ML20085E655
| ML20085E655 | |
| Person / Time | |
|---|---|
| Site: | Saxton File:GPU Nuclear icon.png |
| Issue date: | 03/31/1965 |
| From: | SAXTON NUCLEAR EXPERIMENTAL CORP. |
| To: | |
| Shared Package | |
| ML20083L048 | List:
|
| References | |
| FOIA-91-17 NUDOCS 9110210095 | |
| Download: ML20085E655 (123) | |
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1 SAFEGUARDS REPORT FOR THE SAX 7ON REACTOR PARTIAL
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PLUTONIUM COI:E II 1
i MARCH 1965
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TABIT OF CONTE!TPS ELEE.
I1 I.
I!TrBODUCTION.....
I-l A.
Objective and Scope..
I-l 3.
Program Description........
C.
Progrun Schedule......................
1-3 II.
NUC LE/Ji DESI GN.........................
11-1 A.
Introduction........................
11-1 B.
Re a c t ivity Summary.....................
II-6 5
II-ll
.atrols Summary.
C.
D.
4.netic Characteristice..........
II-12 III.
COBE 10fDRAULIC AND THERMAL DESIGN................
III-1 III-l h
A.
General..........................
B.
Coolant Flov............
III-2 C.
Variation of Primary Dystem Temperature and Prescure....
III-2 L.
Engineering Hot Channel Factors..............
III-3 III-5 E.
Daparture from Nucleata Boiling.
F.
Hydraulic and Thermal Design Ihrameters..........
III-12 G.
Central Temperature of the Hot Ibliet........... -III-15 IV.
MECHANICAL DESIGN......
IV-1 IV-1 A.
Core Icading.
B.
Fuel Assembly De sign....................
IV-3 C.
Fuel Rod De s i gn......................
IV b D.
Justification for Re-use of Control Rod Followers and L Assemblies.
IV-10 V.
INSTRUMENTATION.........................
V-1 A.
In-Core Inntrumentation.................. V-1 B.
Plant Site Monitoring.
V-2 VI.
ACCIDENT ANALYSIS........................ VI-l A.
General.
VI-l B.
Reactivity Accidents.................... VI-3 C.
Mechanical Accidents.................... VI-17 D.
Muimum Hypothetical Accident..
VI-20 VII.
SAFETY CONSIDERATIONS...................... VII-l A.
Justification for Inclusion of 9x9 Assemblies of Vibrationally Cogacted Fuel in Saxton.........
VII-l B.
Operation with Defect ive Fuel............... VII-5 VIII.
CONCLUSIONS, VIII-l l
1
_-__-_i
LIST OF TABL.E_S.
Table No.
Title page 1
7 Pressure and Void Coefficiento............... 11-10 d2 Isatopic Power and Neutron F2nctions, Delayed 11-12 l
Neut x Data,/f,ff and Proept Lifetime....
III.1 Engineering Hot channel Factors..
........... III-4 III-2 It/drauliu and Thermal Design Parameters.......... III+12 IV-1 Fuel Assembly Types.................... IV-2 IV-2 Pluton!.um Fuci Rod..................... IV-$
IV-3 Core II Fuel-Rod Dimensions................ IV-6 IV b Maximum Allovable Gas and Vapor Content for Flut, onium Fuel Mixturco............... IV-11 IV-5 Maximum Pressure Stress Plus Thermal Stress IV-12 in Tuel Cind at Beginning of Core Life.
IV-6 Control Rod Follouer and L Assembly B,irnup Analysis....
IV-13 ii
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LIST OF FICUFIS Figure No.
Title II-l Inctalled Reactivity vs Lifetime for the Saxton Plutonium Core II-2 Saxton Plutonium Core Power Dict ribution 11-3 Center Assembly Power Distribution 11-4 (a)
Radial Nuclear Hot Channel Factor vs Hours Operation at 27., MJt II-L (b)
Operating Sequence - Maximum Thermal Power vs Months Operation for a 16 kv/ft Linear Power Limitation 11-5 Saxton Temperature vs Surface Heat Flux of Pellet II-6 Saxton Doppler Coefficient vs Effective Fuel Temperature II-7 Saxton Power Coefficient vc Power Level II-B Variation of Moderator Temperature Coeff5 cient vith Temperature II-9 Reactivity Worth of Boron vs Boron Concentration 11-10 Stuck Rod Power Distribution II-11 Dropped Rod Power Distribution III-l Comparison of H-D!B CorreIntion with Measured Data in Quality Region (p = 800 to 2750 psia)
III-2 Co:::parison of H-DIG Correlation with Measured Data in Quality Region (p = 2000 poi)
III-3 Comparison of q"-DNB Correlation with Measured Data in Subcooled Region (p = 800 to 2750 psia)
III-4 H-DIG Probability Curve at 2000 psia III-5 q-DNB Probability Curve at 2000 pain 111-6 Stable Film Boiling Heat Transfer lata and Correlation III-7 Thera 11 Conductivity of ' Uranium Dioxide III-8 Thermal Conductivity of Vibmtionally Compacted Fuel IV-1 Sarton Grid Design V-1 Saxton In-Core Instrumentation i
iii l.
LIST OF FIGURES (C.tt'd) he ac N.
Thie VI -E -1 Cold Startup Incident:
Power Response 2 5 x 10'b bk/see Int.crt:c.
V' -M C: _$ Sur'.p Incident :
Power Response 2 5 x lO*b bk/see Insertien VI-B-3 Cold Startup Incident: Average Fuel, Clad agd )later Tereperature Responses, 2 5 x 10- Uk/seeInsertion VI-P '.
Cold St aMup Incident' HotSpotHeatFluxResponse,25x10*kbk/see Insertion V1-B-5 Hot Star- -tp incident:
Power Response 2 5 x 10 bk/see Insertion VI-B-6 Hot Startup Incident Power Response 2 5 x 10O bk/see Insertien V -B-7 Ilot Startup Incident: Average Fuel, Clad andJiater Temperature Responses, 2 5 x 10'4 Uk/seeInst.rtion VI-B-S Hr.t Startup Incident:
HotSpotHeatFluxResponse,25x10"Nbk/sec Insertion V;-P-9 Crett.aous Rod Withdrawal:
Nuclear Flux, H.at Spot Hest Flux and Coolant Pressure Responses, 2 5 x 10"b bk/sc Insertion VI-? -J, S*es Break Ae:ident:
Prit:ary System Pressure Response VI-B-L.
Steam Break Accident:
Steam Flow, Neutron Flux and Hot Spot Heat Flux Responses VI B c.
Steam Break Accident:
Steam Generator Inlet and Outlet and Core Inlet and Outlet Temperature Resporses VI-F-13 Steam Break Accident:
Negative React (vity Insertion Required to Msintain C.5% Uk Shu ;devn vs B' On Cencentratir.
V.1-C-1 Primary Oc 0; Ant Fiov - Coastdown Following Loss of Pu:np Power VI-C-2 Less of Flow Accident DNB Ratios versus Time V!-D-1 In.tantane: ras Release of Main Ccolant l
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I.
I!TTRODUCTION d
A.
OBJECTIVE AND SCOIT The objectives of the Sexton Plutonium Project are to develop infor-mation eencerning the utilization of plutonium enriched fuels in a closed eyele water reactor environmen*, and to develop analytical methods and technfques that vill reliably predict the in-pfle performance and long term behavior of such fuels in large scale i
r reactorn.
The scope of the program covers the design and fabrication of mixed PuO -UO fuels, evaluation of in-core performance of these fuels and 2
2 post irrsdiation examination and evaluation of fuel and clad samples.
These fuels vill make up part of the core in the Saxton reactor with the remainder of the core made up of UO fuel.
The scope also includes 2
pre-irradiation criticals to evaluate the predicted nuclear design.
B.
PROOR/M DESCRIPfION The progrom goals of extended irradiation exposure and high fuel burnup plus the nuclear and control characteristics of the Saxton reactor are the basic parameters which have been considered in the design of the fuel for the plutonium core.
The selection and analysis of various combinations of critical factors assures a fuel design which vill not only meet the goals of the program but vill also be within the operational limits of the Saxton reactor.
Based on the fuel design and enrichment established, the thermal j.
. and hydraulic characteristics of the core have been analyzed and l
evaluated.
In addition, a set of critical experiments have been designed to verify the nuclear paraueters of the fuel specified i
l I-l l
l
during the design pheme of the program.
The results of these critien1s, expected by May 1%5, vill be analyzed and coepared with predicted values from the design study before the partial plutonium core is operated.
Additional tests prior to operation at full power in the Caxton reactor include a series of zero and lov power tests and physics measurements to obtain and verify actual in-core characteristics and performance. The infomation to be cbtained from these tests a
and measurements include core excess reactivity, control rod and soluble poison vorth characteristics, and temperature, pressure, 4
moderator and power reactivity coefficients.
These nuclear characteristics as well as the core thermal and hydraulic charac-teristics and neutron flux and power distributions vill be measured and evaluated at intermediate power levels before operating at full power.
Present Saxton in-core instrumentation consisting of pitot tubes, themoccuples and flux vires vill be used in obtaining the data.
Daring the course of the project, a re-determination of there charceteristics vill.be made during several shutdown periods in order to detemine characteristic changes with irradiation exposure.
Comparisons vill be made between measured values and those predicted by analytical methods.
Following the in-core irradiation period, an extensive post-irradi-atien examination program vill be conducted to evaluate fuel per-formance and to co= pare actual and predicted performance.
Dcamina-tions vill be made for signs of distortion or year marks on fuel assembly cans and for broken or distorted grid structures. Fuel rods vill be examined for abnomal appearance, vear and dimensional changes.
Selected high burnup rods vill be examined with a stereo-tieroscope for evidence of cracking, distortion or other abnormal appearances.
I-2 5
___i______. _ _ _ _ _ _ _. _ _ _ _ _ _
i s
Some rodo will be cectioned and examined metallographice uy to evaluate cladding perfornance, fuel cracking and radial redistri-bution. Some selected high burnup rods vill be punctured so that gao pressure and clad stress can be calculated and a quantitative analysis performed by mass spectroccopy to determine fission Eas releases.
Samples of cladding from these rods vill be taken for metallographic and raechanical testing.
Ib11ets from several rods vill be sampled and examined using mass spectrographic techniques to determine the change in U and Pu isotope concentrations as a 4
function of burnup.
Radiochemical analyses vill also be performed for selected fission products to determine burnup and radial iis-tribution.
C.
PROGRAM SCHEDULE
\\
A tentative schedule for installation-of Core II is as follovst s
3/18/65 Begin critical experiments 6/4/65 Complete critical experiments 7/10/65 Start receipt of fuel at Sexton 7/1/65 Complete fabrication oi' all fuel assemblies 8/27/65 Core loading completed Iov po w testing period - 9/ 1/65 - 10/1/65 I
I 1
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II.
NUCLEAR DESI0tj A.
I!CTIODUCTION 1.
Ob.jectives As outlined in Section I, the purpose of the Saxton Plutonium Program is to develop infomation concerning the use of plutonium enriched fuel in vuter reactor systems.
This purpose is to De fuel in the Saxton reactor achieved by the irradiation of Pu0 -UO2 p
and by a supporting program of analysis, experiment and evaluation.
The core, composed in part of plutonium fuel, is designed for 8250 megawatt days of operation.
Since it is desired to achieve a high burnup in the plutonium fuel, the reference design is one in which nine plutonium fuel assemblies vill be installed in the center of the core with the twelve tranium fuel assemblies installed in the peripheral locations.
2.
Nuclear Characteristics Summary When installed in the center of the reactor, the plutonium assemblies influence the nuclear characteristics of the system to a greater extent than if they were installed in peripheral locations.
- However, the analysis has shown comparatively small changes in reactivity and kinetics characteristics (as compared to an all uranium core) are introduced by the plutonium.
The following qualitative statements briefly summarize the major reactivity and kinetic effects that occur with a partial p,1u'tonium core.
a) The Sarton plutonium core vill have a more negative moderator temperature coefficient than with a conventional uranium loading.
II-1
b) The negative Doppler coefficient vill be larger with PuO 'UO 2
2 fuel than with.)unt UD I"*1' 2
c) The part-plutonium core vill have a lt.rger positive pressure coefficient and a more negative void coefficient than the conventior.a1 uranium core.
d) Boron worth and control rod worth are decreased in the partial plutonium core.
e)Thedelayedneutronfraction([ff)andpromptneutronlifetime when averaged over the core at the beginning of life vill be smaller with a partial plutonium loading than with a full loading of uranium fuel.
They both remain essentially constant through-out the core life because of the-change in radial power distri.
bution and the Pu buildup in the outer tiranium assemblies.
As the central plutonium region has a larger absorption cross section than the outer uranium region, the thermal flux undergoes a marked change at region boundaries and peaks to a greater extent than normal in those water slots near the plutonium.
In addition, since the fission cross section is larger, local power peaking occurs in the plutonium rods at the boundaries s.nd water slots.
Diffusion theory analysis has been used in determ'ning power distribution and power peaking effects.
The critical experiments as outlined in paragraph A-4 vill determine if the available reactivity and power distri-butions vill permit installation of the plutonium assemblies in the center of the core or if it is necessary to change the reference design to one in which the plutonium assemblies are installed in peripheral locations.
Such a change vould minimite the differences between a core cLntaining plutonium and a conventional Saxton core.
II-2 j
3 Analytical Methods The nuclear design of the plutonium loading in the Saxton reactor was based on th* use of standard analftic methods developed in the design and operational analyeis of a number of pressurized water reactors.
