 i2 g ,, 205 213 2 S 8 4 ' 206 200 i,2 47 27

• 12 C/A 7 2 5 2s 11 207 37 u S 1 F 8 y f 6 ,, 13 201 2 I 1 67 2 l Sg 25 202 S7 3,,

@ F 7 1, 2

24 2 ,

33

' I I 1 33 60 10 15 7 C/A 38 59 4 128

• = ZONE ORIVER 33 F = LIMITING ASSEMBLY IN ZONE FOR FIRST CORE g g 3

. S = LIMITING ASSEM8LY IN ZONE FOR 8 SECOND CORE g

8 I ,

t l Figure 10 Core Orificing Scheme 166 & 61  ;

142

ASSEMBLY NUMBER

[ 308 313

$ musa AVERAGE L'j O.8 06 mann PE A K 1.1 1.1 1.5 mana 3 5 IGM A 1.6 307 312 1.7 1.0 1.7 08 muss 3 SIGMA e OVERPOWER 18 1.5 311 302 26 306 2.2 1.9 17 2.4 0.9 2.9 4.3 39 1.7 5.4 Jul 5.0 305 16 2.8 310 '

R.0 3.2 5.6 2.3 10 6.4 5.5 18 52 8.1 'O 304 2.6 210 50 7.4 91 6.8 7.8 2.4 2.8 1.0 10.9 56 1g 10.4 12.6 51 12 0 49 7.2 203 2.6 211 14.1 8.5 13.4 7.2 8.0 2.5 2.8 0'9 11 8 10.4 5.8 37 13 2 45 12.0 43 7.3 204 25 43 13 4 8.2 1

9.0 14 7 8.7 7.3 2.4 258 0 12.9 11.8 to 6 56 1.5 44 12.3 7.2 14 0 15 6

/ 9.

13.2 14.7 ,46 13.7 80 213 f

C/A

\:" ':

9.3

s 12.2

!U ,s 10.4 3

e39 U =

101 12 207

,u 37 11.g 1- ,u ,13.9 u ,12.0

- / ::

10.4 n a n' 14 1 68 \ 11.6 201 2.6 208 98 13.9 36 g3 15 8 95 13 0 32 2.9 06 4.3 15.5 9.2 8.7 11 12.0 12.3 11.7 64 y3 a

32 7.6 98

" / \ :n /:2: X:n 100 cia 9

g g \ 11.9 / g

/ e. 11 7

\ 13.1 n

24

>n 4.3 5.4 [

3. 7 10.9 4.7 g,g \ 13.3 / g4g 14 6 7.4 10 8 6.0 /

5.7 7.5 12.0 12.7 9.7 10 15 12.5 7.3 33 60 13.7 13.8 14 0 8.2 7.9 10 8 46 15.3 8.7 1$ 4 8.9 10.3 7.5 12.6 12.6 0 11 P 34 9.7 13 8 13 9 61 13 8.0 10 8 15.4 15.6 4.3 CIA f 94 1 0 3 gf 62 4.3 1

I34 77 10 9 4.4 I

'O I 68 11.8 9'b 7

30 87 59 4 7.5 42 13.2 3.3 10 8 g8 97 66 52 11.2 29 85 128 66 l 12.5 7.3 95 3.8 7.4 POWER RATINGS IN KWiFT I

9.5 5.4 11.0 2 93 7.0

'~ 73 3.5 7.8 6.3 94 gy 57 10.9 gg 3.5 12.1 6.2 69 I8 24 88 3.1 40 4.5 Figure 11 Core Assemblies Linear Power Ratings at HOCl 1668-52 143

. . - . , w-

i i

I i

1 O

AS$f MHL Y NUMHf H 308 313 ,, T 09 08 aman g AVfMAGE 16 15 a=== j PE AM 22 y* 21 312 25 1.4 24 24 11 2I k\ *m*a*n="

I f 3 SH MA 3 51 'MA

. OV1 HPOW6 H 302 37 306 30 3gg

2. 4 41 21 34 1.3 53 48 25 67 301 61 305 35 310 7.5 39 68 28 39 13 78 .

66 26 52 98 50 83 304 36 210 69 10 9 6.4 92 3.0 41 13 10 0 95 68 26 11 6 51 11 0 8.5 203 36 111 49 12 8 79 12 3 67 95 31 41 13 10.8 9. 5 69 25 12 3 45 gi c 4g 87 204 3b 212 43 39 85 13 7 81 12.3 68 97 30 1.1 11 4 to 8 97 68 21 12 2 12.9 / 44 87 137 46 11.2 12.5 28 68 SS 95 205 28 30 34 213 08 7.1 14.4 / 11 5 11.3 *7 66 i5 13 0 13 7 47 11 2 27 83 n y3 ?M 102 14 5 77 15 3 83 12 5 67 92 2i y4 oe 11.1 95 48 10 C/A 11 9 14 37 14 5 12.7 gy 10 9 26 61 207 tot 16 86 16 2 14 2 1 I 63 68 14 88 70 26 11 5 11.3 11.3 95 12 8 12 g 68 13 le 0 201 37 208 98 36 33 7 41 14 3 14 5 88 15 3 81 12 2 39 09 i 8.2 87 18 11 2 11 3 10 8 16 10 4 98 99 12 8 12 9 67 12 2 25 202 22 12 7 10 9 ya 3g 14.1 63 14 3 14 4 77 13 6 79 to 5 11 8 10.8 53 C/A 12.2 67 32 12 8 100 11 14.4 14 34 14 3 16 1 8,7 13 6 69 75 56 66 87 84 10 4 *g, 11 5 10 0 10.2 33 12 7 60 'O 12 8 gg 11 g 12 8 85 14 3 8g 17 g 18.4 78 14 1 66 g, y 99 10.4 11 4 11 3 11 4 34 12 6 61 II 8 12 9 12 8 78 14 1 62 '42 14 4 C/A 10 0 10 4 CA 94 *15 3 12 7 62 6.1 *28 76 I41 62 93 98 10 3 30 11.3 gg 11 4 4 12 6 7.4 12 6 gg 12.7 54 i40 95 91 78 10.9 gg 1y i yyg 95 12.2 73 12 4 gg 10 6 93 93 10 7 y ,7 yg 5.1 11 9 7.3 10 8 1 85 92 10 4 57 10 6 11.6 50 11.9 84 69 10 3 39 11 5 POWE R R A flNGS iN MW Fi 54 66 F.3 Figure 12 Core Assemblies Linear Power Ratings at EOCl 1681 35 144 0

ASSE MBLY NUMBE R 30g 313 ,,

'  ! """" AVERAGE f6 22 """" 3 PEAK 2.3 307 312 /

1.4 24 """" 3 SIGMA 26 1.1 f 27 2.2 f 3 SIGMA OV E RPOWE R 3.7 306 3.0 39, 302 4.2 22 34 2.5 g3

5. 4 49 25 68 61 305 35 310 301 76 4.0 69 35 3g y3 79 82 yg 5 gg 50 to 4 304 36 210 6.3 i t ,1 5.9 11 6 30 40 13 91 88 68 25 10 5 51 10.2 49 85 20.1 3.6 211 11.7 7.2 11 4 63 95 30 40 1.3 98 90 68 2.5 43 11 1 [ 45 10 4 48 85 204 35 212 75 12 4 7.5 11 6 66 96 30 3.9 11 10 3 / 10 0 95 68 2.2 44 11.3 ep 10 9 28 8.5 205 JO 213 11 6 13 7.8 12.6 6g 12 2 67 95 35 34 09 10.7 11 1 96 82 gg 11 1 10.4 206 22 209 11 7 ggy 135 47 y, 6

13 1 7,3 15.1 81 12 4 6. 7 22 24 0.6 11 0 9.6 49 10 CIA 11 4 "

12 6 11 1 26 61 207 1.4 37 13.9 101

7. 7 14 1 12.3 64 69 1.4 16 15 5 8. 7 U4 11 0 96 2.7 11 2 12 0 12 9 68 13 11.1 201 37 208 gg 36 11 5 12.4 42 0g gy 13.4 g4 14 4 gg 82 40 11 4 10.9 79 1.6 92 11.0 99 99 12.6 13 0 1 -

5 202 23 11,3 67 11.1 12 6 60 14 1 14 5 79 13 8 79 25 26 10 0 cia gy 4 10 9 54

\ 12.2 100 1g gg y 14 12.4 24 68 13 6 66 g7 1 8 0 pg 55 89 16 9 10 1 10 4 11.4 11.9

'7 1' 33 32-14 2 60 13.1 10 88 13 3 [ 15 8.5 13 0 g 7.8 68 14.7 14 9 f 10 0 10 y 12 2 11.6 11 6 13.3 13 0 34 13.0 61 12 9 14 6 14 8 14 6 80 7.1 CIA 10.3 11.7 CIA 94 11 9 I33 3

14 3 62  !

6.2 gi 16 0 g1 /

95 10.4 12 4 30 7.6 11 5 12 9 59 12.0 63 4 14 0 [

6.1 13 4 15 6 /

9. 7 g5 g4 11 2 29 11 6 128 11 5 12 6 75 13 0 5.9 12 8 96 84 g3 11.1 2 10 2 52 12 4 76 11 4 POWER RATINGS IN KW Fi ,

88 87 l 10 7 67 Il I gg 9 5.2 12 4 88 10 7 69 g l .g i 4.1 1 56 68 f 76 l

1 1

Figure 13 Core Assemblies Linear Power Ratings at BOC2 1681-36

(

145 l

l

O AS$1 MBL Y NLIMHE H 308 313 ,,

1.2 1.2 2.3 2.1 anna AVERAGE 3.2 307 29 asaa PE AM 3.5 1.9 3.3 3, anna f 3 5HiMA

• • • • f 3 Sit.M4 OVi HPOWf H 29 305 306 4.1 311 3' 7 28 1.7 3.2 45 6.7 6.0 3.3 8.3 301 7.4 305 46 310 9.3 5.0 B.3 3.5 5. 2 1.8 9.6 7g 35 11.9 50 9.6 304 4.8 210

$2 18 59 13.3 5.4 10.7 3. 7 53 7g 8.0 3.5 8.2 48 95 51 90 49 9.9 203 211 10.6 6. 7 10.1 5. 7 11 0 3. 7 53 17 8.9 79 80 3.3 43 10.2 45 91 48 99 46 212 11.4 6.9 10.2 59 11.0 2".

3 7 5.2 15 7.3 80 2g 9.1 89 8.1 10.2 94 99 205 41 10 6 / 44 46 10.5 28 11 0 35 45 '

11.4

\ 11.8f 7.5 9.2 34 13 0 8#

60 82 78 y

y, 10.7 102 14 5 95 27 9.6 206 29 209 9.4 16.1 7 10 6 60 10.7 28 33 08 12.0 14 C/A 14 0 83 60 79 8 95 26 74 207 19 37 15 6 101 12 97'9 10 6 58 19 22 7.6 17.4 77 88 83 94 96 13 4 84 3.7 10.8 36 11.1 68 le g 13 97 201 5.1 2t*

gg to 8 g4 12.1 7.7 12 4 79 16 6 73 50 57 12 12.9 9.5 96 95 96 23 gg 11 0 11 3 67 11 0 25 11 9 202 32 14.3 12 3 13 3 12.3 12 6 10.2 7.2 32 35 16.0 86 13.2 C/A 15.4 9. 7 6. 7 32 14.7 100 11 17 2 14 11.1 24 83 8.1 18.4 92 8.1 19 2 82 12 4 64 P3

( 12.0 13.5 'O.1 10.2 g.g l

114 33 15.0 60 11 6 to 11.9 15 to 3

[

15.0 16.8 95 13 0 8. 4 13 3 80 11 5 7.4 9.1 13 9 10.3 10 0 10 5 34 15 5 61 Il 8 11 1 76 17 3 10 0 13 3 13.1 11 8 C/A 93 15 3 CA

( 94 8.7 10 8 12 0 3

7,7 17 1 19 1 62 gg

[

l 12 9 94 -

10 8 f

l

' 30 7.4 14 4 16.1 59 gy 10.9 12,1 93 4 12 6 [

90 13.7 14 0 /

13.1 10.3 29 14 6 128 15 2 11.5 7.3 16.3 86 17 0 89 12 3 10 3 2 13 7 g3 7, y 11 5 7.4 15.3 2 POWE R HATINGS IN kW. F T 57 1O 7.7 11 6 15 2 12.2 69 13 6 66 15 2 9.2 10.2 11.4 Figure 14 Core Assemblies Linear Power Itatings at EOC2 1681-37 146

l

\

ASSI MBL Y NUMBE R Jag 313 ""