These nethods have been cceptred to a large number of water moderated critical experiments. They were also used in the design of the initial uranium loading in the Saxton reactor (Saxton Core I) and in the subsequent ecm:parison of analycis with experimental data from Saxton ole ration.
In addition, the methods were used in the analysis of six mixed-oxide (Pu0pV0 ) critical and approach-to-critical experiments p
conducted at Hanford and an allowance for the discrepancy between analysis and experiment was included in the reactivity calculations for the plutonium core. Thus the analysis is based on proven practice, the reactor is one for which experimental information is available, and the extraplation to a part-plutonium configuration is based, in to far as possible, on applicable critical experiments.
i The following paragraphs contain a brief description of the computer programs used in the analysis.
LEOPARD (1}
The LEOPARD co=puter program determines fast and thermal neutron spectra based on a modified MUPI-SOTOCATE model.
The thermal spectrum is the same as that given by a Wigner-Wilkins SCFOCATE calculation except for the treatment of disadvantage factors.
Disadvantage factors are determined by using a modified form of the Amuayal-Dencist calculation at 172 energy levels from tero to 0.625 ev.
LEOPARD computes a non-thermal spectrum based on a consistent B-1 MUPT IV calculation. The resonance integral for U-238 is determined frce a correlation that is in good agree-ment with Hellstrand'c veasure.ments for uranium metal and urunium (1)
- 8. F. Barry, " LEOPARD - A Spectrum Dependent, Non-Spatial Depletion CodefortheIIM-7094,"WCAP-3741(1963 11-3
dioxide at any temperature.
The calculational procedure contained in LT:OPARD has been compared 4
to 116 critical and exponential lattices, $$ using UO fuel and 2
61 using uraniu= metal fu n.
The calculations result in an average k of 0 9931 t 0.0086 for the 116 cases where the quoted gf errors correspond to one standard deviation.(1 The code also computes Cero-dimensional fuel depletion effects and providen a tiroe dependent microscopic cross-section library for subsequent spatial burnup calculations.
The burnup portion of the LEOPARD code has been compared with Yankee Core I spent core data.
The plutonium isotopic composition as a function of U-235 depletion is in good agreement with the data.(2)
PDQ-}
Solves the fev-group, time independent, neutron diffusion equations in X-Y geometry.
AIM-5 Solves the fev-group, time independent, neutron diffusion equai ton in one dimension.
TURBO
- Solves the few-group, two dimensional (X-Y geometry) neutron diffusion equations in combination with a point-vise burnun calculation to determine reactivity-lifetime relationships The microscopic library generated in LEOPARD is used to determine time dependent group constants for use in TURBO *.
(1)
L. E. Strawbridge, " Calculation of lattice Parameters and Criticality for Unifor=WaterModeratedLattices,"WCAP.37h?(1963).
(2)
"L rge Closed Cycle Water Beactor Research and Developant Program a
Progress Report for the Period July 1,1964 to September 30, 1964" t
WCAP-3269-5 l
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+=y vy---s-yv&,,wwi7ew7-6W---
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LUX __
/s modification of the CANDLE one-dimencional few group depletion code and is used to determine reactivity as a funetton of lifetime.
The LEOPARD microscopic library is used.
THUNOS A cell transport theory code in both space and energy.
This prog am was used as a check on the thermal group calculation contained in LEOPARD.
The results from the two calculations were in excellent agreement.
h.
Experiment Sequence Prior to the operation of the plutonium fuel in the Saxton reactor, a critical experiment program with both the PuOg UO2 #"'I # 08 ""d the UO fuel r de vill be carried out at the Westinghouse Reactor 2
EvaluationCenter(WREC).
The proposed program consists of the fellowing six basic configurations:
Type Configuration Type Core Type Puel Lattice 1
Single-region, clean Pu0gUO2 b 08" core 2
Single-region, clean Pu0 -UO DesignH/Pu p
2 core 3
Two-region, clean Pu0gUO,Inside Design H/Pu core UO Outhide 2
4 Single-region, Pu0 'U0 DesignH/Pu 2
2 borated core 5
Two-region, borated Fuo -UO Inside DesignH/Pu core UO Out$1de 2
2 6
Inverted, two-region, UO Inside DesignH/Pu 2
clean and borated Puo -UO Outside cores 2
2 As shown by the table, single-region criticals with the plutonium rods and two-region criticals concisting of separate plutonium and uranium II-5 l
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-.r-v.-
s
.-. -,...-y,.,e
,,-.,.~.-r.
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eer--
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zones vill be carried out.
The experimente vill include the measurement of cyatem renetivity,,'ver distribution and power peaking effecto, flux dist ributions, control rod and boron vorth, and kinetic parameters.
Tnese experiments vill provide a " clean configuration" check on the enslytic methode used in the design of the fuel and reference core.
Configuration 6 may not be checked if the first 5 closely verify the analytical results.
later, low power testo conducted in the Saxton pressure vescel vill provide an additional check of the design prior to power operation.
There tests vill follow procedurec developed in the course of previous Saxton stsrtups.
The lov power tests vill include such measurements as the ruoderator teeperature coefficient at several boron concentratione, boron vorth, control rod differential vorth in a normal withdrawal sequence, and flux distributions.
B.
REACTIVITY SGWJiY 1.
Available Reactivity The excess reactivity of the system was determined by means of FDQ-3 two group diffusion calculations in X-Y geomet.y.
The hot, clean reactivity at power is expected to be 103+135Ax/x.
This exceso reactivity estimate includes an allowance for the discrepancy between analysis and experiment that is based on the analysis of six variable lattice, mixed-oxide critical and approach-to-critical experiments conducted at Hanford.
In this comparison, an average analytteal over-prediction of 2.6% A x/x vas found. The listea uncertainty or t 13% Ax/x is an estimate as to the size of the possible error that may occur in initial excess reactivity.
To provide a check on the methods of analysis used in determining-the system reactivity, similar PDQ-3 calculations were carried out for II-6
Saxton Lore I.
A calculated k of 0 9997 was determined for the g7 hot, rods out, just critical configuration containing 180k ppm boron.
Calculations were also cade to determine the clean Core I reactivity The calculated reactivity,1k.825Ak/x, is in good at power.
agreement with the reactivity as determined from boron vorth and power coefficient measurements, Ih.4%bk/k.
2.
Reactivity Lifetime Expected The available lifetime for the Saxton reactor containing a partial core of plutonium fuel was determined by both one. dimensional (radial) and two-dimensional (x-y) burnup calculations.
The reactivity as a function of lifetime as determined from a LUX radial calculation is shown in Figure II-1.
This figure also chove the possible variation in lifetime that vould result for the maximum and minimtun orpected values of initial reactivity, t
Duplicate lifetime calculations using LUX vere carried out for Saxton Core I and vere compared to the projected lifetime based on the measured depletion rate.
In this comparison, the calculated lifetime is about 1000 hours0.0116 days <br />0.278 hours <br />0.00165 weeks <br />3.805e-4 months <br /> less than that indicated by the experiment.
3 Lifetime Power and Burnup The peak power in the core vill occur in the plutonium region at the beginning of life.
Figure II-2 shows the core radial power distribution ac calculated by a PDQ-3 two-dimensional diffusion analysis.
The analysis was originally carried out for a plutonirn enrichment of 6.5w/oPu0.
The difference in power levels for i,he new reference 2
designof6.6v/oenrichmentandthe6.5v/oenrictmentdesignis small, approximately 1% in local power.
The local to. core average power in eacn fuel rod in the central plutonium assar.bly is shown in Figure II-3
(
II-7
Tne follewing list sunnarises the maxitum vaiutre of pear. red radini pcuer tc ecre average radial power fer bot h the plutenlum and urani :
regions na a function of time. The values were det ermined by meant of two-direncienni burnup calculations using TURID*.
PMK R3D TO CORE AVEMGE IMD.T.AL P0h'ER Tine. H:u r t, at ??.5 ISit Po Rechn U Rarien 0
2.36 1.22 6, COO 1.68 1.26 12,000 1,63 1.23 Average for 0250 U.ID 2.00 2.24 Tht-initial peak power values in this list and in Figure II-4(a) are higher than those shown in Figures II-2, II-3, in ceder to cerrect the calculated peak value to an expected experimental peak value based on a comparison of analytical and experinental results from the cricinal uranium loading. The initial peak power also includes correc-tions for an increase in coolant boron content to corres},ord to beginning of life conditions and for the extrapolation frca 6.5 w/o enrichment to 6.6 w/c enrichment.
To prcvide the desired lifetime of 8250 )MD, the specified enrichment is6.6u/oPu02 and results in an increase in the nuclear het channel factor over that for a conventional all uranium Saxton core. As a recult, it may be necessary to limit the reactor power level initially in order te avoid exceeding the linear power density design of 16 Kw/f t.
F.icure II-l.(a) ebowe the radial nuclear hot channel factor in the plutenium region as a function of operating time at 23.$ int.
A react:r cperating sequenee that does not ex:eed 16 kw/ft is shown in Figure II-L(b). This figure shows that a power level of 23.5 iMt will be allowed following six months of operation at 21.6 )Mt and a load fact:r c f 0. 5.
The actual initial power level limit to ensure that the design of 16 Kw/f t will not be exceeded will be determined following experimental neasurements and calculation of het channel factors.
F Thercaf ter, the maxinum linear power density decreases with the reactor a*. a constant pcwer level of 23.5 }Mt and II-B
dr, expected to be approximate 3y 13 0 kv/ft af ter 82501&D of operation.
A design lifetime objective of 8250 KG and the listec averaga value for the peak rod pover ir. each region results in the following predicted peak rod everage burnup values,.
20,000 }&D/ tonne 14t.ximum burnup, Pu fuel rod
=
11,500!sD/ tonne Maximum burnup, U fuel rod
=
4.
F.eactivity Coefficients 4.1 Doppler Coefficient Doppler coefficient calculations vere carried out as a function of the effective fuel temperature where the effective fuel temperature it, defined as that temperature which gives the correct experimental power coefficients of reactivity when employed in standard design calculations. The relationehip between pellet temperature and effective fuel tempert,thre is based on the work vhich correlated effective fuel temperature for Doppler broaden-ing of U-238 resonances with experimen'.a1 power coefficient measurements in the Yankee, Saxton, an:t 3R-3 reactors. The temperature power relationships developed for the Se.xton reactor are shown in Figura II-5 The Doppler coefficient was determined by completing a series of LEOPARD claculations for values for T,ff ranging from 800*F to 2000*F in 200-degree increments.
The resulting group constants were then used in AIM-5 radial diffusior. calculations. Figure II-6 su=marizes the results for variations in T,ff alone and for the variation in temperatures shown in Figure II-5 The difference between the two curves 18 due to the-change in moderator cor. tent resulting from the small change in clad dimensions with temperature.
Figure II-7 shows the derived power coefficient based on the Doppler calculations and the temperature relationships of Figure II-5.
II-9 c4
________-_.-_------.._-..+-J
Calculations vere also made to determine the fraction of the
. total Doppler effect due to temperature changes in each region separately. Approximately 70% of the reactivity change due to Doppler is the result of temperature changes in the central plutonium region while 30% is the result of temperature changes in the outer uranium region.
4.2 Moderator Coefficient Tne moderator temperature coefficient was determined by a sequence of LEOPARD and AIM 5 calculatione.
Figure II-8 summarizes the results for partial plutonium core.
The figure also incluies a comparison of the measured values of the moderator temperature coefficient with the analysis for the conventional Saxton all uranium loading, k.3 Pressure and Void Coefficient i
Table IT-1 compares the pressure and void coefficients for the conventional uranium loadin; vitt that of the core containing i
nine f.lutonium fuel assemblies.
L'he analysis was carried out using a LEOPARD and AIM-5 sequence Mermures ci 2eu cf the pressure coeffinent at various boron concentrations are Aneluded in the table. The analysis and experiment are in good agreement for comparable conditions, TABLE II-l PRESSURE AND VOID COEFFICIE!rTS 6
f
( A x/x / psi) x 10, g,,,,,,,,, 330 7 Pressure Coefficient:
Pu Core CONVENTIONAL CORS 1000 ppm 1000 ppm Clean Boron Maximum Boron Boron 0 Boron Redded, O Boron
+3 5 Analysis:
+1 30 (2000 ppm)
+2.1
+3 1
+5 6 (all rods)
Erperiment:
+1.k5(1600 ppm)
+2.4
+4.6 (partial rods) i II-20
l f
(% Ak/k/n void), Moderator = 530 F Void Coefficient:
Pu Core COWENTIONAL CORE 1000 p;c 2000 ppm 1000 ppm Clean Fully Rodded Boron
, Baron Baron O lbron O Boron
-0.27
-0.10
-0,18
-0.26
-0.43 k.h Bottn Worth The LEOPARD - AIM-5 calculated boron vorth rs a function of boron concentration is shown in Figure II-9 C.
CONTROIS
SUMMARY
The nominal mode of control for the Saxton Plutonium Core II is expected to be chemical chim.
However, at times during the program, it may be g
nececcary to employ control rods alone.
In any mode of control, soluble poison vill ta required for cold, clean core shutdown.
1.
Benetivity Effects Control rod worths were established by means of a series of PDQ calcu-lations using a method that accurately detemines rod worth in Saxton Core I.
The calculated control rod total bank vork for the plutonium core is 16 9% Ak/k. With a maximum predicted initial hot, clean at j
power excess reactivity of 11.6% Ak/k, the minimum initial available j
dhutdown margin in the hot reactor is 5 3% Ak/k. stuck rod calcu-lations show that the total bank vorth with the most effective rod stuck in the fully withdrawn position is 117% Ak/k.
The minimum shutdovn margin for this condition at the beginning of core life is
'therefore only 0.1% Ak/k.
Careful evaluation of the critical tests vill detemine the amount of initial excess reactivity available with the new core.
If the initial excess reactivity is greater than 9 7% A h/k, then the shutdown criterion of 2% A k/k shutdown margin availsble with the most reactive rod stuck II-ll d
.q
Vill not te net, and it vill be necessary to do one or both of the following at the beginning of core life.
a) Pertrict rod withdrawal initially by cetting existing limit evitches to hold the required reactivity in the core no was practiced with the conventional all uranium loading.
b) Ure chenical shim for other rea;tivity effects such ac xenon in u.idition to that required for the temperature defect.
5 2.
i'o.:er ?ffect Figures 11-10 and II-11 chov the effect on the beginning of life power distributions for stuck rod and dropped rod conditions. The peat rod poucr in either figure will be about 3$ less for zero bcron cencentration.
D.
17TfTIC CHARACTERISTICS
[,7 was evaluated throughout life from a TURBO
- calculation.
A constant value of rv' O.00L9 was found.
Tr.e ggf was determined by weighting the delayed group vields of the various fissioning materials by the fraction of neutrons from each icotope.
An 1:tportance factor, derived from LEDTARD calculations using different source spectra, vac cpplied to account for the dif.'erence in importance of delayed and prompt neutrons.
Table 11-2 sammarizec the power and neutron cource fraction in each of the two regions for each isotope at the beginning of life.
The table also includes delayed neutron data and the region and core average [df and prompt neutron lifetime.
TAE.E II-2 IS01'OPIC PO'n"ER AND IEJTRON FRACTIONS, DELAYED NEUTRON DATA, /
AND PR34PT LIFET1ME d'f Beginning of Life Icotopic Power Fractions Neutron Sour _ce Fractions Region U-235 U-238 Pu-239 Pu-2h1 U-235 U-238 Pu-239 Pu-2h1 Ccre Average 0.4301 0.050h 0 5078 0.0112 0.4008 0.0514 0 5352 0.0125 Pu Region 0.1302 0.0503 0 7998 0.0177 0.1188 0.0490 0.8132 0.0190 U Region (
0 9390 0.0520 0.0088 0.0001 0.9340 0.056 0.0099 0.0001 i
II-12
- _ _ _ _ _ _. _. -. ~.. _ _ _ _ _. _.
Delayed deutron Inta
-1 relayed i'
1 g
Neutron Group, i U-235 Pu-239 U-235 U-236 Pu-239 Pu-241 1
0.012k O.0128 0.000215 O.000201 0.000074 0.000091 2
0.0305 O.0301 0.00142h 0.002151 0.000626 0.000775 3
0.111 0.12h 0.001274 0.0025h3 0.000443 0.0005k9 L
0 301 C 325 0.002568 0.006092 0.n00685 0.000848 5
1.13 1.12 0.000748 0.003533 0.000181 0.00022h 6
3 00 2.69 0.000273 0.001178 0.000092 0.000114, 0.0065 0.0157 0.0021 0.0026
/,ff and Promut Neutron Lifetime Pecion gg Lifetime, psee Core Average 0.0049 11 7(5)
Pu Eegion 0.0035( }
8.6' U Begion 0.0075 15 7 (1) Includes particily burned uranium fuel followers.
(2) Includes partially burned uranium L-sections.
(3) Relative abundance for Pu-239 used.
(4) Calculation made with core containinE 1000 ppm boron.
(5) Weighting based on power fractions of U =.439, containe L-sections and followers, Pu =.561 II-13 1
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PDQ Potte. Distribution, Configuration: 9 Pu0 -UO Assemblies (6.5 v/o Pu0 ) in Center Positions 2 2 2 No rodo,1000 ppm Boron, Depleted Fo11 overs and L-Sections A = 1.0SB4B B C D E F I I I 0.6h3 0 79h 0.6h5 l 1 q' -, I l ll6,g.118 l , __ _ I < 1.6hT '1.6061 'r r O.619 1.192 1 1.L3h 1.2h8 O.6hh 2 ~ h 2 i, _L y 1.982, J.9Bh F, q 1.616,1.125 q '8 71 1 978' '2.033 7-1.656' 1.155 4, ~ 0.7h1 7 1.k18 1.718 1 1.h50 0.802 L,w hh 5 2.16 3 J 1.072,M[3 e w itw2.058,<P.010 g' i1 1 s, 2 , -- q, r - i, 1.0h9 1.520 2.007 2.013 l u w 4 0.6?8 M 1.h6h 0.66h 1.225 6 1.52 J.62h j, i s a st 1.100 '1.131 I 'I 0.660 5 O.66h 9 0,818 w w ,w i The underlined value in each assembly is the average power in that assembly re'ative to the average.pover in the core. The relative power.is also shown for individual fuel rods near the uncontrolled corners of assemblies (where hot spots usually occur). Maximum Rod Pover = 2 36. (This value includes a correction for the discrepancy between analysis and experiment for Saxton Core I and for an increase in boron content to 2h00 ppm). Saxton-Plutonium Core Power Distribution v r "RE i;.-
l Conditions : 9 Pu0 -UO Assemblies (6.5 v/o Puo ) i" e'"t'# 2 2 2 No Rods,1000 ppm Boron Depleted Followers and 1,-Sections 1 = 1.088kB 4 [ h 2.03 1.85 1.88 2.1h 2.1h 1.91 1.y0 2.10 1.8h 1.66 1.66 1 76 1.89 1,76 1;67 1.68 1 90 1 o7 1.66 1.63 1.6k 1.66 1.6h 1.63 1.67 1 92 2.15 1.T7 ' 1.6k 1.62 1.62 1.63 1.65 1 77 2.14 1 94 1.66 1.63 1.62 1.63 1.66 1 90 19h 1.673 1.63 1.63 1.63 1.65 1 77 2.15 1 93 1 71 1.68 1.67 1.65 1.6h 1.68 1.89 2.00 1 9h 1 95 1 95 1 78 1,67 1.68 1.87 \\ 2.16 1.89 1 96 2.06 / Maximum Rod Power = 2.36 I' (This value includes a correction for the discrepancy between analysis and experiment for Saxton Core I and for an increase in boron content to 2h00 ppm). Center Assembly Power Distribution FIGURE II-3 l
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500 1000 1500 2000 Boron Concentration, ppm
PDQ Power Distribution Confi6uration: 9 pug 'UO Assemblies (6.5 v/o Puo ) I" **"*'# 2 2 2 i l positions. Bods 1,2,3,b,6 in, 1000 ppm boron, I Depleted Fo11overc and L-sections B C D E F c >I I I y ;.. _> i 1 l :r ,lr t l I 1 g 0.L3 O_.18 d 0.62 .l I 1 31 0.96,<0.96 l sif , 1 36l.36 I L-M r It I, i a H, 1.00 s. 0.62_ 2 0.L8 0.05 1.11 e m q $1 h 2 i. 5 7 ] 1.89,1 93 1.37,p.97 i(. -4j ( "{ -)I 2.0f 5.07 '{- JI 1 36'0 96 l u
- i 1 77 1.11 L
0.50 -3 0.88 1.81 s 3 -- ) i 1.60,J.Vu 31 2.10,<1 95 4 --3 i y 2.071 92 1.65'#2.37 q I 1.9h 2.86 0.98 0.LS .4 1.12 y 6 h. ic 2.LS,3.b1 7 ,s Fc- } Of 1 73'1.69 'i l 1.00 0 0.52 5 1.18 C ~ u:. A mi K i w Tne underlined value.in each assembly is the average power in that assembly rela-tive to the average power in the core. The relative power is-also shown for in-cividual fuel roca near the uncont. rolled corners of assemblies (where hot apote usually occur). The peak power includes a correction-for the discrepancy between l-analysis and experiment from Saxton Core I. l Stuck Rod Power Distribution H GUF3"I:-10'y g 2 .,.. _.. ~..
PDQ Power Distribution Configuration: 9 Pu0 -00 Asser.blies (6.5 v/o Puo ) in Center Positions, Rod 6 in, 2 2 2 1000 ppm Boron Depletec Followers and L-Sections B C D E F l l 1 ,i 7 ,; g _..; i l .i L 0.76 0 9h d 0 76 l ~ 1 - )i 1 31,<1.26 r I<- P q , 1,831,83 I t __ _ r {. i 6 0.72 1.37 h 1.63 .6.hD h 0.73
- - 2
~ ~ i 4 2.h5 2 Q" 2 3 <2.18 6 g 1.7h, Q _j ( 'I 2.16' '2.22 'I 17f"1.22 [ l 0.82 .J 1 55 1 79 1.h9 9 0.8h 3 e i 5 3 [---)l 1.1h,<1.63 31 1.93,<1.88 1 y 4 3 I[ 1.10 Y 1.821.77 I' [' 1 59 y 1.09 2 0.61 C O.66 1.17_ h 1.1h ~ ~ q m 6 h 1.36,1 35 ir kt ic-- ss ic- [ _..D 1 0 93' 5.92
- 'f
) 0.62 0.3h 5 k 0.60 q 94 A _ 31 2 The underlined vslue in each assembly is the averer,e power in that assembly rela-tive to the average power in the core. Tne relative power is also shown for in-dividual fuel rods near the uncontrolled corners of assemblies (where hot spots usually occur). The peak power includes a correct. ion for the discrepancy between i analysis and experiment from Saxton Core I. FIGURE II-11 Dropped Rod Power Distribution
_ _.. _ _ _. ~ _ _ _ - - _. _ _ III. CORE HYDRAULIC AND THDMAL DESIGN A. GENER/1 Introduction The power level at which the reactor may be operated without core I damage depends on the ability to remove heat from the core without exceeding limiting clad and fuel temperatures. The hydraulic and thermal design of the core cust provide sufficient margin so that the appropriate safety limits as defined belov are not exceeded during steady state operation and reactor transients. The maximum te=peratures are evaluated by using the concept of hot channel factor, i.e., the ratio of maximum to average conditions within the core. For the evaluation of this ratio, neutron flux distribution, coolant flow distribution and the core fabrication tolerances are considered. Thermal Desien Safety Criteria The thermal output of Core II has been established on the basis of the following ground rules: No bulk boiling is allowed in the hot channel during normal steady state operation even with the core inlet temperature at the upper limit of the control deadband with maximum nuclear instrumentation error, and with the prima:/ system pressure at steady state minimum. The reactor is not permitted to reach a condition corresponding to a calculated departure from nucleate boiling ratio (DNB ratio) lover l I TI; 1 v I
than 1.25 vithout actuation of an automatic scram. D2parture from nucleate boiling is defined in Section III-E. Center melting of the fuel is not permitted to occur during normal operation or normal transient conditions. Daring core life, the core power rating may be increased as the maximum to average power density decreases. In no event vill the maximum heat flux and linear power density conditions, as given for .e 22.1 MWt power rating, be exceeded. 4 B. COOIldTP FIDW Coolant Flow Rate The total primary coolant flow rate at operating temperature and pressure with one pump operating is 2 94 x 10 lb/hr. Eighty-five percent (85%) of the flow entering the vessel is effective for heat removal which is the same as Core I. l Pressure Drop Core II pressure drop is calculated to be approximately 0.1 psia greater than Core I even though the spring clip grid design has been modified. C. VlaIATION OF PRIMARY SYSTEM TDiPERAT' RE AND PRESSURE J Pressure i l l The maximum steady state primary system pressure variation including instrument errors and deadband, is +50 and -50 psi. The minimum III-2
steady state pressure is used to calculate the conditions in the hot channel under the vorst steady state conditions. Temperature The maximum steady state temperature variation, including secondary steam pressure control system errors and deadband, is 5 0 F. The maximum inlet temperature is used in the calculation of the departure from nucleate boiling ratio (DNBR). D. ENGINEERING HOT CHANNEL FACTORS The engineering hot channel factors used in the Core II design are 1.0k5 for F and 1.22 for F The subfactors used in obtaining q g. these values are described below. D>finition of Engineering Hot Channel Factors The total hot channel factors for heat flux and enthalpy rise are defined as the maximum-to-average ration of these quantities. The first com iders the local maximum at a point (the " hot spot"), and the latter involves the maximum integrated value along a channel (the " hot channel"). Each of the total hot channel factors is the product of a nuclear-hot channel factor describing the nuclear flux distribution and an engineerin5 hot channel factor to allow for variations from nominal design conditions. The engineering hot channel factors account for the effects of flow conditions and fabrication tolerances and are made up of subfactors accounting for the influence of the variations of fuel pellet diameter, fuel density, enrichment and eccentricity; I III-3 e I
fuel rod diameter, pitch and bowing; inlet flow distribution; flow redistribution; and flow mixing. "'able III-1 is a lict of the subfactors comprising the engineering factors. Basic for Confidence The enE neering hot channel factors are estimated using the i maximum allovable manufacturing deviations to determine each engi-neering hot channel subfactor. These subfactors are combined by multiplication which gives a resulting engineering hot channel factor with the maximum deviations occurring shuultaneously at the hot spot or hot channel. These estimated engineering hot channel factors are combined with the nuclear hot channel factors to establish thr design. The total engineering hot channel factors listed belov apply to both the pelletized and vibratory compacted fuel. TABLE III-l ENGINEERING HOT CHANNEL FACTORS Fq , Subfactor Saxton Core II Pellet Diameter, Density 1.041 Enrichment, and Eccentricity l Rod Diameter, Pitch and Doving M1 Resulting F 1.045 q t l l (1) Nucleonics, Vol. 20, No. 9, September 1962, " Engineering Hot Channel Factors for Open-Itttice Cores," H. Chelemer, L. S. Tong. l I III h l 1
[g Subfactor -Saxton Core II Pellet Diameter, Density, 1.037 Fhrichment, and Eccentricity Rod Diameter, Pitch and Boving 1.10 Inlet Flow Ma1 distribution 1.07 Flow Redistribution 1.05 Flow Mixing M Resultirg F 1.22 E. DEFARTURE FROM NUCIZATE BOILING General Westinghouse has developed correlations to predict the condit1'ons at These are the W-2 DNB correlations (1which which DNB vill occur. are based upon an extension of a dimensional analysis performed,by Griffith(2) and verified by a parametric study of about30$0 pub-lished DNB data points. Since the data points were obtained by various independent investigators, the correlations include both systematic and random, uncertainties. The data points used to develop the W-2 DNB correlations cover the full range of points for NR cores. Therefore, this correlation may be confidently used in the cesign of WR cores. i l (1) Tong, L. S., Currin, H. B., and Thorp, A. G. II, "New Correlations Predict DNB Conditions," Nucleonics, Vol. 21, No. 5, May 1963 (2) Griffith, P., " A Dimensical Analysis of the Departure from Nucleate Boiling Heat Fa.ux in Forced Convection," WAPD-TM-210, 1-1959 l \\ l [ III-5 l l \\ l l-
Develorcent of the W-2 Correlation The correlataur.: of DNB data obtained from various investigators are not in agreement.(1) Jens and Iottes(2) and Zenkevich and Subot$ n reported that DNB is strongly influenced by local sub-cooling and local heat flux.