13 13 2.5 24 aman g AVERAGE 35 307 33 312 ====

f PEAK 39 21 36 16 ====

f 3 SIGM A 41 33 ==== f 3-51GMA + OVE RPOWE R 302 56 306 4S 311 35 6.3 30 61 1.9 i4 67 38 9I 301 83 305 310 10 2 93 39 5.,3 5 yo 3.4 10 5 89 40 52 50 11 0 304 55 210 13 0 i 74 is 5 69 12 3 42 6. 2 2.0 1 10 5 10 3 93 40 J 12 2 53 11 9 49 11 4 203 55 211 1 13 6 83 13 2 73 12 8 43 6.2 19 <

11.2 10 3 95 38 1 43 12 5 45 11 9 48 11 7 204 63 212 l 85 14 0 86 13.3 74 13 0 4.2 5.9 16 i 12 0 g 3.3 10 5 93 3.3 13 1 44 12 6 46 12 I 11 4 205 45 213 13 5 28 5.1 l

14 7 85 14.1 56 74 12 8 3.9 1.3 12 3 88 10 5 89 24 12 9 11 3 12.1 27 11 0 206 33 209 14 4 102 88 12 3 36 57 12 6 13 5 7.2 30 0.8 C 88 to 3 6.7 1.5 1 33 4 37 is 3 101 15 0 12 11 9 26 83 207 21 l 81 12 6 91 56 13 2 6.9 93 2.1 2.4 )

11 8 11 8 87 10 3 41 98 12 4 36 13.5 68 13.3 13 11 8 201 56 208 3g 13 8 87 15.1 91 12 6 85 13 2 5.4 6.3 1.3 6y 11 4 11 8 11 2 10.5 2.5 13 1 13 5 67 12 6 25 13.0 202 35 86 99 96 40 14 6 15 1 5.7 14 0 83 14.5 35 3g CA 8.7 11.1 74 7.0 100 1t 11 2 14 12 5 24 91 32 90 13 9 10.0 44 8.6 12.5 85 74 10 2 35 53 70 II 4 12 3 10 5 69 90 60 13 0 to 12 8 15 12.1 3 10 1 14 5 80 14 3 84 13 5

7. 7 I4 44 yo 11.8 12.0

,96 34 90 61 12 3 \ 13,1

/

75 13 8 14 6 C; A jy 4 97 10.1 40 69 C/A /

II 2 89 94 3 6 12 5 go n 3 ,2 42 73 64 95 67 30 82 11 0 4 8.6 59 71 92 40 12 3 33 96 91 62 48 10 5 29 go 128 62 11 7 69 gg 36 69 89 51 10 3 2 6G 11 5 G9 7.3 POWER R ATINGS IN KW F T 3

59 88 77 57 10 2 85 33 11 4 58 69 75 22 8.3 30 38 e2 Figure 15. Core Assemblies un ,r Power Ratings at BOC3 1

i 1681-38 1

147

i A%E M8L Y NUMBEH O

)

308 313 ** I 1.8 1.7 31 nas AVERAGF 3.3 4.5 307 42 47 312

} PE AM 3 S:GMA 50 27 53 21

4. 3

==== ,f 39GMA

• Ovf Rt*Ows H 3ng 7.2 306 5. 7 311

' 8.0 38 6.4 25

4. 3 88 80 49 301 96 305 66 310 10 6 27 64 10 7 47 y4 11.8 5.2 12.2 10.1 1

52 14 6 50 12 2 304 70 210 4

6.4 60 13.6 5.1 7.8 27 16.4 9.0 86 10 6 52 1 12 8 203 70 lli 10.4 51 10 0 49 j 11.1 6.2 14 2 5.2 78 25 Il 6 7.2 49 9.6 86 10 8 99 13 0 66 212 I 43 10 9 48 48 204 #0 74 11 1 6.4 14 5 51 71 7.7 12 2 96 8. 7 10 6 43 99 12 8 5. 7 213 11.3 44 to 9 46 10.1 28 205 7g 12

• 82 11 2 6.3 14.2 4.7 64 iy 12 6 10.1 12.7 g7 TO 1 31 4

11.4 15.3 47 10.1 27 12 2 206 42 209 )

102 38 47 1.1 12 7 17 1 75 11 2 62 13 6 4

90 98 80 20 CIA 93 4 86 37 II 3 99 26 96 207 28 16.1 101 78 12 6 l g ,2 11 1 60 10 7 27 31 17 9 80 10.1 86 5.3

10 0 12 7 11.3 36 11.5 68 13 to 0 201 72 70s e 98 95.3 11 1 6.4 80 i8 7.5 12 7 7.9 12.8 30 97 0 7.4 99 10 0 96 12 2 33 11.8 14 6 202 11 4 11.5 10 9 25 45

14 2 99 67 16 4 43 t, 0 15.9 7.y 12.7 12 8 gg 12 1 7.2 4

C/A 96 8.8 12.1 13 4 10 6 14 6 100 11 16 0 14 10.9 24 32 17 g 7.9 12 2 64 11 8 7.0 16.3 80 79 12 2 99 10.0 g.0

10 4 11 4 11.4 15 10 4 12.4 33 14 6 60 to 16.3 80 12 7 y8 12 7 77 11.6 13.g 7.3 9.1 12 1 10 1 99 14 6 11.2 11.3 10.5 34 61

1. 3 12 5 12 6 11 7 73 76 l C/A 91 12.1 C, A 94 10 5 3 14.5 62

' 73 11.7 7.2 16 2 7.5 11.1 9.0 11 8 30 13.3 59 to 4 4 14.1 l

7.1 14 8 7.2 11 6 68 15 8 88 10 9 g.8 l

, 10.1 29 13 0 129 11 8 l 11.3 69 14 6 69 13 2 8.7 99 l 11.8 l

g3 10 0 2 I 11 1 70 13 2 63 86 POWL R R ATINGS IN MW F T l 10.2

,73 g7 10 0 Il I

. 3. 7 6.3 10.2 69 12 2 53 136 7.3 8.7 98 Figure 16 Core Assemblies Linear Power Ratings at EOC3 1681-39 148

A%$f MH1 Y NUMHf H 308 313 ,, 1 17 16 \

31 29 maan AV t. H ACE 42 30; 40 312 ====

f PE AK 47 2g 44 20 naus j 3 SIGMA 50 40 sans f 15tGMA OV t. RPOWI H 302 6. 7 S4 3II 306 61 40 75 36 24 82 75 47 99 301 91 305 64 310 11 0 60 10.1 4g ?1 2G 11 4 g7 50 13.7 50 11 6 3% 67 210 52 15 3 5. 7 13 0 .9 75 26 60 50 84 82 10.2 95 49 12 3 203 67 211 97 51 10 8 6y 10 6 60 11 7 5.0 7.5 24 gg 83 10 5 47

6. 4 43 10 1 / 45 70 96 10 7 48 6.1 12 6 14 0 204 71 212 49 20 69 96 11 3 / 91 86 10.2 40 10.3 '8 12 3 54 213 10 8 [ 44 Il 5 46 11 0 28 13.7 205 45 61 16 72 12 If 10 0 79 6.1 O? 38 12 8 8r i M

11 2 joy 87 gg 27 M '6 206 40 is 4 1.0 12 5 86 17 2 #3 11 0 60 13 0 36 44 C.A 97 83 75 I' 33 g 37 16 2 101

' 12 96 26 91 207 26 74 12 5 y9 10.7 5.6 10.1 2.5 29 18 1 78 10 5 99 12 8 8.2 50 11 6 36 11 4 68 15.4 13 95 6. 7 208 93 I 9 201 79 12.7 78 17 2 7.0 10 6 60 7.5 1.7 82 91 11 3 99 99 11 4 3.1 12 5 11 4 11.4 67 10.3 25 13.7 202 42 99 ILS 6. 7 14 0 82 12.'s 12 7 86 15 3 40 47 13 0 C/A 13 5 89 82 16 2 14 10 1 24 9.9 32 15 7 100 11 y3 i7g g.2 7g 18.1 71 II 3 60 11.0 11 2 12 6 99 la0 g4 13 5 15 1 60 gg 4 to 11 2 15 g7 33 15.1 16 9 81 12.7 7.4 12 5 6.9 10 8 l

76 j qg 12 6 10 5 96 34 69 l 1 7 8

• i CIA 13 0 C, A 96 94 11 0 3 15 6 62 j

?.7 12 3 7.5 17 4 81 11 6 95 1L2 l 30 13 9 59 10 9 4 12 4 74 15 6 ?S  ?? ? 75 13 9 l 93 11 4 11 1 10 7 13 7 I33 j 29 128 12 0 y3 15 3 14 8 1 77 92 10 4 I 93 10 6 2 12 5 66 11 9 7.3 13 9 10 8 92 POWE R R ATINGS IN K W F1 12 9 57 10 6 14 4 66 11 3 10 7 69 II 9 g r, 14 4 )

76 91 10 2 Figure 17 Core Assemblies Linear Power Ratings at IlOC4 1681-40 l

I 149 l 1

'\

ASSEMBLY NUMBER f

308 313 2.0 1.9 xxxx AVERAGE 3.7 3.5 anna PEAK 4.9 307 4.6 312 xman 3-SIGMA 5.5 3.0 5.1 2.4 umnu , 3. SIGMA

• OVERPOWER 5.8 47 302 7.7 306 6.2 311 4.7 8.6 4.1 6.9 2.8 9.1 8.3 5.4 9.7 305 7.1 10.7 301 310 11.9 10.8 5.1 8.0 3.0 6.8 10.3 12 4 5.7 50 12.1 304 7.5 210 52 14.5 5.4 5.1 13.5 5.5 8.4 3.0 16.2 7.4 7.1 10.8 5.7 8.1 203 7.5 211 8.5 51 49 12.7 91 14.2 5.6 8.4 2.8 9.5 6.0 5.3 5.4 7.0 11.0 7.9 71 45 12.9 204 2M 43 9.1 0.1 48 6.2 1- 5.5 8.0 2.4 6.6 10.1 9.1 5.4 8.1 7.9 7,1 10.8 4,7 9.4 9.1 12.7 6.2 213 44 46 8.2 28 205 10.5 II 9.2 5.4 14.2 6.9 1.9 6.8 9.7 5.1 14.5 7.1 10.3 3.5 i 8.3 4.6 9.6 102 16.2 [ 47 8.2 27 12.1 206 2C9 l 6.3 9.2 5.3 13.5 5.1 1.2 C/A \ 10.7 11.1 16.3 18.1 f 8.1 7.0 4'i 8'3 2.3 l 18.2 9.3 12 8.1 26 207 3o ,

98

/7i 37 86 9.9 20.3 36 101 6.9 8.4 9.6 10.4 68 9.7 14.5 9.1 13 5.1 7.1 8.2 9

10'y 201 8 5.8 77 86 3.4 208 l 18.1 6.2 9.1 6.8 2.0 7.3 11.1 10.8 6.9 9.5 8'l

.4 8.4 7.9 12.4 3. 7 67 9.1 25 14.5 202 d.9 9.6

~

11.0 99 11.1 10.1 6.1 16.2 4,7 5,5 12.3 11.3 10.8 C/A 16.3 7.9 9,i 11 18.2 14 9.1 24 10.7 32 16'.5 18 4 100 20.3 6.8 10.1 54 31,9 11.0 20.6 11.2 7.1 15.8 16.1 8.4 8.3 7,4 9.7 10 9.6 15 8.5 17.6 33 17.9 60 19.7 20.0 10.9 7.1 10.8 6.6 9.5 7.1 11.2 8.1 8.4 16.1 8.6 9.7 34 17.9 61 93 Il l 9 10.9 7.1 20 0 11.3 C/A 8.5 16.5 C/A 9.8 3 18.4 62 94 10.9 7.0 20.5 7.9 10' 8.4 9.5 POWER RATINGS IN KW/FT 7 jo 1 9 10.8 8.4 15.6 15.7

9. 7 29 17.5 1

10.8 37;4 ig 4 19.5 30 9.6 14.8 93 10.7 2 16.5 9.8 7.0 18.5 14.9 83 16.7 57 96 18 6 g,7 10.8 I4'9 69 8.8 16.6 18.6 12.2 13.6 15.1 Figure 18 Core Awemblies Linear Power Ratings at EOC4 1668-51 150

l l l I

O ASSEM8LY NUMBER 308 333 ""

21 20 AVIRAGE 39 36 g 48 anna PE AK

$3 3oy 3,3 j 53 """" SIGMA 5.7 32 26 ====

/

6.2 go f 3 SIGM A OVE RPOWE R l

82 66 311

$ 302 306 91 73 30 5.8 301 305 310 85 37 61 50 304 81 210

$2 90 32 6.1 49 203 81 211

$1 90 3.0

$8 45 204 76 212 43 [ 48 85 26 j 50 44 46 28 205 66 213

/ 102 47 27 7.3 206 2.('

36 48 209 i 5.3 13 l 24 i (CIA ) 37 101 12 26 20; 3p 62 3.2 36 l

i I

98 36 68 13 201 Si 208 9.1 2.1 39 i 67 25 20; 5.1 99 5.7 100 11 14 24 ,

32 i II 33 60 10 34 61 Ce A (C!A )94 3 62 l

j 30 gg 4 29 128 93 y POWE R R ATINGS IN MW F T 57 l

69 j

\

l l

I l

Figure 19 Core Assemblies Linear Power Ratings at IlOCS 1681-47 O 151

4 i

AutUHLY N4 IMH4 it O

3og 313 ""

25 23 42 ==== AVtHA4.t 45 \

S4 314 Pt An 57 6.4 307 37 60 ,2.9

===,=j

,, , 3 g,gga 3 sicvA ov, Hmnt ,1 70 .... f 7y 3

1. 81 3 6 'a 301 305 8. 4 310 f

93 37 69 j

rA 304 88 210 i

$2 99 37 69 i 8 2II S1 49 203 9

65 204 84 717 43 / 45 48 93 7 e3

[ *n 6 12 46 28 205 jit 44 81

/ 47 2# 206 y3 42 S4 7typ 702 60 i e, C/A 28 12 26 207 36 37 got 31 41 10 68 13 201 90 208 58 3G 10 1 y r, 4 r.

l'3 202 5y 99 67 64 CrA j4 100 11 14 32 60 to f5 33 61 14 C/A CIA 62 94 3 59 4 30 29 128 l'UWE R R A TINT,5 IN s ys Fr 93 y 57 69 Figure 20 Core Assemblies Linear Power Ratings at EOC5 1681-43 152

a d

D I

i l

g ASSEM8L Y NUM8ER j 113 '" \

.10 8 792 787 amas -NOMIN AL 822 815 suas -00 827 307 821 p2 mm - go 827 801 821 804 mann - 30 835 841 302 841 306 846 311 820 841 841 846 783 849 877 823 855 301 882 305 829 310 855 g49 883 838 829 796 887 873 828 52 892 50 879 304 833 210 1093 893 1072 879 g44 833 796 1135 1112 881 828 1161 1136 49 887 203 833 211 1171 1146 1090 887 847 833 793 y

1131 884 823

~

d!\liii/::i!\= hii s s i

= / ::!! = e x= 951

,e 1129

g73

== 815 C/A \ll 37 102 885 936 942 957 958 101

1. 47 1110 1149 1976 1154 1164 12 p 379 1089 \ g79 g

26 206 841 877 882 821 821 207 209 769 787 792

)

I 1104 943 1104 1186 896 1070 883 801 yg3 1143 1143 950 1110 835 98 874 1169 1180 36 1092 1169 1180

/ 68 1104 957 957 13 1099 1134 1144 201 848 841 841 208 792 920 1129 / 1142 1129 887 822 926 927 99 1155 \ 1169 67 1155 25 892 893 202 827 842 1165 885 1165 1088 820 827 878 \1179 95 11g7 g4g C/A 32 884 100 11 \ 941 / 14 1141 24 855

( 826 884 851 1090

\ 942 f 1089 1152 1091 855

,( g57 33 890 1128 1126 1133 863 5

896 8% 6 1153 H63 gg 1151 / 15 1158 1169 863 g,g 1075 1162 f gog3 1128 889 1111 1119 1138 895 61 1136 1145 IM 895 840 1146 1155 C/A 1107 875 C/A 94 H 31 3 881 62 846 1141 1058 882 872 884 1095 918 30 889 59 1119 4 924 PL ANT THDV CONDITIONS 1047 890 841 1129 820 924 TEMPERATURES IN 'F 1083 878 850 ADI ABATIC BO'Jf uARIES 1106 29 833 128 855 1115 1056 884 810 855 1094 863 93 1118 2 868 835 1127 1058 869 869 1095 875 1119 5,

1129

/

875 333 869 7 69 374 800 875 823 829 829 l-igure 21 Core Awemblies \lised \fean Outlet Temperatures at BOCl 166N-50 153

O ASSEMBLY NUMBE R

?A 313 A14 807 ==== NOMIN A L 853 844 maan 1 00 858 301 849 J12 ,,,, / gg 859 825 850 830 / 30 gyn gyy ===*-

302 876 Job 883 311 846 876 870 683 816 881 913 856 887 301 918 305 861 310 875 920 862

$ 887 920 Sc3 904 821 883 926 50 910 304 869 210 53 821 1071 926 1051 911 872 869 1088 916 863 1110 49 869 1135 51 1111 922 203 211 1086 1121 1066 923 875 869 816 3144 J 1120 1106 920 SS6 43 \ 5145 gg 1'29 1139 48 926 927 204 861 862 212

, 830 1094 1997 1063 872 1133 \115$ 1131 1102 916 877 1959 / 44 itgy 116; 46 938 1926 1136 28 1063 922 923 205 863 883 883 213 807 1170 f 1101 1141 995 1102 904 844 1168 102 1002 47 1126 27 910 206 84g 20g 1179 g30 1003 1082 1136 1065 911 870 850 781 C/A 985 1120 1104 913 80g 37 992 101 114S 12 1128 26 919 207 810 1095 993 1081 11 % 938 1138 1049 920 825 810 113g 1119 996 1066 870 1"Al 1109 98 1161 36 1943 / 68 1002 13 1119 201 8?b 876 876 208 814 939 1972 1076 1080 1075 996 1913 1164 / 1118 1107 920 853 1003 99 1938 / 1143 67 1831 25 926 202 858 1003 892 1153 929 1141 106S 926 846 859 937 1148f C/A 984 :096 881 943 100 11 991 14 1120 24 887 32 943 1074 991 1074 1129 1069 88#

875 901 giS 948 1112 1111 / 110R g21 33 954 60 1136 to 1136 16 1133 g22 1061 955 900 1146 1068 1146 1067 1943 1099 946 1104 1103 1123 34 95J 61 1129 1128 1132 1063 953 890 1939 1138 C/A 1102 934 C, A 1126 940 62 94 889 932 1I35 1055 1093 3

940 936 993 /[

30 1047 938 939 59 1916 1128 869 4

N 9'"I

/

1083 883 924 907 /

1106 29 932 128 913 1916 1058 932 874 913 1096 913 g3 1119 2 919 880 112g 1059 920 eri 1097 921 51 1121 g28 879 1130 PL ANT THOV (:ONDITIONS 920 TEMPE RATURES IN *F I

69 846 926 927 f

[ ADI ARATIC BOUNDARIES l 878 884 l M4 I

l l

l j l'igure 22 Core Awemblics \liset! \ lean Outlet icuiperatures al 1.001 166M-120 l

l 154 i

1

e

(

ASSIMBLY NUM8ER

.1 tag 113 au 815 809 aman NOMIN A L 855 846 maan 00 861 307 851 312 muun 2 'T 861 Syy 852 831 ,,,, / 39 873 879 302 879 306 885 311 848 879 873 885 818 885 917 856 890 301 923 305 862 310 891 878 823 865 862 820 924 907 862 52 g30 50 913 304 868 210 1041 930 1029 913 873 868 820 1077 1064 917 862 1099 51 1086 49 922 203 8,8 211 1108 1056 1095 1051 923 874 868 816 1086 1088 g18 856 1110 45 1111 48 924 204 862 212 43 1119 1073 1120 1056 925 873 862 831 1053 1105 i094 917 8 79 1088 1111 44 1129 46 11 7 28 972 205 885 213 1121 1063 1139 931 1127 1062 923 865 885 809 1099 986 1101 907 846 1123 102 992 47 1125 27 913 206 851 209 1133 919 993 1079 1134 1068 913 873 852 782 1117 1107 917 806 C/A 971 811 37 978 101 1942 12 1132 923 207 26 811 1062 978 1076 1152 939 1141 923 827 1o54 1098 1113 997 1092 873 1123 1138 68 1003 13 1115 201 879 208 98 36 g17 1132 1148 1084 1004 1080 1124 8 78 879 815 1066 969 1103 1122 1114 924 855 975 99 1127 I14I 67 1938 25 930 202 861 976 886 1937 1157 936 1148 1070 930 848 861 929 C/A 992 1102 885 935 800 11 1126 24

% 32 936 902 1087 999 14 1078 1136 1075 890

\ 873 1000 891 g12 949 1126 1115 1115 J3 955 60 1951 10 1141 15 1139 918 gig 1063 956 907 1161 1083 1151 1069 1149 1102 956 1121 1105 1126 34 962 61 1147 1130 1135 1075 963 915 1137 1140 C/A 1114 %6 C/A 1139 972 94 893 1149 1078 3

973 62 1114 j

[

938 1118 1161 30 944 59 1143 4 1188 1056 945 891 1153 895 1200 1o94 936 g41 1117 29 942 128 947 1127 1071 942 885 947 1111 927 g3 1135 2 933 888 1145 1076 934 929 1115 PL ANT THDV CONDITIONS 935 57 1140 TF'1PtRATURES IN 'F 935 886 1150 mi eABATIC 80UNDARIES 929 69 935 85i 935 884 890 891 Figure 2.1 Core Awemblies \1ised \ lean Outlet Temperatures at 150C2 166 &l21

% 155 i

O

O r ASSEMBLY NUMBE R 308 313 g47 g33 seus NOVINAL 886 assa - U 899 905 592 312 av ,a 20 307 906 862 892 867 uss 3H 921 928 302 927 306 934 311 884 928 913 935 846 928 965 898 934 301 972 305 904 310 934 915 973 898 904 852 967 948 906 52 9 74 50 952 304 912 210 1022 975 1007 953 908 912 852 1056 1039 959 906 1077 51 1060 49 965 203 912 211 1086 1034 1068 1023 966 909 912 846 1C70 1056 960 898 43 \ 1092 / 45 1078 48 967 204 904 212 1045 1086 1024 967 908 904 867 1045 \ 1101 /

1002 1081 1058 959 928 1105 44 1104 46 1080 28 965 205 934 213 1115 1054 1113 979 10b8 1029 966 898 935 838 1092 1027 1063 946 886 ille 102 1034 47 1085 27 952 206 892 209 1126 975 1035 1041 1094 1035 953 913 892 800 1022 1077 1071 965 831 C/A 972 J37 37 1029 101 1100 12 1093 26 207 10s,g 1030 1041 1109 99u 1102 1025 973 862 837 1096 1077 1020 1059 921 98 1121 36 1100 68 1040 13 1081 201 927 208 1009 1130 1041 1109 1047 1048 1046 1090 915 928 847 1063 1077 1084 1083 967 899 1070 99 1100 1107 67 1106 / 25 #74 202 905 1071 955 1109 1116 997 1040 975 884 906 999 1049 1115 [ 928 C/A 1076 32 1005 100 11 1056 14 109g 24 934 943 1006 970 1057 1057 1060 1108 1047 g34 984 1016 1095 1098 108J 1818 990 33 1023 1024 60 1128 to 1069 1122 / Ig \ 1106 991 1050 978 1087 1026 1108 1132 f 1051 1088 \ 1116 1111 34 1033 61 1133 1112 1120 1059 1034 991 1143 1121 1097 1041 C'A C# 1120 94 3 1048 62 ggg 1130 1063 1049 1104 10G4 110i 1147 30 1011 1011 59 1125 1174 /

1043 95g 113s 97j 1185f 1085 1004 1019 1108 29 1011 128 1026 1917 1065 1011 957 1027 1103 1001 g3 1127 1007 PL ANT THDV CONDITIONS 2