- Cook, however, found DNB to be independent of local peak heat flux because, for a given inlet enthalpy and flow rate, DNB occurred at a sa.me power input for both a uniform and cosine power distribution.
These apparently conflicting results are explained by the hypothesis that there can be two types of DNB: namely, a q"-DNB due to excessive local heat flux and H-DNB due to the high enthalpy at the vicinity of heated surface. The folleving nomenclature vill be used in the discussion of the V-2 correlation: A Cross-sectional area for flow, ft /g Heat transfer surface area, ft /S The bubble angle between liquid and solid surfaces (1) Tong, L. S., Currin, H. B., and Thorp, A. G. II, "New Correlations m Predict DNB Conditions," Nucleonico, Vol. 21, No. 5, May 1963 (2) Jens, W. H., and Inttes, P. A., " Analysis of Heat Transfer Burnout Pressure Drop and Density Data for High Preessure Water," ANL h627,1951. (3) Zenkevich, B., and Subotin, V., " Critical Heat Flux in Forced Cir-culation of Water; Subcooled to Boiling," Atomic Energy No. 8, USSR, 1957 (4) Cook, W. H., " Fuel Cycle Program - First Quarterly Report," GEAP-3558,
- p. 35, September 1960.
i III-6 ~
D Equivalent diameter, ft. 2 G Incal mass velocity, lb/hr-ft H Incalenthalpy, Btu /lb li " " N' in H Saturatedliquidenthalpy, Btu /lb f H Heatofevaporation, Btu /lb f AH (H ~N DE DE in ' L Iength in heated chtnnel from inlet to local point, ft. P heated periphery, ft. h p Pressure, psia 2 q"pg HeatfluxatDNB, Btu /hr-ft q" Average heat flux of test section, q" = f[g q"dz. Btu hr-ft b T Temperature, F ) AT Subcooling (T -T1ocal, F y T Saturation temperature, F f 3 pf Weight dens,ity of saturated liquid, lb/ft 3 Weightdensityofsaturatedvapor,lb/ft x Steam quality, percent by veight hxial length, ft. \\ Inletvelocity,ft/pec. p.3 Viscosityofraturatedvapor,lb/br-ft /tf VLscosityofsaturatedliquid,lb/hr-ft Sirfacetension,lb/ft cy-III-7
DNB Correlation in the Quality Region DNB in the quality region is due to excessive enthalpy in the vicinity of the heated surface. The W-2 correlation for this 3 type of DNB (B-DNB) is given by: M = 0 529 (H -Hin) + 0.825 + 2 36 exp (-204 D,) H DE f fg 6 exp(-150/10)-0.41H exp (-0.00k8 t/D,) - 1.12 Hfgg/r g# fg +0 5L8 Hfg The above equation represents the best fit to the existing data from 800 to 2750 psia. Figures III-1 and III-2 show how veil the-equation predicts the AH e tually measured in tests. Figure III-1 DE is a comparison of the measured and predicted M f r the existing DB lata from 800 to 2750 psia. The locus of points where the measured and predicted AH are equal is shown as a 45 degree line. An DE analysis of the deviations of experimental data from this line shows that measured DNB points fall vithin a range + 25 percent of the AH predicted by the W-2 correlation with a probability of DE 90 percent. A statistical analysis showed that at a 95% confidence level there is a 91 percent probability that DNB vill not occur if M is 25 percent less than AHDE (AH = 0 75 MD E) as predicted by the W-2 correlation. Figure III-2 shows the comparison of measured and predicted g g using the W-2 correlation for only those data points of a pressure 2000 psia. An analysis of the deviation of these data points from the locus of points where the measured and predicted MDB "#* '9""1 (again at h5 degree line) shows a narrower distribution than that obtained when all data points were included. For a confidence level i III-8 4 1 1 1
i of 95 percent, the measured DNB points in this range of pressure falls above 80 percent of the 4H predicted by the W-2 correlation pg with a probability of 9h percent. The effect of axial flux distribution was investigated using date reported by De Borto11 and was found to be insignificant. The ranges of parameters of the correlated experiraental data for AHDE "#* "8 f 11 #8 Geometries: Circular tube, rectan6ular channel and rod bundle Axial Heat Flux Distribution: Unifom and non-uniform Mass Velocity = 0.2 to h.0 x 10 lb/hr-ft Pressure = 800 to 2750 psia I/D = 21 to 656 Inlet Enthalpy: E Q /lb in Incal Heat Fluy = 0.1 to 1.8 x 10 Btu /hr-ft Exit Quality = 0 to 0 90 by weight DNB Correlation in the Subcooled Region i DN3 in the subcooled. reg on is due to excessive local heat flux and i l can be attributed to the interference of the liquid and bubble flows l l normal to the heated surface with influences rom the convection l effect of the cass flow rate. The V-2 correlation for this type of-l DNB (q"-DNB) is given by: (1) De Borto 1, B. A., et al, " Forced Convection Heat Transfer Burnout Studies for Water in Rectangule Channels and Round Tubes at Pressure above 500 psia," WAPD-188, 1958. l III-9 t ,m, r y e r y'
r! l 6 g,gggg) (3,g g,gy 37,e) q"De - (0.23 x 10 0.435 +.23 en (-0.0093I/D,)I 1 7 - 1.k exp -0 532 (H -H in}! g g!d} ( } f f l This equotion represents the best fit to the existing unifom heat flux data. Figure III-3 chovs how well the equation predicts the q"DE a tually measured in tests. It is a comparison of the measured and predicted q"DNB "8 "" analysis of the deviation of the data points from the locus of points where the measured and predicted q"p g are equal (45 degree line) shows that the measured DE points fall within f 20 percent of the q"pg predicted by the V-2 correlation with e probabiljsy of 90 percent at a confidence level of 95 percent. This means that, fcr the entire range of the data correlated, there is 66 percent proba-bility that q"-DNB vill not occur if q" is 20 percent less than 9"pg (q" = 0.8 q"pg) as predicted by the V-2 correlation. The following parameters are covered by this correlation: Geometries: Circular tube, rectangular channel and rod bundle Axial Heat Flux Distribution: Uniform Mass Velocity: 0.2 to 8.0 x 10 lb/hr-ft Pressure: 800 to 2750 psia L/D : 21 to 365 g l Inlet Subcooling (H -Hg): 0 to 700 Btu /lb f Incal Heat Flux: 0.4 to 4.0 x 10 Btu /hr-ft Subcooling at DNB: 0 to 228 F l l [ III-10 l v. ~ n
Application of DNB Correlations. in Design Definition of DIG Ratio H-DNB ratio is the ratio of the predicted enthalpy rise at DNB to the enthalpy rise from inlet to the corresponding point in hot channel. Local q"-DNB ratio is the ratio of the predicted DNB heat flux to the local heat flux at the corresponding point in the hot channel. Statistical Combination of DNB Correlation and Engineering Hot Channel Factors The occurrence of DNB depends' upon the statistical nature of the hot channel factors and the DNB correlation. The nuclear hot channel factors are not treated statistically because of the complicated treatment involved. Instead, the maximum calculated vslues are used with an added 10 percent margin to account for uncertainties. Engineer-ing hot channel factors are calculated on the basis of the vorst tolerances permitted in fabricatton. However, as was discussed earlier, due to the statistical nature of deviations in fabrication, the engineering hot channel factors can be treated statistically. The statistical convolution of the probability functions for the W-2. correlation for AH and the engineering hot channer factor f " b py3 results in Figure III h sho' ang the probability of not reaching MB i as a function of the minimum AH-DNB ratio calculated from the W-2 correlation using design hot channel factors. It can be seen that l usingadesignbH-DNBratioof1.25resultsinaprobabilityofover 99 percent of not reaching DNB. l Because the engineering hot channel factor F is very nearly 1.0, q l there is not very much to be gained by convoluting the probability functions for F and q"-DNB. Here the probability of DNB not being q i l ( III-11 l Y ~
exceeded is 96 percent when the design engineering hot channel factor for F is used together with a minimum DNB ratio of 1.25 as shown by Figure III-5 giving the probability function of q"-DNB. Film Boiling Heat Transfer Coefficient s Heat transfer ummediately after departure from nucleate boiling is conservatis ely assumed to be limited by film boiling. A period of transition boiling is neglected. The heat transfer coefficient in film boiling was obtained by correlating the existing data as shown in Figure III-6. F. INDRfdJLIC AND THERMAL DESIGN PAR /#ffERS S The hydraulic and shermal design parameters for Core II are given in Table III-2. TABIE III-2 HYDRtJJLIC AND THERMAL DESIGN PARAMETERS
- TOTAL CORE Total Heat Output (Initial) 22.1 MWt Total Heat Output (Initial) 75.h x 10' Btu /hr
= Heat Generated in Fuel 974% System Pressure, Nominal 2000 psig System Pressure, Minimum Steady State 1950 psig
- Based on a radial peak rod to core average factor of 2 30 and a linear power density of 16 KW/ft.