960 1137 1068 1008 TEMPERATURES IN *F 1M 1107 ADIAB ATIC ROUNDARIES 1012 57 1931 1013 961 1141 1006 69 1012 928 1013 967 973 974 l-igure 24 Cure Awmtilin \liwal \ lean Outle? lemperatures at 1 OC2 16 tim-I l 4 O

156

N

\

a gASSEMBLY NUM8ER as 108 854 313 846 asaa \g NOMIN AL gog 897 nana 00 915 307 903 312 mass 20 915 870 904 879 unas 30 933 945 f 302 939 306 951 311 894 940 929 952 858 941 986 915 947 301 992 3% 92I 310 948 928 993 916 921 866 984 969 g2y 52 ggo 50 975 304 g33 410 1090 ggi 1076 976 g30 g33 866 1131 1118 987 927 1157 51 1141 49 993 203 $2 211 1967 1102 1151 1o94 994 934 933 858 1133 1136 9*3 915 43 1158 45 1162 48 M1 204 921 212 1089 1169 1115 1172 1093 1000 930 921 8 79 1126 1146 1135 987 945 1152 44 1973 46 WI 28 993 205 951 213 1163 1093 1983 gg4 1971 1093 994 916 962 846 1130 948 1135 969 897 '

1156 102 954 47 1160 27 9 75 206 903 20g C/A \ 1167 878 926 955 1105 1144 1971 tog 3 1135 976 929 986 904 804 837 4 37 932 101 1170 12 1161 26 992 207 843 '

1073 933 1091 1181 894 1171 1075 993 870 843 1108 1128 947 1115 933 38 1133 36 1954 / 68 954 13 1140 201 939 208 860 1943 1072 1090 954 1150 940 903 1107 1164/ 1128 1092 1122 928 984 854 909 909 99 1132 \ 1153 67 1147 / 25 990 202 915 909 834 1142 1163 878 1081 991 894 915 868 C/A 926 1157/ 1109 941 32 873 100 11 932 14 1934 24 947 820 874 844 1071 932 1066 1144 1090 948 849 881 1106 1101 1131 s 854 33 886 60 1130 10 1125 15 1157 854 1045 887 843 1140 1047 1135 1063 1167 108i 880 1079 1098 1104  :;4 885 61 1103 1122 1114 1049 886 832 1112 1132 C/A 1085 866 1108 CIA 94 3 871 62 839 1918 1038 872 859 875 1074 901 JO 880 59 1096 4 goy 1029 ggg 835 1105 314 1064 907 869 342 1086 gg 8 75 847 128 1094 1039 875 847 824 1075 855 93 T NW CONWTIONS 829 06 1040 1076 TEMPE R ATURES IN 'F 6

6 57 1098 ADI ABATIC BOUNDARIES 868 829 1'07 861 69 867 797 867 818 823 824 Figure 25 Core Awemblies Mimi Mean Outlet Temperatures 3t ilOC3 166M115 O

d 157

I i

O 7ASStMHLV NUMBE M g 8 km NOMINAL n'a - 00 963 947 ""

969 J07 953 20 312 """ 3" 970 914 954 325 988 1004 302 995 306 1011 311 935 998 978 1012 898 986 1040 966 993 301 1047 305 972 J10 994 967 1048 9% 973 909 1026 1012 983 52 1033 1018 304 989 210 50 990 1047 1034 1032 1019 972 909 1083 10e y 1034 983 1106 51 1089 49 1040 989 211 703 1116 10 % 1098 1045 1041 g78 990 898 1088 1082 1040 966 43 1060 1112 [ ag 13PS slog [ 48 1047 204 g73 912 / 212 925 1043 1048 973 f 1096 1121/ 1098 1114f 107g 1034 1004 1121 44 1121 [ 44 IW2 28 1040 206 9%

1011 1012 213 8114 1130 1067 G73 till 1043 1041 1104 1131f 1036 1079 1012 947 4#

102 1042 1052 1101 27 1018 [ 20s, 953 954 209 829 1043 104S C/A \ 1128 1938 962 1022 1087 1HO Illo 1081 1019f gy8 1040 870 37 1029 101 12 1104 26 1047 207 SFS 98

/ 1063 1100 1124 1030 1048 1083 1106 1119 68 972 1035 1042 1183 13 1031 1066 1088 1048 20s 914 988 995 8 76 208 36 982 1047 1134 1045 1079 II19/[ 1048 1082 1043 104g 1076 109F 967 1026 996 895 963 1064 09 1102 H 111605 [

67 \ 1099 / 25 1033 202 969 970

%2 1037 1034 9 35 10 % 9,8 980 gitt CsA

/ 1022 \ 1108 / 10hF sub 11 1028 14 1090 993 32 986 100 24 012 987 g37 1044 3079 1042 1098 1047 994 959 ggt 1079 1076 1

/ ,0n N :::/ 9:\ =

1016 990

,0::

1073

= / ,0: \ =0g3 /

1070 1108 34 1043 997 61 926 IM [ 1093 [

C/A

\ 1099 1079 998 977 110h / Cs A 1102/

94 1102 3 983 62 \ /

/ 922 972 till 1036 1072 984 979 1044 k [

30 978 979 59 917 1094 1103 906 4 1051 1%2

[

1032 1064 966 951 7

IJ8 29 972 128 957 1097 1044 973 909 958 1080 956 93 1103 2 962 917 1113 104% 963 966 1082 7 g PL ANT THDV CONDITIONS 96b if MPERATURES IN "F 69 986 971 [ ACIARATIC HOUNDARl[S SPF 972 /

932 933 figure 26 Cure he111blies %Iigest \leali Otitlet Ictiiperatiires al 1003 1664l18 158

\j i

p ASSEMBLY NUMBER 308 313 85 88 7 877 anna NOMIN AL 951 937 nsaa Ou 957 30; 943 3gy saan 2 rl 9'>8 90g 944 gty mass 30 976 994 302 882 306 1000 311 924 883 967 1001 892 973 1026 958 979 51 u133 305 965 310 980 954 1034 947 965 904 1010 1001 975 52 1017 50 1008 304 981 210 1028 1018 1020 1009 966 g82 904 1063 1053 1025 975 1085 203 981 1094 gg 1075 1083 / 49 1031 972 982 211 892 1038 1068

/ 1036 1072 1032 1032 958 43 1090 45 1094 48 1039 965 212 204 1099 /[ 1052 1031 1103 1037 1040 965 917 1064 966 1083 1072 994 1086 1025 1106 1095 [ 44 1115 46 968 1095 1104 / 28 1037 1031 205 947 1000 1001 213 877

/ 1041 1075 1029 f 1072 1032 1001 937 102 1036 47 1095 27 1008 206 943 209 C/A

\ 1098 1107 957 1016 1037 1048 1083 1104 1036 1071 1009 967 1026 944 824 863 37 1023 101 1106 12 1094 26 207 868 1033

/ 1052 1088 1024 1047 1082 1115 967 1029 1103 1020 1053 1034 905 976 869 93 till 36 \ 1105 / 68 1036 IJ 1075 201 982 208 1121 1047 1037 1078 1120 1048 1083 \1114 / 1081 1033 1063 1083 954 1010 983 887 951 1146 gg 1106 1104 67 1085 25 1017 957 202 1156 946 1115 II13 957 1094 1020 1018 924 958 1002 C/A 1016 1048 g73 100 1070 32 834 1008 945 11 1022 14 \ 24 973 1009 1047 1023 101g 1029 980 e 987 1001 1082 1050

\ 1078 1063 CD 993 iOO7 d

/

994 33 1008 60 1105 1114 10 1072 / 15 1085 /

1060 944 1029 1098 1000 1062 1081 / 1009 1040 1094f 1122 34 1006 61 1084 1061 1132 1061 1007 943 1093 1069 C/A 1100 999 CIA 9 105 0 10 987 3093 111g 30 994 1053 996 59 934 1117 1126 927 4 1143 [

1090 gay gyg 1154 /

1813 1123 29 993 125 921 985 /

1066 994 986/

1105 87I 93 1129 2 877 9 1139 978 PL ANT THDV CONDITIONS 987 57 1931 TEMPER ATURES IN 'F 988 928 1140 ADIABATIC BOUNDARIES 980 69 986 895 987 937 943 944 i

Figure 27 Core Awmblics \lised \ lean Ouilet Temperatures at HOC 4 l 166 *]19 rh

( l 159 L/

i 4

8

O r ASSEM8LY NUMBER 308 921 313 g0g

,,,,\

,,,, \ . 0 0 gog, ggt M2 gy4

- " 20

'I 307 9 39 980, 98 312, 95 s """..

.. 3 ,,

1016 1036 302 1023 306 iO43 311 956 1024 1003 1044 922 1006 1063 993 1012 301 1070 30g m 310 1013 1071 975 1000 935 gg4 1039 1028 1011 52 50 1035 304 101; 210 1046 1004 1047 992 1036 995 1018 g35 1038 1022 1053 1011 IL42 49 1059 203 101y 211 51 1001 922 1064 1011 1050 1003 1061 10'8 1060 gg3 1043 1034 1055 48 1067 204 999 2 43 1064 45 995 1063 1000 1068 1000 y 1019 1072 1019 1053 1052 1031 1036 1052 1052 28 1059 1074 [ 44 1073 46 1061 975 i

1082 1021 1060 1001 1028 l0 M 1083/ 10281063 1078 1032 1035 W

yy

/ 102 \ 1085 [ 47 1052 1060 1003 1036 C/A

\1085 1094 / 1022 1078 \ 1086 f 1007 1039 1034 1055 1063 207 888 893 37 1086 tog 1060 12 1070 1068 1063 1071 939 894

/1042 1077 1087 1010 1042 1021 1078 9

1016

d= d =/d = d \ E 10 1 g 1

=

99 10 1078 1026 1006 1087 C/A 1046 1012 11 1086 14 24 32 1094 100 1018 1087 1053 1005 1013 1C08 1021 10e6 1028 1037 10g4 1052 1040 1077 1073 15 1057 33 60 10 1061 1084 1091 1081 1069 999 1065 1086 1040 1092 1026 1020 1030 1075 1083 1053 1090 61 1075 1051 1098 34 1059 1107 1040 1091 1028 1084 1086 i 1076 c3 C/A 1099 3 1093 62 94 1018 1108 1038 1094 107e 1074 1074 1109 30 1081 Sg 1097 4 1134 [

1083 1106 1016 1038 1014 1144/

PL ANT THDV CONDITIONS 7g 1 / yyg 9

1009 1079 Tf MPER ATURES IN OF 1905 1054 1077/ ADI ABATIC BOUNDARif S 1

g3 1025 112 1054 1071 1083 1092 1090 57 1115 1091 1075 1125 1083 69 1090 [

995 1091 /

1047 1054 1055 figttre 28 C u rt.' Awelliblic% \liged \le311 Otillel let11per3ttartw at i 004 166449 160 I

l i

l

t a

U r ASSEMBLY NUMBER 308 **

313 924 913 mana NOMINAL 997 mann O ff

,00, => 98,1

,, 3,, .... 2 .,

==== 3" 1004 ,', 988 ,[

302 1030 1051 311 306 1031 1062 9 in: 'a 1022 i

• • ** 30" 102 #$

1022

$1 49 203 1002 43 45 48 204 2 l 1044

/ 4-102

~\ 47 3-27 2-206 981 987 209 988 847 891 (C/A 37 101 12 26 207 944 1024 896 897 98 36 68 201 1030 208

/ 13\ 25 1031 924 997 99 67 202 IN3 C/A 11

/ ,

32 100 14 24 O 33

/ x /1x /1x /  ;

C/A C/A 94 3 62 30 $9 4 29 128 j PL ANT THOV CON 0lT10NS sy TEMPERATURES IN 'F ADIABATIC 8OUNDARIES 69 Figure 29 Padial lliankel Scomil Row \ssemblics \lised \ lean Outlet iemperatures at 11005 166M-I l 2 l

161 l

l

)

l i

O r ASSEMBLY NUMBE R i ma a NOMIN AL 1031 1010 mana 00 i 1030 307 1017 312 msna 20 1039 gyg 1017 993 n=== ~30 1083 1067 302 M6 1088 311 1064 1066 1063 956 1031 301 305 1038 310 1039 971 1052 60 304 1059 210 1060 gy1 2

$1 49 203 9 211 1 0 N 1031 43 45 43 204 1038 212

/ 1039 993 1081

/

/

" - =  :::

1010 102 4y 27 206 1017 209

/ 1017 867 913 (C/A -\ / N 12 1057

\ / "

\

= 22:

1031 99 67 25 202 C/A 32 Icg 11 14 24 33 60 to 15 C/A C/A 3' 4 59 29 12e

• PL ANT THOV CONDITIONS TEMPERATURES IN 'T 5,

A0lABATIC BOUNDARIES 69 l

! Figure 30 Radial filanken Second Row Awemblies Mised l

Mean Outlet Temperatures at 1005 166M-107 162

b t

f

(

)

i e

1400 1400 l-C c- }

[

g 1300 - -

1200 j E = l

" f

$ 1200 - -

1000 $

g w  !

m a:  ;

w 1100 - -

800 5 m .i E p' -

E

' {

c.o 1000 -

/ 600 E ' r o ]

@ 900 - '/ -

400 z

w a

u s' W

) 800 -

- PLANT EXPECTED CONDITIONS -

200 2 2a HOT SPOT FACTORS E j k

g 700 0

CY1 r = CY2

128 DAYS 200 DAYS Figure 31 Iifetime Cladding Temperature / Pressure IIistory in Fuel Assembly Tare, Orificing Zone i 1668-32 163 O

O 1400 1400 D

C 1200 5 1300 -

5 -

E 1200 -

7 1000 b a: -

E / $

3 110r - / -

8# 5 w / W o / =  ?

1000

/ -

600 E

o /

/ 3 Z

o 900 -

p -

400 $

a / PLANT EXPECTED CONDITIONS

/ 2a HOT SPOT F ACTORS 200 z 800 -/ -

=> #

z I

k 700 O z

= CY3 - -

CY4 =

275 DAYS 275 DAYS Figure 32 Lifetime Cladding Temperature / Pressure Ilistory in Fuel Assembly *tl01, Second Core. Orificing Zone i 1668-33 O

164 1

h i

C 1400 1400 w

cc 1300 - -

1200 h

5 1200 - -

1000=[

t w D

• 1100 - -

800 @

e z E 1000 -

m' 600 m

=

m " " z o

,+ w 5 900 -

400 d u

H ANT EUECTE0 CONDIDONS

[3 800 -

20 HOT SPOT FACTORS 200 x

I

$ 700 O l

!= OY1 128 DAYS

!c CY2 200 DAYS r! l i

i Figure 33 Lifetime Cladding Temperature / Pressure IIistory in Fuel Assembly #43, First Core, Orificing Zone 2 i

1681-13 O

V 165 l

l,

O 1400 1400 C

c-w cc 1300 - -

1200

=

r' E 5 1200 - + -

1000 w E s' 5 5 - M H 1100 -

g' -

800 g 9 ' "

/

C g 1000

,'/ + -

600 o #

/

d 900

'- PL ANT EXPECTED CON 0lil0NS 20 HOT SPOT FACTORS 400 E 800

'/ 200

,/-

x I

700 O

!= CY3  ;-!' CY4  !

275 DAYS 275 DAYS Figure 34 Lifetime Cladding Temperature / Pressure Ilistory in Fuel Assembly a 37.

Second Core, Orificing Zone 2 1681-14 166

c 1400 1400

'O W  ;-

m 1300 -

1200 -

m + -

i;5 E 6 5 1200 - -

1000 *

o. m I =

w M

• 1100 - -

800 m e E E1000 -

600 5 5

o , 5 a

j "a

• u 900 -

, +

400

" ,, - AN D ONDm0NS i [3 800 , , 20 HOT SPOT FACTORS 200 x

l < l 0

4 E 700

!= CY1 =l= CY2 =l j 128 DAYS 200 DAYS i,

4 4

i i

I i

Figure 3$ Lifetime Cladding Temperature / Pressure Ilistory in Fuel Assembly #45, First Core. Orificing Zone 3 1681-15 167

O 1400 c .1400 p 1300 1 00 m

E m 5 5

n.

1200 -

1000 6 w

a cc w

• - 1100 -

/

/- 800 o g

e / y

/ a.

E 1000 - -

/ m 600 i5 / o 5 # 5 900 - -

400 a

/

f# g

{ PLANT EXPECTED CON 0lil0NS

,,,,, # p 200 2 2a HOT SPOT FACTORS x

2 700 0

CY3 CY4

275 DAYS 275 DAYS Figure 36 Lifelinic Cladding Temperature /Prewure Ilistory in Fuel Assembly :t2, Second Core. Orificing Zone 3 166M-34 ,

l I

l l

168 O

l l

l l

,C 1400 1400 w

m 1300 2 -

1200 g - -

< =

5 1200 - -

1000 $

k w

=

m

• 1100 - -

800 y 9 PL ANT EXPECTED CONDITIONS w 2a HDT SPOT FACTORS E 1000 - -

600 5 3 0 "

5 u

m 900

$ 800

'y

r 400 200

."j

• I 3 /00 0

< CY1 = = CY2 =

128 DAYS 200 DAYS i

Figure 37 Lifetime Cladding Temperature / Pressure Ilistory in Fuel Assembly

1. 52. First Core, Orificing Zone 4 1668-35

\

169 1

> s

O 1400 1400 C

1200 1300 r

5 1200 1000 E r "

a N 1100

- w 800 h

• #', E o

's "

z 1000 -

600 2 5

O /'- = @

5 900 -

400 PLANT EXPECTED CONDITIONS E 20 HOT SPOT FACTORS 2

x 800 200

=c I

700 O I= CY3 =I='

cy4 &

275 DAYS 275 DAYS Figure 38 Lifetime Cladding Temperature / Pressure llistory in Fuel Assembly # 24.

Second Core.Orificing Zone 4 1681-18 170

g 1400 1400 c-E 1300 _

1200 = ,

? E

~

< w oc w 1200 1000 cc

& 3 b E

• 1100 -

PLANT EXPECTED CONDITIONS 800 g C 20 HOT SPOT FACTORS n.

4 E g

1 E 1000 600 a w a " "-- '

u

< 900 400

"' =

d

! 800 ,,,,,,

""""~ -

200 m .

• l 0  !

$ 700 l= CY1 128 DAYS rl CY2 200 DAYS

=l l t

a V e i

i l

1 1

l Figure 39 Lifetime Cladding Temperature / Pressure Ilistory in Fuel Assembly #49, First Core. Orificing Zone 5 l l

l l

1681-19 )

i I .

I L

I I I l

t

. l i

171 )

O 1400 1400 C

1300 2 -

1200

$ 1200 -

1000 g km 8 a

• 1100 -

800 g 9 ,s* a-c /

/ -

600 b

g 1000 -

f z W

o /

o /

/ -

g 900 - ## -

400

/ s ',,, PL ANT EXPECTED CONDITIONS l

I 800 ,e'f 2a HOT SPOT FACTORS -

200 y / ,

3 1 0

700

CY3

=l= CY4 -

275 DAYS 275 DAYS l

l Figure 40 Lifetime Cladding Temperature / Pressure Ilistory in Fuel Assembly #49 Second Core. Orificing Zone 5 l

1681-20 0

172

O j

u.

1400 1400 l E 1300 1200

s =

E E cc 1200 -

1000 ~

E E E m

1. 1100 - " -

800 E e E ,

y 1000 600 g I i5 PLANT EXPECTED CONDITIONS z 900 - 2o HOT SPOT FACTORS -

400 m

g 800

200

=--

x

$ 700 0 CY1 = =

CY2 128 DAYS 200 DAYS INNER BLANKET INNER BLANKET 4

I Figure 41 Lifetime Cladding Temperature / Pressure Ilistory in Assembly #98, First Core. Orificing Zone 6. Row 6 Alternating Position 1668-36 f

173 l

l

. . .J

9:

1400 E 1400 i N

R 1300 1200

$ =

E 1200 - -

1000 E a

w -

m s -

oc o 1100 800 g

<a cm PLANT EXPECTED CONDITIONS w 2a HOT SPOT FACTORS -

600 E

$1000 E o o d 900 -

400 3 a-o ,s =

3 800 200

-,,,,,__--J-y 0

700

= CY1 - -

CY2 =

128 DAYS 200 DAYS INNER BLANKET FUEL i Figure 42 Lifetime Cladding Temperature / Pressure Ilistory in Assembly #62, l First Core, Orificing Zone 6. Row 6 Alternating Position l

1681-22 l

l l

l 174 l

O

i t' I l

's i

_) { \

l<

i f,

t 1

.i 1400 1400 }'

c v v

o_.

w 1300 - - 1200 m =

(1 1 o

• 2

$ 1200 - -

1000 -

w w

_ m x

o a 800 N i

1. 1100 w a g PLANT EXPECTED CONDITIONS g }

1. 20 HOT SPOT FACTORS ';

ts 1000 600 m '

E / $

@ / w at 900

/ -

400 Z$ #

, /

E 800 -

/ -

200 m - - d g

y 700 0

= =

CY3 CY4 275 DAYS 275 DAYS 2

INNER BLANKET FUEL Figure 43 Lifetime Cladding Tempera ure/ Pressure Ilistory in Assembly 62.

Second Core, Orificing Zv.ie 6. Row 6 Alternating Position 1668-37 175

.q

O 1400

,{1400--

1300 1200 h -

r

• w E 1200 -

1000 m $

g -

w l0 1100 800 g .

i

-ci E E 1000 - PLANT EXPECTED CONDITIONS -

600 $

5 20 HOT SPOT FACT 0ftS ".j o

ca 5

o 900 - -

400 I

) 800 - 200 I

x

,,,___ , -------- g --

• f 0 x 700

!= CY1 128 DAYS

=!' CY2 200 DAYS

=!

l l

1 O

l Figure 44 Lifetime Cladding Temperature / Pressure IIistory in Inner Blanket Assembly n 67.

First Core. Orificing zone 7 1681-24 l

0

- 176

?

__________m

t k

i i

N 1

i 1400 1400 c

?

1200 .

E m

1300 -

E -

1000 h

5 1200 a e m

4 w

• 1100 -

800 g  ;

i E

9 n. <

PLANT EXPECTED CONDITIONS -

600 s E 1000 -

. s HOT SPOT FACTORS m

3 z w

$ 900 400 g u

I

=

2 x

800

-"" ~

200 '

< f 0

3 700

= = = "

CY3 CY4 275 DAYS 275 DAYS i

5 l

i Figure 45 Lifetime Cladding Temperature / Pressure Ilistory in Inner Blanket Assembly #99 Second Core, Orificing Zone 7 1668.18 .!

1 l

I LO

! 177 l

_ . _ y

O 1400 C 1400 E

m

@ 1300 1200 e

> E'

=c 5 1200 -

1000 m a  :

W -

E 80G 1100 +

E 1000 -

600 E PLANT EXPECTED CONDITION 3 z 5 U 20 HOT SPOT FACTORS "

400 g 900 I

= 800 -

200

< l 0

. I 700

$ = CY1 =I=

g CY2 &

128 DAYS 200 DAYS Figure 46 Lifetime Cladding Temperature / Pressure Ilistory in Inner 131anket Assembly # 12, First Core, Orificing Zone 8 1681-26 178 9

h

h t

i

?

. e

1400 1400 t

r -

o

! w 1300 - -

1200

=

m

>= ~

5 1200 - -

1000 @

i I C [

a m

'60 1100 800 g l-

-en w i =

i c.3 PLANT EXPECTED CONDITIONS n-1 z 1000 - -

600 E g 2a HOT SPOT FACTORS

' o z

< atC w d 900 - -

400 g i zm i

! E 800 - ~ "" 200

p g ___ r 2  ;

l, 700 0 1

! = CY3 = = CY4 &

275 DAYS 275 DAYS i

I l

l i 1 k

i  !

1 i

s 1

) Fieure 47 Lifetime Cladding Temperature / Pressure Ilistory in inner Blanket Assembly #46. Second Core. Orificing Zone 8 16H 1 -27 I

i 179 l

E b

1400 1400 E 1300 -

1200 IE

=s 1200 - 1000 w

$ 4 w 1100 -

800 $

o

=

m CC O '

E 1000 600 y $ PL ANT EXPECTED CONDITIONS b O 3 20 HOT SPOT F ACTORS $

" 900 -

- 400 g 2

2 E

~"~, - 200 N 800 -

+

m 700 0

• CY1 **- C Y 2 :  : CY3  : CY4 ~

128 DAYS 200 0 AYS 275 DAYS 275 0 AYS Figure 48. Lifetime Cladding Temperature / Pressure Ifistory in Radial liianker tssembly =201. Orificing Zone 9 4

9 9 e

4  ; <' t 0 -

1 e

n o -

Z g

i n

i c

w$EmC' E3Zwg C

f i

r O

0 0 0 . .