III-12 ( 7:
Coolant Flov 6 Total Flov Rate 2 94 x 10 lb/hr Effective Flow Rate for Heat Transfer 2 5 x 10 lb/hr Flov Area for Heat Transfer Flov (Unit Cells) 2 51 ft Average Velocity alon6 Fuel Rods 58ft/see Coolant Temperatures N:minal In3et 520.0 F Maximum Inlet, including Instrumentation 525 0 F Errors and Deadband Average Rise in Vessel 20 7 F Average Rise in Core 24.8 F Average in Vessel 530 7 F Average in Core 532 5 7 o Average Film Coefficient 2540 Btu /hr-ft'-F Average Fi k Temperature Difference 58.0 F Heat Transfer a Active Heat Transfer Surface Area of Fuel Rods 498 ft Average Heat Flux 147,200 Btu /hr-ft Average Thernal Output 4.4Kv/ft bhximum Clad Surface hperature at 642 F Nominal Pressure Pressure Drop Across' Core 4.1 psi-Across Vessel, including Nozzles 11 3 psi III-13
l l CENTRAL CORE REGION (UG -Pu0 Fuel) 2 2 F Heat Flux Hot Channel Factor 3 61 q F Enthalpy Rise Hot Channel Factor 2,81 g 1 Nominal Outlet Temperature of Hot Channel 586.7 F Maximum Outlet Temperature of Fot Channel 591 7 F Max 12::um Outlet Enthalpy of Hot Chennel 5958 Btu /lb Saturation Enthalpy at Minimum S'.eady State 6659 Btu /lb Pressure Meximum Heat Flux 553,700 Btu /hr-ft Maximum Thermal Output 16.0 Kv/ft DNB Ratios local q"-DNBR at 100% Power, Nominal Conditions 2.62 Incel c"-DNBR at 120% Power,1800 psia, Max. T 1.87 g N.A.(1) AH-DNBR at 100% !Wer, Nominal Conditions N1-DNBR st 120% Pover, 1800 or 2200 psia, N. A. Max. Tg l OUTER CORE REGION (UO Fuel) 2 l F Heat Flux Hot Channel Factor 2.04 q l F Enthalpy Rise Hot Channel Factor 1 59 g Nominal Outlet Temperature of Hot Channel 558 9 F lhximum Outlet Temperature of Het Channel 563 9 F [ Maximum Outlet Enthalpy of Hot Channel 5589 Btu /lb l Saturatlon Enthalpy of Hot Channel 6659 Btu /lb 2 M.tximum Heat Flux 301,600 Btu /hr-ft Maximum Thermal Output 905xv/ft l DNB Ratios Iceal q"-DNBR at 100% Power, Nominal Conditions 4.86 Incal q"-DNBR at 120% Power,1800 psia, 3 47 Max. Tin III-14
DNB Ration (cont'd) M-DNBR at 100% Pover, Nominal Conditions N.A.(1) M-DNBR at 120% Power, 1600 or 2200 psia, N.A. Max. T1n G. CE7TntAL TEMFATURE OF THE HOT PELLET 1. Uranium Dioxide Fue) The temperature distribution in the pellet is mainly a function L of the uranium dioxide thermal conductivity and the local power density. The absolute va]ue of the temperature distribution is affected by the cladding temperature and the thermal conductance of the gap between the pellet and the cladding. L The occurrence of nucleate boiling maintains maximum cladding surface temperature below about 647 F. The cold gap between pellet and cladding is specified to obtain contact during full power operation at the core hot spot. A contact conductance of 1000 Btu /hr-ft-F,atthehotspotwasestimatedforstainless steel cladding.(2) For Zircaloy clad a contact conductance of 1200 Btu /hr-ft wasused.(3) m . :2 (1) Not applicable - no bulk boiling in channel. (2) E. A. McCabe, J2., " Thermal Design Aspects of the Yankee First Core Puel Rod," YAEC-106, November 1960. (3) R. A. Dean, " Thermal Contact Conductance Between UO and Zirealoy-2,* 2 CVNA-127, thy 1962. l- .r [ t III-15 L
The themal conductivity of uranium dioxide was evaluated from published recults of recent work at ORNL(1}, Chalk River ( }, dnd WAFD(3)b) The recommended curve (Curve B) for thermal con, ductivity is given in Figure III-7 The section of the curve at temperatures between 0 F and 3000 F is based on the data of Goafrey, et al( The section of the curve between 3000 F and 5000 F was based on two factors: a) In-pile observations of fuel melting dictate a positive temperature coefficient for conductivity above approxi-mately 3000 F. The temperature dependence in this range should conform to an exponential curve since this reflects the most credible physical interpretation of the high tem-perature ephductivity increase, (1) T. G Godfrey, et al, " Thermal Conductivity of Uranium Dioxide and Armco Iron by an Improved Radial Heat Flow Technique," ORNL-3556, June 1964. (2) J. A. L. Robertcen, et al, " Temperature Distribution in UO Fuel 2 Elements," Journal of Nuclcar Materials 7, No. 3,1962, pp. 225-262. (3) R. N. Duncan, "Rabbitt Irradiation of UO," CV.9A-142, June 1962. 2 (4) J. A. Christensen, " Thermal Conductivity of Nearly Stoichiometric UO..- Temperature and Cor: position Effects," WCAP-2531, November 1963 2 (5) G. R. Horn and J. A. Christensen, " Identification of the Holten Zone 7.rradiated UO," ANS Winter Meeting Transactions,1963, p. 348, No. 5 2 III-16
b) The area under the recommended curve in such that the integra) it d t is equal to approximately 97 v/cm as given by Robertson, et al(1) and Dancan(2) This value is based l upon the interpretation of fuel melt radius as deter:nined i at Hanford(3) and chan River (l). The thermal conductivity curve can best be reprenented by the following equatior.o: a) Temperuture Range - O f T $ 1650 C k= + 1 32 x 10' exp (1.88 v 9' T) b) Temperature Range - 1650 $ T 3 2800 C -3 k = 0.019 + 1 32 x 10' exp (1.88 x 10 7) 2 withk2.nv/cmCfor94percentdensefuelnudTin C l 2. Uranium Dioxide - Plutonium Dioxide Fuel I Pelletized Fuel I The ther:::a1 conductivity of the UO -Pu0 sintered pellets was 2 2 taken to be the same as that for UO fuel as given above. The 2 hct, rod in Sutton Core II will be of this type with either { i l (1) J. A. L. Robertson, et al, " Temperature Distribution in UO I"*1 2 Elements," Journal of Nuclear Materials 7, No. 3,1962, pp. 225-262. (2) R. N. Duncan, "Rabbitt Irradiation of UO '" CVNA-142, June 1962. 2 (3) G. R. Horn and J. A. Christensen, " Identification 6f the Molten Zone in Irram;ed U0," ANS Winter Meeting Transactions,1963, p. 348, No. 5 2 III-17 i .\\ l i
Zircaloy or stainless steel clad. The maximum central tempera-ture of the hot pellet at steady state for each type of clad is: a) Zircaloy, Pu0 -UO 3400 F 2 2 b) Stainless Steel, Pu0 -UO 3440 F 2 2 c) Stainless Steel, UO 2550 F 2 Flux depression compatible with enrichment and density was included. All temperatures are vell below the irradiated PuO fuel melting 2 point of 5050 F and the irradiated UO melting point of 5000 F. 2 Daring a maximum overpover transient the centerline temperatures are: a) Zircaloy, Pu0 -UO 4000 F 2 2 b) Stainless Steel, Pu0 -UO 4060 F 2 2 c) Stainless Steel, UO 2930 F 2 Vibratory Compacted Fuel The present core loading plans are such that the vibratory compacted fuel assemblies vill not contain the highest power density rod. However, a maximum central fuel temperature was determined assuming maximum power density conditions. The curve of thermal condue-tivity is shown on Figure III-8. The maximum steady state-center- , line temperature was 4060 F. Again, a flux depression compatible with theoretical density and enrichment was included. For m<txhnnn U overpover transient, the centerline temperature was 4600 F. i III-18 ( ~-,-.,-.-y
6m 25 % 500 , s.,#- .t u = '"u .d 400 O 3 / ' ".. ".' " ". " [25) = g Jb'.g 300 h, 3 =' ok s 2 ,. s.~. m a 200 F j s/ l00 V; o' 0 0 100 200 300 400 500 600 Predicted AHDNB,BW M, COMPARISON OF H-DNB CORRELATION ~ WITH MEASURED DATA IN QUALITY REGION (p 800 to 2750 psla) j FIGURE III-l e e4 v i
600 7 20% 500 p .= f"' (, .j?." 205 ^ "/ g 300 7 g om P = 200 = E T 100 / l 0 100 200 300 400 500 600 Predicted AHDNB, BW. COMPARIS0N OF H-DNB CORREL.ATION WITH MEASURED DATA IN QUAll1Y REGION (p e 2000 -4 i FIGURE III -2 I
4.0 l .M 3.0 m:c L. 5 y a. "r i ~ = p e E 2.0 a ge w g s u y ..OF M I l.0 ,[ l l 0 20 4.0 1.0 2.0 3 6 Predicted q"DNB,10 Btulhf-ft l COMPARIS0N OF q"- DNB CORRELATION l WITH MEASURED DATA IN SUBC00 LED REGION (P = 800 to 2750 psla) l FIGURE III - 3 i I c w
1 z I.35 $s E l g I.30 o ?3 f.25 E E 'F = l.22 g AH u_ Confidence Level - 95% g ,,g 5 8 u,1.15 c o _k l.10 a g ?c g i.05 _s _Z_ 2' I I I I I I I 1 1 l00 99 98 97 95 94 93 92-91 90 I" PROBABILITY OF NOT REACHING H-DNS, PERCENT H-DNB PROBABILITY CURVE AT 2000 PSIA FIGURE III k'
1 f, 1.30 L -3 8 1.25 4 3e So E I.20 Confidence Level - 95% aW E F - I.'045 5 4 s m3 1.15 O i 52 l a: y 1.10 o ir 8 1.05 z .9 1-1 I l-1 I i 33 100 95 90 85-80 75 70 65 60 55 PROBABill1Y OF NOT REACHING g-DNB, PERCENT .q-DNB PROBABILITY CURVE AT 2000 PSIA FIGUFE III-5
,.l-l i f!? l l M 8 I I 1 O O 3 2 2 2 E - - - 3 3 3 3 3 6 C A A A 3 3 3 3 3 z3 N WW W 0 0 0 0 0 s0 IIl E 4 4 4 4 4 7 e R 1 I I E 6 G G 7 3 4 r r r r r 4 - 2 F t t t t t - P 6 E E E E - - - - - R MM M C C C C C A - S S S E E E E E wE R I A A A A A A A A H G T N S S S S S O O E I C N N N N N R E E E E E O 2 T I A R wW wM M M M M r . A L U N E O 88 8 E E E E E I I I I I A . S E S 8 8 0 S S S S S H G M R R E O C M RUA 0 0 0 4 0 0 0 4 O 0 0 SI 0 0 3 4 9 4 6 O 0 0 D 1 S S 0 6 1 8 9 8 5 4 s 41 2' 0 N 8 , A EP 3 2 3 2 2 2 2 2 i R j P A ^ T 6 X A 8 pk D E D A B C O E F G H J M c C N \\ R I E I 4 E F j S IN II p A R E , u T R V U T G D, p A 2 I E F -/,f ( H 2 w G 1Wrl/ 7 i I d A N X S L P I M O 1lg 0 0 B. 0 0 f ) 0, 2, 8 M Y 0 t e 5 2 I i j Y F 5 s I 6 E /g E L R, B. C A 9 ss 4 4 TS 0 0 O J 0 0 O O 0, s, / s i E / R M D w O g h F 2 / k r M g 6 4 Qg 6 4 2 I l ogC- ~ 1 !i il j', i ,l"
- I 1
<a ,+ TEMPERATURE, C 0 500 1000 1500 2000 2500 3000 9 I I I I I L i i i i i I I I i i i I i l I i 5 l I t o ,, cRnt 355s: m E i364 O.08 l O D: BMi-1448; JUNE 1960 V: KINGERY, J. AM.CER.30C.,37 (1954) I Y I 7 A: ARMOUR DATA, 1956 O I v: uuEA IG REPORT 51,(1960) 0.07 4 7 I Y .& : URAEA REPORT AERE-M/R 2526. (1958) g-Or >: REl3 WIG, J. AM CER. SOC. 44 (1961) p g g -l: GEAR-4624: (1964) ] y - A: Cuatt river, J. nuc. wi't3 7 90.3 (19s2)- dT=9fWID l 0.06 o U r y a: WAPD DESIGil CURVE - d 97W/CW l g C: G.E. DATA,tTOMS ET.Al, TRANS. AM. NUCLEAR 30CIETY, JUNE.19H, l 3 g3 gy f % W/CM 0.05 PAGE 106 - l g s ny e i \\ o -!E y E \\ ,o ?'v i 7*
- l
-i f- ,E N e" z 2 N dr o 0 \\ ' b_ on- ? ~ ~ ~ ~ ~ ~ / A 0 03 t w ~6 I" E A b i o' _ _ _M> 1. d - 7 >7 i iE ) = ~u h"> pt.e 4.< ' / is o.o2 y 1 y4 l 4 I 0.01 e i i O 0 1000 2000 3000 4000 5000 TEMPERATURE, OF THERMAL CONDUCTIVITY OF URANIUM DIOXIDE FIGUPE III - 7 .. -.. =...
.12 4 .11 o EURAEC (88.0f 1.5% T. D., VIPAC, UO ) 2 o EURAEC (88.0f 1.5% T. D., VIPAC, UO -5% Pu0 ) 2 2 ,;0 x Fiat (90%T.D., Swagedt 002) a Chalk River (90% T. D. Swaged, U0 ) 2 y. 09 a Daniet (87% T. D., VIPA.C, U0 ) 2 52 e.08 ~ 5 h.07 i- $.06 z O.05 a a x Proposed Curve for VIPAC g5M
- x in-Pile, improved Conductivity iE
'x Results from In-Pile Sintering O o ouo s 0c go o o .02 U U a Propcsed Curve for [ .01 VIPAC at Start-Up I I I I I I 500 1000 1500 2000 2500 3000 3500 TEMPERATURE C THERMAL CONDUCTIVITY OF VlBRATIONALLY COMPACTED FUEL FIGURE III-8
i I IV. MECHANIF 's.__D_ED_IGN i A. CORE ICADING I The il main fuel acsc miies (9 x 9 array) in the Saxton Core II fuel loading vill be made up of 9 plutonium fuel assemblies,11 new UO2 fuel assemblies, and the special hollow 51 rod UO assembly presently 2 installed in the reactor for the supercritical program. Tne 11 nev UO assemblies vill further consist of four assemblies of the existing 2 Core I design and seven assemblies of the new Core II design. Ac vith the Core I fuel assemblies both the plutonium and new UO2 asseruolies vill be comprised of two groapsJ one group of the 72 rod design, nr.d a second group of the hollov 63 rod design to accomodate 9 rod removable subassembly. A breakdown of the number of plutonium ard UO fuel nasemblies in each group is given in Ibble IV-1. 2 1 As with the present core in Saxton, two rod positions in each of the cain fuel assemblies in Core II vill be used for either in-core instrumentation as indicated in Section V, source rods, or removable fuel rods, depending on the location of the fuel assemb.ly in the core. In addition, the special L-shaped assemblie: presently installed in l the control rod slots in the peripheral fuel assemblies vill be reused with the fuel assemblies in Core II. Tne remaining fuel luniing for Core II vill be made up of the six control rod fol30 vers presently in the reactor, the supercritical assembly, und to tr of the removable 3 x 3 type fuel subassemblies .to be used in conjunction with the hollov 63 rod design fuel assembly. 1 IV-1 1 l l ~..