0 0 0 0 0 0 0 3 4 2 0 0 0 0 0 1 1 1 8 6 4 2 0 0 2 .

- - - - - r y l

b

- S Y

m _

w

- -~

4 A s ,

Y D A C

" 5 7

t e -

2 k

- l n

a

- = B _

l

- = i a

d S

N

- S R a _

O Y I

T - 3 A n .

I D Y D i N R O

S

- C 5 7

y r

C O T 2 o t

D C i s

E A l I

T F =

C T e E

r P OP = u X S s E S s T Y e r

T N O A

A H L 0

- 2 Y D P

/

e C 0 r P 2

- 0 2 t u

a r

c e

~

1 Y

sYSA D T p

m e

C .

=8 2

1 i g

n d

0 0 0 0 0 0 0 0 d 0 0 0 0 0 0 0 0 a 4 3 2 1 0 9 8 7 1 1 1 1 1 lC e

C Eo 2# c- 3 E5o<aoEoEE<E i m

t f

e i

L ,

9 .

4 .

e r

u g

i F

_$v _5 g- .

l > ;  !!

5 E

3 1400 1400 C

- 1200 a 1300 m

3 1200 - 1000 h w = w a

w E

• 1100 -

800 5 9 x

$ PL ANT EXPECTED CONDITIONS -

600 g 2o HOT SPOT FACTORS g1000

< w 400 k s

s 901 -

E E 800 2

__------""~o 200 0

700

• CY1 = c CY2 = r CY3  : CY4 -

128 DAYS 200 DAYS 275 DAYS 275 DAYS Figare 50 Lifetime Cladding Temperature Pressure Ilistory in Radial Blanker Assembly =206. Orificing Zone 11 8 9 e

~ .

~

e- .$ ~

. O__

- F gi u Ec 9r$ DCm $-

r e

• NE:  :

s < C 5 1 1 1 1 1 1 7 8 9 0 1 2 3 4 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 L

i f l e

t 1

2 _ - - - - -

i 8 C _

m D Y e A 1 C Y  : _ 4 l

a S d  : _

2 d 0 i

n 0 _

g D C Y _

T A 2 _

e Y m S p & _

e r

a  : _

t u

r 2 _ P L

e 7 A

/

P 5 C _ 2 N aT r D Y es s

u r

e A

S Y

3 _

S HE OX T P P

O E

C T

O I

l = _ T ED i F s

t c _ AC o

r C O T N y

i 2

7

- OD RI T n 5 C R D Y - S IO N

a A 4 S d Y -

i a S l -

B  :

l a -

. n k

e -

t 2

A 7 -

s 5 C

s e D Y -

A S m Y -

b S y

l 2

1 1 1 1

,2 2 4 6 8 0 2 4 0 0 0 0 0 0 0 O

r 0 0 0 0 0 0 0 0 i

f i

c i Z=$tbmb z g

n Z

o n

e 1

2

_ O

O ASSEMBLY NUMBER 3 na mann TIME OF OCCURENCE

/ 307\\ /

/ 312\\ namn MAX.CLA00tNG to TEMPERATURE IN FIRST CORE ('F) 302 306 311 7 301 305x 7 3,0x

,eJ,'\

" 12

.0C., / '"\ / N

,,j, x u a / ,,j, x 1

/ 0,x / 211 x

.C.,

"" / ,0l, \ " /,et,' \ / '"\ / '"\

/ .0",\ "" / E0/N" / 0<"\ / '"\ / '"\

( ,,,/S\ "" / ,0;',\"" /,0l,

.0 \

, / '" \ / '"

37 1152 101 1308 1 1298

/ 26 \ / 207\\

80C-1 '[ BOC-1 EOC-2 / BOC-1 \ /

98 1270 36\1239 68 1170 13\1285/ [ 201 \\ /

/ 208 EOC-2 BOC1 \ 80C-1 BOC-1 \

1206 99 1274 1285 EOC 2 67

" 77

/ 25 \

/ 80C1 \ /

/ 202\\

E3C2 C/A 32 E O C-2 1127 EOC2 1

BOC2 1194 [ 14\1271/ /80C1

/ BOC-1 \

24 \\ /

/

1106 33 1140 60 1267 10\ 1247 / 15 \ 1309/ /

BOC1 EOC 2 E O C-2 \ / BOC-1 \

1233 34 1156 61 1254 1245 BOC 2 E OC-2 C/A (C/A EOC2 94 1248

/ 3\ ug2 /

/ 80C 2 \ / EOC 2 62 80C2 30 till 59 1253 4\1307 E O C-2 \ /

[ 62 EOC-1 E OC-2 1217 29 \ 1114 128 1158 1120 80C2 \ EOC-2 g3 1239 2 1813 EOC 2 BOC2 1126 57 1243 EOC2 69 1127 EOC2 1057 l'igure 52 I melope of Assemblies Ntasimum Clashling ID lemperature at Plant 1.spected Conditium During the l' int Core ami Time of Occurence ,

l i sox.4 i O

184 l

l I

i ASSEMBLY NUMBER 308\

\ /

[ 313 \\ xxxx TIME OF OCCURRENCE

/ 307\

\ /

/ 312\\ xxxx M AX. CL A00 LNG 10 TEMPER ATURE IN SECOND CORE (*F) 302 306 311 301 310 BOC 3 BOC 3 EO 5 BOC-3 BOC 3 EO 4 EO 5 1311 43\1293/ BO[ C-345 48\1228 [ 204 \ 1153/ EOC / 212\ 5\

l BOC 3 \ BOC 3 \ / EOC4 \

213 1251

/ 44 1300 46 1307

[ 28 1218

[ 205 \ 1216 //

/ BOC 3 EOC4 / BOC 3 \ / EOC-4 \ EOC-5 \

209 102 1235 47\1307 / 27 \ 1189 [ 206 \ \1120/ /EOC5 C/A

\1258 E O C-4 B O C-3 \ / BOC-3 \ / EOC4 1306 26 1228 20 989 101

[ 12 \ 1310 f /

/BOC 337\1245

- \

36 B O C-3 / EOC4 \ / BOC 3 \ / EOC5 \

98 1236 1273

/ 68 1235 13\1296 / 201 \ 1194 / 208 E O C-4 BOC 3 \ / BOC-3 BOC4 \ / EOC4 \ / EOC-5 1247 1271 1271 1258 EOC4 9

/ \ EOC4 67

/

/ BOC 3 25\1205

\

/ 202\1150

/ EOC4 \

f 32 1262 100 11 1245 14\1264 / 24 \ 1153 /

( E O C-4 1244 EOC4 33 1247 BOC-3 1245 10 BOC 3 \

1?21

/ BOC3 \

1309

/

BO C-4 60 E OC-4 BOC 3 /

/ 15 1234 1246 1201 C/A / BOC 4 34 EuC4 61

/

f C/A BOC 3 [

1217 3 1260 62 \ /

/EOC494\1236 \ 80C 4 4 1258 EOC-4 \ /

30 BOC-4 1221 EOC 4 59 1229 EOC-4 /

[ 62 EOC 3 1222 29 1213 128 1234

[ 1167 B O C-4 EOC 4 /

1240 2 g 1209

/

EOC 4 BOC-4 /

1242 57 1241 [

EOC-4 /

69 1241 98 EOC-4 EOC-3 10i9 1171 l~ipure 53 1:inelope ul .\ssemidies \lasimum Cladding ID Temiterature at I'lant l Esglected Conditions During the Second Core and Time ol' Occurrence k./ 1668-42 185 I. _

O e ..x /r "

s ,'l, %r sk . %,':q.x

'" sX

" +N,'l, % >r M ,", %%

r M 1 % 'r M ,,, 4 s h ss 'i' >?, %'16 '"41% 're ll M;' 1

, +

'+ ,I,  %'e s X

,Yl,'" a ll-MlY% '

" ater's ,'yg "t~ +%,l,,tu sj,t,  % +x 1 % ',tls

,% a ,1,se>:' e ,,,

s es '"i s, 1% y4 ,spl"l,*s: s e "' 4 ll, % '"" %%% "'l'W," o

~ts

&~

sTl,s , , g~ll:p ll%)~" 4~,',', y," Wll,4 ~",gll

% "'% ,%,sTw "42%vn ""\s4l,% ,%~>W ~">Wy,% ~l,,5 :, e r:l"s;l,,

s l,'%';;e .  % '74 'll% X Ve"'% 1%,' O ey %;% "14 .g, % ~ %,,,'% "' ,,,y s," % ~4,",% ~% \

e/

l?%;l":,%t+Y s'"et

,,lr% '"%%"ll'st c% ~>n L % y rn e>r%t%;"w;;,y - , . ,

s,1,%"l.%":V e ~ y ,,, y

- 1/

l'igure 54 Core hiite lhict \listwall lemperatures as .l2" I lesaiion.

IM)Cl. Uniler Nenninal Conilitions 1668 II)')

O 186

~

C%

\.

w/

l

1. ' +X

$ ps, % ?

4/'s\

'e *yo $

ersk);,%'r5,':,%%X tI%>v415't'e',+y,'"x

• % * 'e  %  %

$I 's 6 #v g is3

'+16'"416'"#41 % 'r ell, %# *h

$ $0 g, t '" ### 85  % 72 #4 k\

+-T,6;ll+4">%'ll4,"/sa66'"@ l# 'll+N 7l,%'ll s\

h $ 922 6# '# 873 D h82 74: D s 748 so9 4 #

1. g E* #

$ go h 768 U 5 743

$ #a s " s75 #O si e *4 75

+ '" De 899 Y, [ $e 60 " 'S + 898 0 5 " 8' 'e ' 750 46 6 5 7

$ 'e, 7,3 & 4 45 @ T #s g,3 @h, go, , 7 8

C) *)l,' % 'll% ll(% ';,' % 24 'l,' 4 i % '" W",

8 Y #

1. # 77 # # 863 0 Y 907 771 4 48 8 D 782

,. u t , % ,' s t , k ' 75% r n,;,.y s <s$ "'n le% '">>11 791 4, , e g 860 , e, & TEMPER ATURES IN 'F st,s lll, ,y r"w g:y

• l '#

7 Figure 55 Core hide Duct slidwall temperatures at 32" 1.le$ation, 110C1. Accounting for Uncertainties 166M-110

(

v i

187 l

O

,"llA , % "'4;l"l'+A& , %3X r e ll, # -

s 'r> @", % 'i: #%>'tl% A ,,,l,'A

'4 s ', 4,i'M,L%Tel,l,#pK

, ,,,,6'I'@,

i:M,L

' sI #/"% !l%'r 4,'l, %#+X -

p""{'al#" 4,",#'""/r 4 ;l:%'!;s ," e,ll(A "

'+ g"st%" l ,:+ g":n, as v + ,%" x

' s ;,", s ." ,

s"l#r'sy" &"'%k yl,, og....sg,s~ s ); % ".s , ,

" a " if," ++ "lglw"%g'"

y ,% 2+g 4bk"lM", -y,,,y,"$% "lM,,

• st e'~ s

%yll,  % " -

s ';, p% "":n,,;;,5  %"ub' s ",>

% rp",s16"X,Yr sg,sy">,y,,, oq s,' a"<%';l

,; F +w .1,5 't"!w%'i: e...

% n%1ogl%"' *,,.

• c'"(,1 %"'  %'"% #7,6'"'#1. %'l: %Y

%7ll,4tf

-),

r" s 1, +%'l"*e' +d, Y & r%"l %,l!,9 ,_,,-,,..,

s,\l,e'ils,;,l,'y s

, F & ,,, y s,;,f I igure St- (. ore hide Dutt \lklwall Temperatures at 60" llesa, ion.

11001, Under Nominal Conditions 1668 43 188

1 j'"'N J O,

$ $$ #NSk

$ ,'l, o e "* ss ,',', %2s\ .