TABLE IV-1 FUEL ASSDGLY TYPr.S Number of Assemblien 72 Rod Assembly 63 Rod Assembly Plutonium Assemblies 8 1 Core 1 Design UO Assemblies 3 1 2 Core II Design UO Assemblies 5 2 2 Supercritical Arisembis' ($1roddesign) 1 IV-2 ~
B. FUEL ASSEMBLY DESIGN 1. Overall Construction The construction of thte plutonium and uranium fuel assemblies remains essentially the same as that for the Core I fuel assem-blies. No chance has been made in the overall fuel nseembly dimensions or fuel rod pitch and the fuel assembly cross section remaine identical to that shown in Figure 2031 of the original Saxton Final Defeguards Report. Theon3v_chtngeJn_the.. basic Core I fuel assembly design made for the Core II assertblies is in the type spring clip J g used for radiel support of the fuel rods in the fuel ascembly. 2. Core II Orid Ibsien The improved design spring clip used with the Core II fuel assemblies consists of interlocking sheet matal grid straps s brazed together in an egg crate type construction similar to the Core I grids. With the improved grid design, however, the fuel rods are supported _vith a six point support instend of a four point support as in the previous design. In the old design, the fuel rods were supported luterally on all four sides at each grid location by spring f:.ngers which extended out, in the form of cantilevers, from the mein body of the straps. This arrangement provided ample latert.1 support for the rods but offered no moment to restrain bending of the rods between the grids. ~ In the new grid design, the fuel rods are supportud ut. thin each grid lattice by a ecmbination of spring fingers ana pairs of stiff fomed dimples. In this case the spring fingen are t IV-3 =
located on only two ciden of each fuel rod, at 93 from each other, and are formed within the main body of the grid straps. - The pair of dimples associated with each spring are formed in the parallel grid strap on the opposite side of the fuel rod from the spring and are spaced evenly above and below the spring. Insertion of a fuel rod between each spring and its associated dimples preloads the spring,vhich, in conjunction with the dimples, produces both a force and a couple to restrain the k ser.mme - 1 rod against motion. The magnitude of the force and couple is sufficien+ to maintain contact between the fuel rod and the support surfaces under all anticipated combinations of assambly misalignment and reactor operating conditions. This built-in condition of the fuel rods at the grid locations vill minimize any possible fretting between the fuel rods and supporting grid straps. A comparative sketch of the old and new spring clip designs is given in Figure IV-1, C. FUEL DOD DESIGN 1. Dverall Designs The Core II plutonium fuel rods are composed of either pelletized or vibratory compacted fuel encased in either cold worked 304 l stainless steel or Zircaloy-k cladding (all four cambinations are utilized). The fuel rods are of the same overall size as the Core I fuel rods (nominally 0 391 inch in outside diameter and 39 051 ( inches long). Table IV-2 lists the number of rods, type of clad l and fuel configuration. IV-4 l
T/G1E IV-2 ITf1LI" TION OF PLlTf0NIUM FUEL ROD IDCATIONS IN CORE II d Fuel Conficuration, Cindding Number of Rods Telletized Zr-4 h73 ib11etir.ed 30h SS 20 Vibratory Compacted Zr h 136 Vibrntory Ccmpacted 304 SS 10 7 Ir:ctrumentation locations Source Rods DT b Total 651 The Core II uranium fuel rods consist of UO pellets encased in 2 cold worked 304 stainless steel and, like the plutonium rods, are of the same overall size as the Core I fuel rods. A total of k83 of these fuel rods are used in the Core II design fuel a s s er.iolie s. The remaining fuel rods vill be of the Core I design. /dl of the fuel rods are provided v.th end gaps to allow for I differential axiul growth between the fuel and clad and to pro-i vide void space ibr fission gases, moisture and other gasec { contained within the fuel. \\\\M,\\ (v:J ), f To compensate for radial grovtb resulting from thermal expan-p ~ sion and swelling of the fuel under the high burnups, the pelletized fuel rods are also provided with diametral gaps between the fuel and cladding. However, because of the lover density of the vibratory compacted fuel, no diametral gap is required. ihe cladding dimensions, pellet sizes, and radial and end gaps required for'the Core II fuel rods are given in Table IV-3 I IV-5 i l
l T/JILE IV-3 f CORE II FUEL ROD DIMENSICAS f Puo, Fuel UO, Puol Clad Type Zircaloy-b 30h SS 30k SS Clad inside diameter, in. 0 3445 0 361 0 361 Clad vall thickness, in. 0.0233 0.015 0.015 Rod overall length, in. 39 051 39 051 39 051 Pelletized fuel rod Pe11et diameter, in. 0 337h 0 3558 0 357 Pellet length, in. 0 366 0 366 0.600 Diametral cap, in. 0.0071 0.0052 0.004 Pellet stack height, in. 36.6 36.6 36.6 End cap, in. (rdninum) 0.726 0.609 0.612 Vipe fuel rod Fue3 column height 36.6 36.6 End gap (minimum) 0.784 0 784 IV-6 f l ~- - - ~, - N-"""*T+e"*-v: w-
-. -. ~. - -. - -. - - J In order to enst.re that the fuel remains r,olidly packed in the fuel rode, a retaining device is used at the upler end of al]
- clletizea plutonium and uranium fuel colutans.
The retainer le designed to prevent axial motion of the fuel under the loads expected during handling and shipping of the rods, but permits free axit.1 expansion of the fuel relative to the cladd'.ng duttnc reactor operation. 2. Puel D3rden and Charseteristice i The fuel for both the Ielletized and vibretory compacted plutonium fuel rods is composed of a mixture of natural uranium dioxide and plutoniam dioxide powders. The enrichment for both types of fuel rods has been set at 6.6 v/o pu0. The mixei oxide 2 0 vill have an initial celting point of 5170 F. The fuel pellets are produced by sinte ring cozIncted powder to obtain a final density of 94 12% of the theoretical solid fuel l density. The Iellets are centerless ground to the required l diameter after sintering. In order to provide axial space within the pellet stack for differential cylancion resulting from the l radial thermal gradient in the fue2, one end of each pellet is dished. The pellets are then stacked with the dished ends oriented in one direction to s stack height of 36.6 1 0.183 inches (one half a pellet height). Although the tolerance on the pellet otack is fairly lar6e, the end gaps in the fuel rods are con-trolled to a ninimum of 0.609 inches by the addition of thin sluminum oxide spacers at the top end of the ste k. The vibratory compacted fuel is compacted within the clad to a density of 8711% of the theoretical solid fuel density. The fuel colutans for the vipac fuel vill be 36.6 1 0.188 in., and the end gaps vill be controlled, as in the pelletized rods, to a 1 minimum of 0.784 inches by the addition of the aluminum oxide spacers. IV-7 s -r-e -v rww--e~-=-->va-mer -r,,---+,r r
n.w-,r-r,,-e,
.,w-s-e,
n-,--vw,
-w v.
The uranium ide3 for the core II design fuel rods is composed of $.69 7 v/oenrichedUO2 pelletr. As with the plutonium pellets, these pellets are produced by sintering and are centerless ground to size. Because of the longer length of the UO p 11ets, h vever, the pellets are 2 dished at both endo and the tolerance on the 36.6 inch stack height is 300 inch. 7ht tolerance on the end gap in this case is held t o 4,.068 inch with the aluminum oxide spacers, The eoisture content and allowable gas vo unes specified for the Core II fuel are as lirted in Thble TV-4. 3 Cladding TWsicn The fuel cladding c3nvicts of either 10% cold worked type 304 stainless steel or cold drawn Zirenleyah tubing. She fuel rod design for the stainless steel clad rods was established with the criteria that the cladding be free standing under des 1Cn pressure and temperature conditions, that diametral contact between the pellets and cladding occurs only under the worst expected tcleranes, pcVer and burnup cambinations, and that the internal gae pressure at the end of life is less than the reactor operating pressure. The fuel rod denign with the Zircaloy-4 cladding was established using a criteria similar to that used for the stainless steel with the exception that diametral contact between the pellets and clad is not limited entirely to the vorst tolerance, power, and burnup combinations. Because of the creep properties of Zircaloy-4 at high temperatures, it is expected that some reduction in clad dianster may occar in the hot zone of the fuel rod.
- However, since this creep will'be limited to the high temperature region of the rod nr.d vill cease upon contact between the fuel and the clad, it vill not affe:t the integrity of the clad nor vill it introduce any undue hazard into the operation of the reactor. The clad thickness, diametral clearance between the pellets and cladding, and the fuel rod internal void vclume for both the stainless steel and Zircaloy-4 clad were established consitt*nt vith the foregoing criteria..
I IV-8 I
The maximum clad stresses calculated for the pelleted Plutonium fuel are compressive and occur at the beginning of core life when i the internal pressure is taken to be zero. These stresses are calculated on the assumption that the clad cross section is oval, and the external prersare then induces bending strestes at the major and minor clad axes. Table IV-5 lists the values of these atrecres for both the ott.inless steel and Zircaloy-h cladding. The values listed in Table IV-5 are naximum values and will be reduced after a short period of operation at power because of internal pressure generated by the water vapor and gases shat are 3 eft in the fuel during manufheture as shown in Table IV-h. The maxin'n internal fisMon gas pressure at the end of life has been calculated to be less than the external coolant pressure for both the stainless steel and Zirealoy-k cind fuel rods. Thus the net pressure acting on the clad vill induce no tensile stresses, and the maximum tensile stress vill be due to thermr.1 stresses only. 7he maximum tensile thermal stresses calculated at the hot spot for both clad materials are:
- 1) Stainless steel clad - Thermal stress at outer curface = 7780 poi
- 2) Zirealoy-h clad - Thermal stress at outer surface = 2380 psi l
l The streanes for the UO stainless steel clad fue1 rods for Core II 2 j are lover than the stresses for the Plutonium stainless steel clad j fuel rods, and thus have not been included in Table TV-5 l As a further safeguard against clad failure, the end plug veld procedures have been established to obtain a veld penetration equal j to 90% or more of the minimum wall thickness of the cladding. s ,,o IV-9 .n.,..,
- a. - - -.-,
...nn-- -n -r-r---- ,n-~~'--, -,~-w~
Tha end plug velds for the UO r ds vill be 100% radiographed 2 and thc pelletized Pu0 r d end plug velds vill be 100% dye penetrant 2 inspected. The velds for the Vipac puo rods vill be ultrasonically 2 incpected. Other procedures that will be used to ensure the integrity of the end plug velds include sectioning and inspection of sample welds and as a final check, the finished fuel rois vill be helium leak tested. D. JUSTlTICATION FOR TE RE-USE OF TE C0KfROL ROD FOLICWERS AND L ASSEMBLIES Because it is planned to continue to use the L assemblies and control rod followers presently in the core for operation with Core II, our analysis was performed which indicated that such continued operation was safe. The analysis vas based upon the worst combination of burnup, pellet and cladding tolerances and fuel ovelling based on a fuel density of 93% of theoretical. The basic criterion used to establish acceptabil-ity of the re-use of the followers and L assemblien was that the end of life internal gas pressure should not greatly exceed external coolant pressure at operating conditions. The results of the analysis are presented in Table IV-6. The results show that in order to meet the above criterion the two center followers (central rods f2 and f5) would have to be moved to peripheral locations during the operation of Core II. In addition to limiting the increase in internal gas pressure, the shift of the fo11 overs also limits the peak pellet to clha pressure due to fuel svelling to very low levels. i IV-10
tf TABLE IV k MAXIMUM ALLO'4ABLE GAS AND VAPOR CONTD0' FOR PLiTr0NIUM FUEL MIXTURES Plutonium Plutonium Uranium Felletired Fuel Vipac Fuel Pellettred Fuel H 0, ppm 30 100 30 2 N, ppm 100 100 75 2 H, from hydrocarbon 2 impurities ppm 15 5 Total gas release at 1000*C .05 .067 (Exclusive of water ) SCC /gm l IV-ll .. _ ~.. -
_ _ _ _. _. _. _. -. _ ~. _ _. _. _.. _ _ _ _. _ _ _ _ _ TABIE IV-5 MAX 3 MUM PhESSURE STRESS PLUS THD NAL STRESS IN FUEL CLAD AT BEGINNING OF CORE LIFE 0.2% Yield Circumferential Stress (Psi) Avg. Strength Cladding Clad Temp. at Citad Minor Axis Major Axis Material Condition (*F) Temp. (psi) Inside outside Inside ou+ side St ainle ss Avg. Rod 600 65000 - 9130 L7k90 -53500 - 3120 Steel Hot Spot 672 62000 -37960 -16640 -41040 -15500 Zircaloy-4 Avg. Nod 605 55000* -13150 -25030 -28700 - 9500 Hot Spot 692 $O00* -23750 -14560 -24710 -13620
- O.2%yieldstrengthincircumferentialdirectionwhere(yieldstrength circumferentit.1) = 1.k75 (yield strength longitudinal - 2000).
I ( i-IV-12 l
6 l 1 TABLE IV-6 ColfrROL ROD FOLWJER AND L AS3DGLY BURITJP A!MIXSIS Puel Rod Peak Durnup N ak Power Fission Gas hllet-Cind (MdD/MTU) Density Pressure (psi) Contact 1 (xW/tt) Pressure (psi) L-Assemby Rod 43,9X) 7.hk 689 0-1 Control Dod Follover Rod Same Locations as Core I 55,000 10.82 2993 2900 }bved to Periphery in 49,700 8.22 2070 20 Core II e i l l s l l r- . t 0 l IV-13 4 i t 1 - y-- ey w , rv 9.-=w. rw.w4..m.,--wm--.m,ym,.. ,.p,_ n, ,,ww,,-,, ,wwe,..,,..,.,,,,,,mg,,wg.,,3._.!.,,e,-4
- w.
yw-.,,,, ye,. e ,.-,w%w%-,e-
t N--] N+)- l { ( -[} - [ ((_ p j u u -- p _ d
- j. -.