,,rs T",# >:> +4,'l, %2s\  ;

1. osf, 8 #o # I # 778 s s'!;s;+;il#'" S

'.,,'s,rr 'v's3l1l,M 6 "+& 'f.% 'I's Ie 'r s,

,',' #s- sa s +s.lll,4u

% 'i'y  ;;,es4' 4.;,"'5'en

, , 4 '+ 5 e, , 3yr  ? +o N '{$ ,g $ $ ,,

s ',1l %l,",6 "'u/ll, # ""sy,"' # ""+) 'il % "'M"O #jl,' s

" 780 [ #

1. ' 907 6 11 D 113 ' ' 20 s 75, 8 g

$ 10 0 '$ to 2 Y ##

84 '$ 111,

[ #s 780 8 98 [ ' 1015 11 5 $'120'[ -$ #

[ p s 773

} $s" sb 864 "# 765 6 ,03 [l+ tiis

1. 160

@ ,,, 796 , 110

$, ** "" $, , 1038,,

,">D #""+3 V.# 'r 4l",%j"-y,",(

b 845 #' '

o 7,1 O 85, e"l%'!" % '"& "% 'r & ll, y

1. Y 8

1. 9, g ,4 s s9 e '" W,',',eg sr ""r ':l, y 7s0

1. ' 107 8 D 185 TEMPERATURES IN 'F ns 1igure 57 ( ore hitle Duce \liilwall Temperatures al 60" 1 lesation.

11001, \ctounting lur Uncertainties l 166M10M o

(. 189 l

' (

O

,p <">-A 803 # II s'" M[,, 20 <4 >; xJr 8,8 4 '" A ll+.\

a909%:!'*9t6;;- 0,83

• 'il d4> 8,,* -l; sx to 7 [ gg 1071 [/ #

96 S # '# #

7

1. Y # ##

4I # D

@ # 81093 1034 [ g8 D sti @

16

1. ' E #

1. 6 f 1 [ #

906 N # #

73 h,# 4 'h g s 940 gg

1. s }g3g h 88 h # $N 3 e

1. ' D $

85 73 e 943 [ 078 [ 09 8 -

• 859 6 #8

[ $2 996 @  % 00 10 0 # #1060 g # 09 [ $

862 d 55 84

1. 1055 65 93 Y 799 3

h Y 883 @ $3 D 8 48 934 4# 065 844 82

1. 90 D D# 984 [ 995 v63 6 5 897 OE #

981 @ .r 041 9

867 2 Il '# "84 "96 y

893 ## # gsg 6 856

1. 964 ' 892 @ #

82 TEMPERATURES IN 'F

1. , cc s, 32 @ # 1017 o "' % e/

'+8?+7 ligure SM Core hitte Duct \lislwall temperatures at 112" l iesaliua, 11001, Uniler Nominal Contlitions 1668-44 O

190

v "3 U*

s, s

'i8 4 389 s 6, s q,1.48'l,y-l, % M

• llr % >"sx tretey:v:y s ?sts, >"%vd:.,e

+,

0, , 8 fO, si :

99 ' #

,"r w ',1,', "' g 'f 4 I % >y M !',, % ~ +x

'sl,",#7%s""e  %% % 'r%% ~k s'" g,ll,%5% ~>s;9 ""s'llo#jll"l,#'r9 g""-n 2%;%ll,%;"s M #

1. 936 178 8 1 4' Y #

832 D '

763 e 8 IUe ,", s t154, 84 p t 4se, 6

3 812 # , 4 '8 4 76 6 p:833:: s " s l: P h 6 ::. s "

n s:" n;ll, % ." %y;" %"e ""g sg,;\; s ">*e""x by s O~ s "%l;'sp lY%'l@,,;#";;, "O6";;;%"';'n o

%, *7 "9 g #

1. t o2o ' '#v 893 1025 # 129 o9 8 '

s1o 4 #

893

@ S 390 Y ##

100 1 #o es 1 4 O[ @ as D

• f' s2@ TEMPE R AT U R ES IN 'F e "' & ,":%""M, 9'

1. 99 O ##

1 s 'e 800

1. 783 '

Figure 59 Core hide Duct \lidwaii Temperatures at l12" I;tesalion.

11001. \ccounting for Uncertainties 1664104

\

'd 191

O 4 s

• \," s, % "' -M.,, 4, e s!-M",4;MS'l,e,,,i,s

,1,% 'E s eX s ,1, # 'P rM Q ,i, #4)~'l,4 ~ A s 'r s 1 e),~ M 1 % 'l e '1 %

sis,"'s,(4";<'sy s~

s

' sk , 4 '" &

i' WY,s4 ~

ey ,,%

"A~k PMk%~%,1.4'P4lL4~e Q',p e ":' dl'% &,ll% "'!% 'i %;sh,;<4 ek~

s ,1, 4 5 %'  %); t s

s>;;* ,,, ef ",\n\;l,  % .1, % %g n% = e'f; f ":sY,,,

s % lll%

s "' sWll, %

s .., % ~ L % j ' & ,;,,, *% " '

a "l% .:, &% ""'(s2 ,'%"ll'M

,,, % ' %\,"l%%' s%

' 4%sh:,.y y n ,,,4:l 9

sb.,,%;;%.%',i4"l4l,\@%;;'*,;;y

, .,.

• T W ,1, .2, ,,,

• mp%"'% %75 %% L V

,p M., %"' e,1, %p R ,.. s/

s ." / ,'r M , b 4 % 4 ,:, 5 s.+s1 tM.:,y

s. ,\;,s, ;p%'1y n -, , - , , ,

s*:,.,9/

% ,YS/

l'igure 60 Core hitic Dut-t \litlwall lemperatures at .4T l lesafion.

I 004. Uncler Nominal ('.rulitions I MiM-l 16 192

.h 839

$ h $

8 5OS ), % 4 88Y g ,1, %

1. 87 h $4 h 93 [

[ $'92 dI46 Ii 2

~ '"g 817 *p"$s'5 !7 s',i,'+X$!8 #s',','t, 8 8 8 [ $ 854 [,$' 5 [ $ 877 [ /

' t' '

• "' $ + e s "' -

%.*r'"s8?+)8.r's'"5,918  ;;s se ':e s 915 se7,'08 #s 'll s 88[#/s ,

7 e"s:+xej;;s8,8 e osgs Sces";s 8!,eo s:swer ';>

7 [ k 900 [ h 856 [ , *$ [ $ 96 [ [O 821

[ Y 877 [ gg 4 38

[4 ' MS [ Ys 853 [ 6 874 8 #

882 [ fs O 865 [ $ 854 d #$ 854 g ## 89

[6 85 h 897 [ **y 783 / [ h 85 *$ 840 88 [ 900 [ D 867 [g h 33 [ $ ' 853 k $ 871 [ # #

$ 900  % 864 [ is 851 772 [ h#871 I k# 773 d 902 [

k $ ' 894 [ ##

868 [ h 885 874 [ O#

s 894 [ k 892

@ k# 878 [ 14 $ 893 TEMPERATURES IN op 13 SH [' 82 i "' $ 895*/

• Ot(*Y Figure 61 Core Wide Duct Midwall Temperatures at 32" Elevation, EOC4, Accounting for Uncertainties 1668-117 193

M O

,,"x 4v

$h7, -1"l & ,?,#,\,, s a",\&,;%'i:&,%';;s\

i

,,, # "" e ,1,% 'il,M ,s.9 ""'s\'e,%'rl+\.4"'

,Mt,(, % "I%'?s Mll,  % 'l+X

1. s Ts T,s\

,7, n')l% " 5 .1, 4 " M % 'lL% ,l1, % '

s";sss*

+>,,, s "' s;,,, % "' e',,, % 'l,' % ,l1, % 'll h ,;/l %

so'i',1, s%,, %;'"l,4,,,'%:l e ,,,'%" % ""'s.s

"'>> s

, %"(' W... %

e ;s "&e", %" %"l

's),,,s ,m , g "'sw,,, ,,L&T.;,&,,,2l,%)""k=::s+

6 "'ss e,"',3,

"  %' ,,/

, T +N ,", # '"

s .T, % '"%'lgkT  % %el,  %"l%1,  %"> V9"lW,,,MT,Y O'" s+e, %",4 q arus,,s r s s ""> u ,,, y

• - ;, % "%&"; ,i,%,,,

,  % sps "

d ,,9' e l%1,, % "'%,1, %"

s),,,%yu,,rg"K;k",*

s">%,,,#"'%,,,Y r

s)l, s,j,,;4,;y 6 4Pj/

[,,6 h ,,, '/

s,o f I'igure 62 Core hitte llut, \lialwall Temtwratures a, 60" I lesatiim.

I 004. L'mler Nominal Comlitiom I 6t*45

l

%.s 983 980

1. s gg 10 0

[6 , 21 -

[ 6 1051 8 7, 059 #'s 1047 Y 58 093 g 1 4 [ $s\

1. s, #1014

[n# s'015 h # #1050 bh6 [26 e, 03 2

9

[ D 04 3

1. # 59 #v #1100 S '1030 82 # # 1060 # #1 35 S 10 6 # #1067 e d #

1. $Y O #'1090 [ #

10 s ' l 76 8 # #

1009 /

76 @ g,

1. #1 7 3

10 3

1. 06

1. 'I O [ / 6 942 -

01 83 0 #s # #

046 $

10g1 1049 029 102g 84 1 23 O #1099 #

06 1 6 ,

O 1 20  % i 103 O # Y1102 D #

1. 057 69 1026 1100 D 1064 ' [# 1017 855 7

1. 1064 0 's 11 5  % 858

• # "# 1 60 #

g #s '109 1 96 10 9 8 4 1092 # "1087

[ ##

1081 #> 1090 0 8 ,

3' TEMPER ATURES IN 'F g 1096'

'# 1079 l igure (> L (' ore hiile lhwi \littwall temperatures at 60" l iesalion, 1 O( 4. -\rrounting f or i!necitaintics 16f>511.1

v 195

O 9,. .x p .9

,h".%"'MT.$x

%'+N 4\ll8 %'l' \

8::49:%:::,'%';pp 9,8 %'l,\b1 %';;s\

8 #' 972 8 33 #

$ 970 @ #> 54 *

, 25

+#

91, % 'l,' #) 9:. 6 'll % 98, n'l,(h 9!. %;ll+.\

8 6 99 1001 [ ##

986 [ $ 942 3

D 52 @ '$ 10 4 8 'S # 989 @

1. 983 @

9i O 8 ':: skHrf8'*s%#8',',\% :!% "l % 9,. % '1l W,,#%ll,'s s

1. 85 8 #

1 08 8 $#988 ##

990 [ is

[ F 98 ,# 880 014 0

1. # #' ' 982 #

[,O # 46 *8 #s #I to ' @ ,

1. 945

• 33 6

1027

1. 992 h 7

' ' 985 h[ 80 * @ D 926

[&% 1017' #s # 89 [ ##

1005 @

, 979

@ ,8s 9b 6 7' 8 5

• 41 9;

42

• 8 , [' ' 982 8 #p 00108

• 1011 ' 8

1. *10 8 !' 972 873 'S 1006 8 #

" 1 12 [ $ 878

[ e 1 6  % 1003 102

1. 'e t s e@ $ to 006' TEMPERATURES IN 8 F

[ 6 N #

toot

' 1008# 6 1022' 8 ##

toit

1. 996 @

l igure 64 Core hide Duce \liitwat' lemperatures at il2" l_lesarion.1-004, t mfer Nominal Conditions 1668-46 l

l O

196 1

A 2,, s.

r .x 1040 [x ## #

7' [

g""'s3.;#e.Xh','O7/rs. s

++,,v/>t+.,",g""41##~',

I % #gggg } 'e, #e

%f #

g y% $ #s 1078 O#1 I #g

• O 076% I Os s

E

$ 10 1 l #,

1. #1098

[ #

k #1 9 S, 22 [  %, #1102 f , $, '1010 N #,$ ' [ #*, Y10 7 f Y# 9

'  % /

'[ O 109 [ k#  %

$ 1038 . o Y [ '4 $1 '# ' $/ 'No 3 0 s # "#10 7 0 0 g $ 977

$ 114 [ #s # 015 ' @ @ 2 1 8 0 063

1. ' 0 #

1133' - 1088 # #1 064 107 1 42 p " $, , y "I

$ $g p N0

o, $1116 @ # 1055'

$ 0' 3' '1 1 094

1. 1. 9 [ 9

[ $ 1096% h 127U N '

/ 046 923 [ $g '; 0g [ k '1140 ' O 924

1. Ts.',",vft;'+X+,,'l,#"rsfYs7 11 %5 8 h 1125D N 11 2 '

[ D, 1 9 '

0 1123

• 1 5U D

1. 14

/, 14Ng TEMPERATURES IN op

[pp $,\#1141D f 6 121%

1 l'igure 65 Core hitle 1)uct \litf wall lemperatures at 112" l lesatiim.