~ / I, () ( h~ j i c c c c-c-c- _g l
- I I
I I I l I (' -+ 1 i l l k l l l _I_ _if-l 1 ( l ) l o vo vo O vovo l v CORE II CORE I i SAXTON GF:ID DESIGN +fs FIGURE IV-1 y, \\
i V. INSTRUMENTATIOl{ A. IN-CORE INSTRUFF.HTATION l In erder to provide experfurtal meacurements and centinuous monitor-ing of conditions in the core, the Oaxton Renctor vus provided with a 1 i relatively extensive in-core instrumentation system. The system in capable of measuring outlet temperature, flow rate, pressure drop j and the magnitude and axial distribution of the neutron flux. No modifications of the system are required or planned because of the installation of a partial plutonium core.~ The in-core inctrumentation system consists of the following compo. nente: Pitot Tubes Inlet and outlet flow rates may be monitored at positione 2E, 2F, 3E, kB, 40, 4D, 4E, kF, SD, and SE and inlet flov rates only raay be monitored at positions 3D, 3B, and 50 In additson, coolant pressura drop acrocs the core may be measured at positions 3D, kF i i and 50. j Thermocouplen l Puel assembly outlet tet.perr.tures are tr.easured over the central position of the following assemblies: 2E, 2F, 3E, 4B, 40,,Y.', hD, kE, 4F, 50, SD, and "". In addities several thermocouplea are provided over the area of asco,bly 2E in orde:,* to monitor the outlet tgmperature distribution. A V-1 1 i l l w ..a-._.,_.v_. u.. a.u.._.,. _ _ __.~.-,. _.. _,_ _,. ., _ _.. _, - _. -. _.. _.,. ~. _. _ _
Flux Viry Activation of manganene in carbon steel vires is used to determine the magnitude and axial distribution of pover jn the fuel rods adjacent to the flux vire thimbles. The vires are moved by remote operation from a st. rage position into the core for activation and thence into a scannira position for scannng with a sodita todide crystal to determine flux data. It is probable thst this system may be equipped with small, moveable detwtors in the place cf one or two flux vires in order to get more rapid results and interpre-tations on flux levels and iesponces to system changes (e.g. control rodmotion). l l Figure V-1 shows the location-of the in-core instruments it relation to the Saxton core. The central nine assemblies (C, D, E - 2, 3, h) vill contain the plutonium fuel. l B. pI/JT SITE MONI70 RING In order to provide a complete radiation detection catability at Saxton, an alpha monitoring system vill be added to the plant site monitoring procedures...tcently being used at thxton. The alpha monitoring system vill be capable of detecting and indicating alpha contamination levels throaghout the Sarton plant. The alpha monitoring system vill be capable of detecting surface contamination and air particulate contamination. Surface contamina-tien vill be monitt. red using two separate systems. One systea con-sists of portable (battery powered hand held) survey meters that have 4 detection sensitivity range of 2 to 2000 alphas per square centi-meter per minute. The second system that vill have a more sensitive detection limit for surface contamination is a portable smear sample hit. The hits are equipped with the necessary laboratory analysis instrumentation to detect alpha activity levels of 0.2 alphas per square centimeter per cinute. V-2
I Air particulate contanination will also be monitored using twa separate syst ms. The alpha contraination of the vapor c9ntainer will be monitored by an on-line air sampling syste. This system takes an air sample from the container through a closed, sealed cystem, passes it throuch a scintillation detector, filter p1per asembly and then returns it to the vapor container. Perticles greater than 1 nicron in diameter are collected by the filter paper. The paper is then viewed by a photomultiplier-scintillation crystal combination and monitored for alpha activity. The system containd sufficient delay time between collection and monitoring to pemit the decay of the masking Radon-Thoron activities. The overall sensitivity of this system is of 1 x 10-12 pc/cc for Pu-239. In order to provide a flexible air monitoring system that can be used to monitor specific areas of the conta1 ment or areas not covered by the installed air sampling system, portable air sampling collectors and associated laboratory analysis instrumentation are also included in the alpha monitoring eystem. This portable air monitoring system will have a detection sensitivity of 2 x 10-12 pc/cc for Pu-239. These portable detectors will be used in conjunction with the routine radintion protection program at Saxton or in opecific instances such as sampling and analysis of the reactor coolant. V-3 I ___.__._..-_____._.m.___
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VI. ACCIDE!C ANALYSIS A. GENERAL The use of a Iertial loading of plutonium in the second core of the Saxton reactor requireo a re-evaluation of some of the accidente previously snalyzed and reported in the Saxton Final Safeguards Report and in the SafeEuerdo Report for Phase I of the Saxton Five-iear Research and Develop-tent Program. The informtion in Section 502 of the Fimi Safeguards Report reinting to the possible causes of accidents and the safeguards provided applies to the accidents analyzed for this report and vill not be repeated. ( The magnitude and rate of potential pywer excursions in the Saxton reactor are reduced by three inherent negative reactivity effects associsted with low-enriched preesurized vater reactors. The negative moderator temperature coefficient brings about a reactivity reduction with incrences in moderator temperature. The time lag associated with the diffusion of heat through the oxide pellet and cJadding maxes this comicratively 3arEe coefficient important primarily in limiting slow power excursions. The negative Doppler coefficient reduces reactivity with an increase in fuel temperature as a result of increased resonance absorption of neutrons in plutonium ?.nd uranium. This reactivity effect is prompt since fuel - heating begins immediately vith an increase in reactor power. Thus the Doppler coefficient is particulnrly important in limiting rapid power transients. Formation of steam bubbles in the core due to net boiling of water also reduces reactor power. The reactivity reduction associated with the forma-lon of steam bubbles is the " void coefficient" which, since net boiling does not normily occur in the core, is important only in accidents that result in a low primary system pressure (i.e. loss of coolant or steam break). VI-1
The reactor safety system monitors reactor operation and prevents core damage by causing a reactor shutdown when necessar/ to keep the fael element cladding in the hottest channel from melting. The safety system functions prinarily in the event of equipment failures or operational errors. Additional protection against a contamination hazard is provided by the containment vessel, which houses the entire primary system. In the event of a maximum credible incident, the contaircent vessel and associated shielding protect the area surrounding the plant from contamination and radia tion. To give assumnce that safeguards incorporated in the plant cre adequate, it in necessary to analyze arid evnlunte their effectiveness under a number of postuhted accident situations. The most adverse combination of system parameters for each particular accident was analyzed. Therefore, it was not necessary to analyze each accident for all possible comoinations of system parameters that may occur during core life. The probability of obtaining this most adverse combination of system parameters simuttaneously with the l corresponding accident is, of course, very small. Therefore, the results of the accident analyses must be considered as a limiting case for each l accident considered. In the same voy, reported DNB ratios for accident l l conditions actually represent the lower bounds of all possible DNB ratios that may occur for a given accident condition. The accident analyses performed for Core II havt be 4 ba' ad on the assump" tion of che nical shim operation at the beginning of 3 ".e as the worst het channel factors occur under these conditions for the l partial plutoniLWore II.. The accident arelysic for non-chemical shim l (rodded) operation of this core vould y1cid less severe result s. The use of contrcl rods for some control of the core reduces t'e peu rod to core average radial power ratio as listed in Section II so hnj.or e. given power level the maximum linear power density vill be below 16 Kv/ft which is the design limit. VI-2
L. REACTIVITY ACCIDE!CS Reactivity accidents are those which result in the addition of more reactivity at a faster rate than is required for normal changes in core power level. ) Such reactivity additions might be caused by control rod motion, by moderator temperature changes or by soluble poison removal. As a rebult 'f such addition, the reactor power level vill increase at a rate and with a saagnitude that depends on the rate and magnitude of the reactivity addition. The inherent negative reactivity effects in the core tuder these conditions vill compensate to some extent for the reactivity addition. In addition, the reactr r safety and protection systems are designed to protect against conditions vhich might endanger the integrity of the core. However, as in any reactor, continued reactivity addition accompanied by failure of the safety systems and operator errors can result in daT. ace tc thu core and release of fission products to the reactor coolant. 1. Uncontrolled Rod Withdrava?dt Cold Startup The condition for uncontrolled rod withirtrwal during startup procedures requitw a series.of multiple failures in the nuclear instrumentation and control systems coupled with simultaneous operator error and violation of procedure. The continuous witharaval of the most reactive control rod group at the maximum rate in its most effcetive region vould result in a naximum reactivity addition rate of 7 2 x 100 b k/sec. The reactor control system is desi6ned to prevent movement of the two control rod groups (the inner two or the outer four) simultaneously so that large insertion rates are not possible. However, for this accident study, a conservative insertion rute of 2 5 x 10 A s/see was assumed. In addition, the following assumptions were also used. \\ VI-3 s E 1
a) thximum expected value of the moderator temperature coef ficient: 4 0 3 x 10 bk/k/r* ~ b) Minimum expected absolute value of the negative fuel temperature (Doppler) coefficient: -1.1 x 10~5bk/k/?' c) Initial average core moderator temperature of 70*F d) The reactor is suberitical by 2% bk at the initiation of the accident. e) Overpower set point initiates scram at 122% of nominal power (7% above scrum set point of 115%) f) Scram initiation delay of 1.1 see (0 5. see instrumentation and 0.6 see for control rod motion in region of little effectiveness) g) Scram insertion vorth of 2% b k/k at a linear rate during a 0 9 see period. (A total delay of 2.0 see from signal to completion of scram) 5) Nominal power level ich03% of the: initial design limit of ~ Y 211MWt .l ~ The results of the' cold startup accident transient are shwn in Figures VI-B-1, 2, 3 and 4 The reactor power transient shown in Figures VIeB-1.and 2 is terminated at the first peak by the Doppler coefficient of the fuel. Tne peak is at 115% of nominal power which would normally initiate a reactor scram signal but the accident analyses assumes no scram until a power level of 122% to account for errors and deadband. The pwer then decreases slightly and then increases again until the l high power level scram is initiated at approximately 137 eec. Figure VI-B-3 shows the avemge fuel, average clad and core water temperature l during the tmnsient. Tne m:.ximum values attained by each of these l variables is much less than normal operating conditions so that the scram occurs well before clad damage could occur. The core water temperature rises oniv slightly above its initial ambient value so that VI-k
it is subcooled to such a degree that DNB ration are vell above those for normal operatirg conditions. The reactor hot epot heat flux, as chown in Figure VI-B-4, reaches a mximum value of about 108% of normal steady state coriditions. The analyci? shows that the reractor and control cystem can safely withstand a rod withdrawal accident from a cold shutdown condition. 2. Uncontrolled Ibd Withdrawal - llot Shutdwn The uncontrolled rod withdrawal at startup with the reactor coolant at operating temperature vas also analyzed. The same assumptions used for the cold startup analysis apply except for the following a) Moderator temperature coefficient of -2 7 x lo-b bk/k/F* b) Fuel temperature (Doppler) coefficient of -1.0 x 10~5 b k/k/F* c) Core vater average tempemture of 530*F The results of the analyses are showr in Figures VI-B-5, 6, 7 and 6. The reactor power level duririg the transient is shown in Figures VI-B-5 and 6. The power transient is terminated after approximat ely 100 seconds tecause of the overpower scram circuitry. Thi. power level rose to 122% of the nominal setting before the Doppler effect of the fuel could become effective. The average fuel, average clad and core water tempentures are shown in Figure VI-B-7 The clad and fuel temperatures are well below normal operating conditiour, at their peak values during this transient. The core water temperature rises only slightly above the initial value of 530*F'so that. it is still subcooled. The heat flux response during the transient is shown in Figure VI-B-8. The peak of the heat flux reaches only 17% of the nominal steady state condition before scram is initiated. As in the cold startup analysis in the previous section, the DNB ration durin[, this transient are very much F greater than those at normal operating conditions. .VI-5
As the analysis shove, the reactor core is protected from damage by the control system during a rod withdmval at hot startup. Only simultaneous system failure and operator error could cause circum-stancer that vould lead to core damgt tod release of fission products to the coolant. 4 3 Uncontrolled _ Rod Withdravn1 at Power Requirements for manual rnd motion when the reactor is at pWer may arise due to the need to adjuut control roi positions to account for xenon changes or fuel burnup. These changes are alwly varying ( effects and therefore the requirements-for manual control rod movements i are infrequent and rod movements are small. s-A / In the unlikely event of an uncontrolled rod vithdrawal at power, the cnximum possible reactivity insertion rate of the most reactive control rod group is calculated to be 7 2 x 10-5 A x/sec. This rate assumes that the most reactive group ie moving through the region of most effectiveness,nt its maximum withdrawal rate. However, in, order to study a worst possible condition (not credible) a reactivity 3 insertion rate of 2.5 x 10- Ax/seewasassumedrorthepurposesof 4 the accident study. In addition, the following assumptions' vere made far the studyt a) Initial power level is 103% of nominal pwer setting because of calorimetric errors (100% = 21.6 wt), b) Primary coolant pressure is at its maximum nominal value of 2050 psia because of instrument errors. c) Minimum expected absolute value of the negative fuel temperature (Doppler) coefficient: -1.0 x 10-5 A x/x/r-t VI-6 4
__~ F i d) Minimum expected absolute value of the negative moderator temperature coefficient: -2 7 x 10 Ax/x/r-e) Scram initiation due to overpwer scrnm at 122% of nominal power (7% over 115( scram set point due to instrumentation errors) i f) Delay of ceram initiation 7or 1.1 see (0 5 cee due tc instrumenta-tion and 0.6 due to rod motion in a region of cenll effectiveness) g) Scram incertion vorth of 2$ A x/h at 11near rt.te durine o.9 see period (total deley of 2.0 Dec # rom signal to completion of scram) Figure IV-B-9 presento the nuclear power esponse, hot epot heat flux respont>e and moderator pressure recponse se o function of time after initiation of the accident. The minimum E!G ratio calculated by the W2 correlation that vill occur during the rod withdrawal transient is 2.05(bil-D13). This minimum P!G ratio condition occurs at 5 2 cec. This minimum ratio is well above the minimum design DNB ratio of 1.25 and indicates that under the most conservative conditionc that the power level scram protection system vill be able to protect the core. Additional safety yactorc which would preclude core daunge,due to an uncontrolled rod withdrawal accident at power are a) Control roi circuitry restricts roi motion so tgt marir.rm reactivity incertion rate could only be a fraction of the vnbe usec iri this study. b) There are three separate power level ceram circuits and operation of any two of the three vill initiate the scram. It r.ny be cot.Aaded from the above analysis that the uncontrolled rod withdreval at power accident is highly improbable and would not result in core damse unless there vere a simultaneous failure of the power level scram circuitry and operator error, VI-7 1 1 ,..._,,,,,_..s,,.__ ., ~...,...