I 004. Accounting f or Uncertainties 1668 111 t

l t

b l

\-- 197 l

O g ASSEMBLY NUM8ER 308 313 a' 752 8 anna - ADIABATIC

,gg // 307 \ 801/ 312\ muun - WITH INTERASSEM8LY

\ / gg \ HE AT TR ANSF ER 302 805 306 313 311 820 841 793 305 310 826 842 802

>x / "X :::/ >"X :/ ;;;x

/ x =7 e x ,853/ 2;> x 5/ 2;;x , ,

45  :::: x,55/ 2 x 802 / ;;;\

8 ex;;;;f::2: o

2

/ 1118 1114 x 02 897 915

7 28x :::/ 20$x 81/ 21N di 1987 1082 2 838 848 206 787 794 209

( >)f :x 90y :,01x =f :12x 26x 2 / ;;; x

,08,7

,80 8:x 878 99

/ ,0,'l\

1984

/ ,,::

1998 67

\

1096 25

/ ::' \ "'/ "' 861 8

202 y9q 8" " 18 32 l54 3ag 826 851 60 MM 02 [ 1089 10 1086[109124 \ 826 15 gag 4 834 860 1083 1085 1065 850 1075 1083 34 g$g 6 1069 1082

( 795 ) 1055 g4 846 1060 1058 3 852 62 872 59 4 875 858 1048 e PL ANT THDV CONDITIONS 1047 841 820 eNOMINAL CONDITIONS 1040 29 855 128 329 e TEMPERATURES IN 'F 1056 830 00NLY VALUES WITH INT E R ASSEMBLY 93 1049 2 842 HEAT TRANSFER REPORTED FOR 1058 CONTROL ASSEM8 LIES 335 846 57 1050 835 69 846 400 804 Figure 66 Comparison of Assemblies Mised Mean Exit Temperatures at BOCl Under Adiabatic Conditions and including Inter-Assembly lleat Transfer Effects 1668-47 O

198

O 1

r ASSEMBLY NUMBER '

g an ADIA8ATIC mann WITH INTERASSEMBLY 921/ 307 \ 918 / 953 312 \ HEAT TRANSFER

' \ "93g /[301\,glll l;;\

956 gi

/ 9H \

930 [ 310 I

975 935 52 974 304 940 1004 982 I

[ 50 992 995 935 1005 51 gg4 ggg 203 940 2

• • 1 1003 1001 922 43 \

1019 \

,1

[ 45

/ 1019 1004 48 ggg 204 929 212\

995 s -

953 '

1018

[ 44

/ 1028 1019 46 1021 1000['

1002 28 ggg 205 g4g 213 1001 975 - '

908 -

1027 102 1018 47 27 gy4 206 209 1002 912 1022 1007 1003 1003 845 101 842 '

[ 1042 37 1019 1008 12 1004 26 \ ggi [ 207 860 4

1010 992 \ #

1021 939

]. 98 36 1009 1 1040

/

/ 1017 gg3 [ 201 \ 937 [ 208 1970 1018 1010 1002 ' 984 \ / 921 1067 99 1018 1009 67 1004 25 \ 982 / 202 \ 922 f 32 102g 100 13 1022 I

995 \ / 956 \

1028 1018 996 24 956 1021 1026 1018 1008 1005 1017 33 1021 to 15 1018 1040 1026 1020 1007 [ 999 1004[

1038 1021 61 10ig[ \ \998 1040 1028 / '

822 838 94 1941 3 1026 62 30 1018 1014 59 1038 1039 4 1070 [O PL ANT THOV CONDITIONS 1038 1067 / e NOMIN AL CON 0lil0NS

?916 / e TEMPERATURES IN 'F 1014 1037 29 128 1014 1016 00NLY VALUES WITH 1054 1009 INTERASSIM8LY HEAT 93 1053 2 100g TRANSFER REPORTED FOR 1925 1054 CONTROL ASSEM8 LIES 1024 1054 37 69 995 1025[

1974 998 Figure 67 Comparison of Assemblies Mised \ lean Exit Temperatures at LOC 4 Under Adiabatic Conditions and Induding Inter Awembly lleat Transfer Effects 1668-48 O

199

O

\ / 313 \

(308 307 312

\ / \

(' ' / o, \ '"'/ o,\ '"/ ,o \

~

3 3 3

\ / \ / 304

\ / 210 \

( 52 51 50 49 203 211 43 45 48 204 212 44 46 28 205 213 102 47 27 206 209 37 101 12 _ 26 ,207 98 36 68 13 201 208 99 67 25 202 32 100 11 14 24 33 60 10 15 34 61 94 3 62 30 59 4 l

29 12 93 2 57 69 I

l l

l icigure 68 Selecteil Assemlilies Anaipeti lhrougluiut I.ile willi Iltl' ION f or 1)nct Ililation anti llunttlell)nct intesaction Analyses )

i l l 166M 10.1 1

200 1

O u

/

2 NUMBER 2 13 36 3 14 21 44 4 15 22 30 38 45 51 4

5 10 16 23 31 39 46 52 57 11 17 24 32 f.7 18 25 33 41 48 54 59 26 34 42 60 l

35 43 l

H I

/ CORNER-TO-CORNER TRANSVERSE FLAT-TO-FLAT TRANSVERSE l l

l Figure 69 Crow Awembl.S Tranwerses where Ascrage Rml Claililine '

Temieratures liase fleen Calculateil by TRITON 166M-102 l

\,_) 201 l

. - _ __ _ _. h

i INNER BLANKET ASSEMBLY #102 AT BOC3 AXIAL POSITION (IN.) 1 2 ASSEMBLY FACE 3 4 5 6 9i' l

1 0.50 730 730 730 730 730 730 6.00 731 731 731 731 731 731 12.00 733 733 733 733 734 733 18.00 743 743 745 744 742 744 24.00 760 760 766 766 758 767 30.00 784 784 796 794 778 795 36.00 808 807 825 826 800 827 42.00 830 830 854 852 821 853 48.00 847 846 875 877 837 875 54.00 858 858 888 887 849 888 60.00 863 862 894 897 856 894 66.00 867 866 898 900 862 899 72.00 869 868 902 905 866 902 78.00 872 870 904 906 869 906 84.00 873 872 906 910 873 908 90.00 875 874 907 911 875 911 96.00 877 875 909 914 878 913 102.00 879 877 910 914 880 916 108.00 880 878 912 917 882 917 112.00 881 879 912 917 883 919 TEMPERATURES IN 8F INNER BLANKET ASSEMBLY #102 AT EOC3 AX1 AL ASSEMBLY FACE POSITION (IN.) 1 2 3 4 5 6 0.50 730 7",0 730 730 730 730 6.00 731 731 731 731 731 731 12.00 735 735 735 735 736 735 18.00 747 748 750 749 748 748 24.00 770 770 776 775 771 775 30.00 800 801 812 809 803 809 36.00 833 834 850 849 837 849 42.00 865 867 888 884 870 883 48.00 888 889 915 913 895 911 54.00 902 904 931 928 910 927 60.00 909 910 938 938 919 935 66.00 913 914 943 942 926 941 72.00 916 917 946 947 930 945 78.00 919 919 948 949 934 949 84.00 921 920 950 952 937 951 90.00 923 922 951 953 939 954 96.00 925 923 953 956 941 957 102.00 926 924 953 956 943 959 108.00 928 926 9 54 958 945 960 112.00 929 926 955 958 946 962 Figure 70 Typical Average Midwall Duet Temperatures Calculated by TRITON for Duet Dilation and Ilundle/ Duct Interaction Analyses 1668-100 202

?

t-LJ CORNER-TO-CORNER TRANSVERSE AXIAL PIN NUMBER POSITION (IN.) 35 34 33 32 31 30 29 28 27 0.50 731 731 731 731 731 731 731 730 730 12.00 745 \ 744 743 742 741 739 738 737 736 16.00 762 759 758 755 753 750 748 745 743 32.00 865 864 862 851 839 826 814 799 790 48.00 923 929 933 921 903 883 861 836 824 56.00 920 928 934 925 907 886 864 840 828 64.00 917 923 929 922 906 886 863 842 833 64.50 917 923 929 922 906 *15 863 841 833 RADIAL BLANKET ASSEMBLY #201 AT BOC2 TEMPERATURES IN 8F FLAT-TO-FLAT TRANSVERSE AXIAL PIN NUMBER POSITION (IN.) 50 42-49 41 32-40 31 22-30 21 13-20 12 0.50 731 731 731 731 731 731 731 731 731 N/ 12.00 743 742 742 741 741 740 739 739 738 16.00 757 756 755 754 753 751 750 748 747 32.00 851 852 851 845 839 832 825 814 807 48.00 912 916 919 912 903 891 878 358 846 56.00 914 918 921 916 907 895 a80 860 847 64.00 913 916 918 914 906 893 878 8 58 847 64.50 913 S15 918 914 906 893 878 858 847 Figure 71 Typical Average Rod Cladding Midwall Temperatures Calculated by TRITON -

for Duct Dilation and Bundle / Duct Interaction Analyses 1668 101 O .

203 -

1

/

O sen w

2e 1 C

• C l N

9 o

T.

n G

0 c2 A

- O

" b

=

.C C

M.

V

~.

1

- 2 t to 3 I

CC W 3

l O .e r

w .$

a 1

y-w  :-

2 U

S m

D N

=

C3 A m Eh 5E C

i c~

l I L' l 3 1 =

i b

b es b

I o

.?

W l i I i O O O O O O O N O CD CE3 #

w

(%) H3M0dlill:11VNIWON I668-2 204 i

l 1

1

i I

k 20 i

POWER TO INCIPIENT MELT, KW/FT 19 -

18 -

a U

a

s 3

u o

17 -

O 16

- O O

' ' ' I 15 15 16 17 18 19 20 OBSERVED l

l l

Figure 73 LIFE-3 Thermal Calibration 1681-2 i

l 205 )

l

CONTROL B00 NOS.

CALCULATE 0

- - - - - MEASURED 10 180ih

,,, k160 ih

- \

8 ,/ N 12

/

120lh \

\

100lh

/

I

/ '

r (1

\ I /

/

6 2 5 3 4

I Figure 74 Control-Rod Worths for EBR-il Run 27A 4

1681-4 O

206

=

g 5

u cc s

E Bu th P \\

i 't{,';[ \

~

S

1. I J

P

\

ll\\

ll a a

==

\

\ 8 t

/ _

' . e. =

4

\ E e

/ \ i s

w

=

1: = , 5 g.

( n I

\ W I O =

a

=

Il

\

g

\ $ $

a g

B B ll \ $

ll \

\ 9

/ \ \ =

n

/ \ -

s.

Il \ w

//

i i i i i

=

9 m *. n 0  ? a WM0d38 1681-3 207

$ 12 2

CAVITAT?ON NUMBER a =2 (P -P,) /(pV,/ 2)

WHERE: P =y ABSOLUTE PRESSURE IMMEDIATELY DOWNSTREA*A OF ORIFICE STACK 10 -

, = FLUID VAPOR PRESSURE P,

- V = AVERAGE ORIFICE HOLE VELOCITY

.. 8 e

cri w 8 - e ai = CAVITATION NUMBER AT THE OPERATING CONDITIONS WHERE INTERMITTENT C AVIT ATION BECOMES CONTINUOUS.

z _

i-

$a E6 e

e o -

n e

• e *

!, g 4 - e ,

g * .

e eeee

• ee .

• e e*e

• e e **

2 - e 0 3.6 1.6 2.0 4.4 2.8 3.2 0 0.4 0.8 1.2 PLATE SPACE TO HOLE DIAMETER RATIO Figure 76 Typical CRilRP Fuel Assembly Orifice Stack Casitation Data G G e

80 5 PLATES  !

70 - )

i 60 -

50 -

4 PLATES 40 -

3 PLATES

=

e

$ 30 -

2 PLATES w

5 u

O E a.

20 -

1 PLATE l

I l i I 10

.0 1.2 1A 1.6 1.8 2.0 ASSEMBLY FLOW RATE x 10-5 (Ib/ht)

Figure 77 Fuel Assembly inlet Norile-Orifice-Shield Pressure Drop for Orifice llule Velocity Limit of 40 Ft/See x

166M-57 209

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