l h. Boron Release Accident t l The results of extencive chemical shim experiments and opemtione both out-of-pile and in-pile ac the Saxton and Yar},ee reactcro as vell as detailed crud behavior follcw and eurveillance in these and other I reactors have not revealed any credible mechanisms for the boron release accident as postuhted in the Safeguards Report for phase I r of the Saxton Five-Year Research and Developnent Program. In addition, l the large amount of data that has been analyzed has shown that even if such an accident vere credible, the mximum resetivity releace rates that could be associated with this accident are less than those that could be experienced in the worst rod withdrawal hecident and are therefore vell within the control capabilities of the rer.ctor. For these reasons, the boron release uccident hac not been re-annlyzed for the chemical chim operation of Core II. 7 The basic objectives of the chemical shim control program and experiment i as outlined in the Safeguardo Report for Phnoe I have been the accumu-htion of reactor operating experience to nemonstmte the feasibility of the concept and the determination that no operating problems exist with chemica) shim control. The najor concern in the use of chemical shim control of a reactor was the possibility that boron might accumuhte i in some manner on the core surfaces and subsequently be released rapidly causing a hrge resetivity transient. It was this concern that led to the postuintion and analysia of the boron release accident. A secondary concern of the program van that boron accumulation on the core surfacec l might occur in an irreversible manner (that ie, the accumulation is not l reduced as the coolant concentration is reduced) and therefore lead to a significant reduction in core lifetime. I 4 Extensive cut-of-pile testa have dioclosed only two mechanisms by which '- significant borou accumulation could occur.(1) The first possibility (1) WCAp-3731, "Radietracer studies of Hideout at High Temperatures and Pressure," (June 1963), L. Ficone, D. Whyte, and G. R. Taylor. f 1. l l VI-8
vas indicated by the study of solutione during nucleate boilirt on the surfaces cf electrically heated reds. By means of a sodium tracer, it vae oemonstrated that, under nucleate boiling conditione in tha presence of unustally heav-/ crud deposits, a significant concentration of the sodium occurred in the crud on the curface of the rods. The accu =ulatic.n dicappeared rupidly when bolling on the surface of the rod was terminated. This concentration effect was observed only in the presence of crud deposits far heavifr than those expected or experienced in normal reactor operation. The cecond possibility indicated van the deposition of alkali borate salts on the surfaces of redo under nucleate boiling conditions. It has been observed that crud levels in pressurized vater reactors are reduced when the reactor coolant is maintained at an alkaline pH. Since boric acid is ensentially un-ionized at normal reactor operating cot.ditions, these benefits of high p11 can be obtained with the addition of relatively small arounts of alkali in the borated coolant. Out-of pile evidence indicates that if the alkali borate salt deposition vere to occur with a lithium additive, re-solution 'f the salt m16ht not occur. A third, less extensive mechanism, one of simple " exchange absorption" of borate by the crud deposits, is also known. The extent of-boron accumulation in the crud is dependent on the boric acid concentration in the coolant and the amount of crud. Bovever, the small amounts of borate that can be absorbed by this proceso make it insignificant from a reactivity point of view. Three tajor techniques have been used to demonstrate the absence of i boron accumulation in the crud during the chemical shim Weriment. The first of these has been to conduct a careful reactivity follow throughout core life and compare the predicted and observed activity. Particular attention was paid to the transition from nucleate boiling to non-nucleate boiling conditions. If boron hideout did occur with i VI \\';
.. - - -.. - ~.. the reactor in nucleate boiling conditions, it vould be observed as a reactivity loss, and in particular, as an abrupt change in the power coefficient vith the onset of nucleate b;,iling. The reactivity locs vould then reverse itself once nucleate boiling opemtion ceased. Tbc second technique was to conduct a careful follow of the alkali to boron ratio during the chemical shim experiment. Demor.stration l that this ratio did not change throughout the test vould indicate the absence of metaborate precipitation. The third technique used vas that of hot cell e:: amination of a test 3x3 subansembly a*ter appreciable operatn under chemical shim cu the fuel rods are conditions. tJud samples from the nur veighed, analyzed and the reactivity 3 et:4 o, the total deposit f estimated. Prior to the beginning of chemical shiu ui-tion, an extansive series of low power physics tests vere conducted followed by a pt,.iod of rodded operation at power. These operations provided the information l required to Eet a base point to conduct the reactivity follow during the chemical shim operation. The che.mical shim operation of the Saxton reactor began in May 1963 ard has-continued, almost without interru" )n, since then. A short period of non-chemical shim operation +' ' ace from January 30, l 19% to March 6,1964 to provide adab, _ information on the effect l of burnup on the reactivity in an unb' v. reactor. This chemical shim test program has cuccessfully demonstrated the feasibility of chemical shim control for pressurized vater reactors.( } No significant operating problems have been encountered during the program. Under the normal operating conditions of the Saxton reactor,. the test program data indicate that: (2) WCAP-2599, "The Chemical Shim Experiment," (August 1964), J..Weisman and I S. Bartnoff VI-10 l l
D a) Ihere is no significant accumulation of boron-containing material on the core surface during normeil plant operation. The deviation from the reactivity predictions were well within the estimnted error and no deleterious trends could be observed.
- e b) There is no decrease in core lifetime because of chemical shim operation.
The hot cell examination and analysis of the central 3x3 subassembly shoved no significant accumulation of high cross section materials on the core surfaces. c) Successful pH control of the coolant can be accomplished and there is no problem vith alkali stability in the coolant, d) The hot channel factors were in reasonable agreement with the calculated values. e) Chemical shin 'ontrol, under normal conditions, causes no hazardcus situations to arise that could affect plant operation. As a supplement to the Saxton chemical shim experiment and results as described above,,further experiments have been conducted in the Saxton reactor whigh demonstrate that the hideout problem does not exist even under:the conditions of abnormal amounts 'of crud deposits on the core surfaces. This experimental program is described in detail in Addendum No. 4 to the Phase I Safeguards Report. The experiment consisted basically of two phases. The first was the deposition of artificial crud on tLa core surfaces by injection of ferrous hydroxide into the reactor coolant while operating at essentially zero boron concentrations and maximum rodded power level. Folloving the deposition of sufficient amounts of crud as determined by calculations and visual examination of the central 3x3 subassembly, the second phase was to observe the reactivity behavior of the reactor during chemical shim operation at various power levels. The results of this experiment are available and provide additional assurance that hideout would not occur even if the careful chemistry control of the coolant as used in Saxton vere not followed. I VI-11
The results of the chemical chim experimental program at Saxton have demonstrated that the boron release accident as originally postulated at the beginning of the program is ntA credible even under the extreme operating conditions of the crud test. It may therefore be concluded that the requirements of an unexplained reactivity limit and a detailed reactivity follow program as applied to the chemical shim operation of Core I are no longer necessary for the normal chemical shim operation of the Saxton reactor. 5 Steam Break Accident Introduction The rupture of a secondary plant steam line is reflected into the primary system as a step load increase. For small break sizes resulting in a step load increase within the transient capability of the Saxton plant, reactor power level vill increase to match the load increase and the decrease in equilibrium average primary coolant temperature. This condition, as far as the primary plant is concerned,is not different from a design step load change and no protective action is required. Depending on secondary plant condition:,, the operator vill decide when to l initiate a manual plant shutdown and to proceed with necessary repairs. If the step load increase resulting from a steam line rupture exceeds the transient capability of the plant, the reactor protection system vill automatically shutdown the reactor by either a low pressure scram or an-overpover scram and prevent-the core from reaching a DNB condition. Following' scram,.a large'pover unbalance exists between heat generation in the core (essentially decay heat) and heat extracted from-the primary loop as demanded by the steca flow through the break. As a result, the primary I coolant temperature decreases and due to the negative moderator ~ temperature coefficient, a' reduction in shutdown margin is to be I expected until operator action terminates heat removal. VI-12
s palysis The maximum credible steam break accident (equivalent to an 0.03h sq. ft, ductile rupture of the steam dome) vas analyzed for che:tical shim operation of Core I at 23 5 MWt with a spiked assembly generating 16 KW/ft in the Safeguards Report for Phase I of the Saxton Five Year Research and Development Program. The calculated reactivity coeffi-..nts for Core I and Core II are listed below: Core I Core II (WorstCase) Moderator Temperature Ak/k/F -4.6 x 10' -4.1 x 10 -5 Fuel Temperature Ak/k/F -0.4 x 10-5 -1.25 x 10 ~ Moderator Pressure Ak/k/ psi +h.6 x 10 +3 5 x 10 Because it has the largest absolute value, the moderator coefficient is the most i=portant in determining the react $vity transient following the steam break accident. As the above table shows, the partial plutonium Core II has a less negative moderator temperature reactivity coefficient and therefore the steam break accident for Core II will be less severe than for Core I. In order to assess the upper limit of possible reactor damage, the maximum hypothetical steam break vas analyzed for Core II. This steam break accident is the complete severance of the steam pipe at the outlet of the steam generator with an internal diameter of 5 5 inches (0.165 sq. ft. area). The steam flov through the break as a function of steam pressure was evaluated by the modified Darcy formula (ref. Flov of Fluids, CRANE Technical Paper No. 410). The formula is of the following form: i 6 VI-13
2 N,,n,g US = 0.525 Y d g where Y = net expansion factor for compressible flow through crifices, noncles or pipe d = internal diameter of pipe, inches 6P = pressure drop, psi E = recistance factor for pipe length from steam generator to break, including entrance and exit losses 3 V1 = steam specific volume, ft /lb 113 = steam flow, lb/see l This accident was anQvned for three different cases with the following reactivity coefficients: Moderator Fuel Temperature Temerature (a) 2000 ppm -2.7 x 10-k ok/k/ F -1.25 x 10-5 g jgjoy (b) 1000 ppm -3.4 x 10'E 4k/k/0F. l.25 x 10-5 Ak/k/F (c) O ppm -4.1 x 10-4 ok/k/F -1.25 x 10-5 h Ak/ / F For each case, the amount of negative reactivity insertion from control rods required to maintain the core suberitical by an 0.5') 4k shutdown margin during the course of the steam break accident was computed acsuming no safety injection flow. The shutdcwn margin will increase substantially with safety injection flow. Since the steam generator feedwater pumps are steam driven, the steam break accident causes temination of the feedwater flew, and as a result, the steam generator is empty in approximately 70 seconds and the cooldown of the primary system is terminated. l VI ~.t1 l i l 1 ) -,.-_l
Besults The primary system pressure response after initiation of the acci-de.nt is shown on Figure VI-B-10. The steam generator flow rate, the neutron flux and hot spot heat flux responses are shown on Figure VI-3-11. The steam generator inlet and outlet temperature and the core inlet and outlet temperature responses are shown in Figure VI-B-12. The minimum DNB ratio calcu3ated for the three cases analyzed is a q" DNER of 2.26 for the O ppm condition (nominally end of life). The reactivity required in control rods out of the core to prevent return to criticality during the course of the accident without safety injection is shown on Figure VI-B-13 Curve B assumes that all rods scram upon signal and curve A assumes that the most reactive rod (equivalent to 5 2% Ak/k) does not scram upon signal. Conclusions The Saxton reactor safety system vill adequately protect the core in the event of the maximum credible steam break accident. The maximum hypothetical steam break accident in conjunction with a stuck control rod and failure of the safety injection system has such a low pro'aability of occurrence that core damage can be accepted. In addition to the above occurrences, the reactor would have to be operating with boron concentrations well belov l thote planned for the nomal chemical shim operation of the plant. Even if the core were to return critical at full power, the stuck rod power distribution of Section VI (Figure II-10) shows that the ( maximum linear power density vould belessthan20KW/ft. 6. Cold water Introduction The analysis and results of the cold vater introduction accident as presented in the Saxton Firal Safeguards Report are unchanged for Saxton Core II. VI-15 l 5 1 -r
_._._ _._.. _.__ _ _ _. _ _.. ~ 7 Xenon Burnout The analysis and results of the xenon burnout reactivity transfer.t as presented in the Final Safeguards Report are unchanged for ~ I Saxton Core II. 8. Conclusionc As demonstrated by the preceding accident analyses, the transient. behavior of the Saxton reactor is such that negative temperature coefficients of reactivity or the reactor scram circuitry vill terninate power excursions well before any core damage could occur. Because of the inclusion of the mixed oxide fuel in Core II, there exists a possibility of some time delay in the effectiveness of the negative Dappler coefficient because of time-delay in the heat transfer from the PuC to the UO. However, the time delay associated 2 2 with this core ic more than an order of magnitude less than the shortest reactor period calculated for any of the transients and thub j it is not effective in alterin6 the Doppler' coefficient. Thus, the accident analyses and results that have been reported. are conservative and demonstrate that core dama6e or hazard to the public from reactivity accidento are highly improbable. I VI-16 g l
.. _ __ _ _.~. C. MECHANICAL-ACCIDEITIS 1. Ioss of Coolant 2 A loss of coolant accident for a 0.0375 ft break in the reactor coolant piping during normal operation at 20 MWt was analyzed in the Final Safeguards Report. This analysis was updated in the Safeguards Report for Phase I of the, Five-Year Research and D3velop-ment Program to cover reactor operations at 2'4 5 MWt. The results of these analyses are applicable to this Core II report and essentially conclude that the core vill not be uncovered if the safety injection system functions properly. 2. Ioss of Flow Accident A loss of coolant flov in the reactor coolant system could be caused by either a blockage of the system or by failure of the reactor coolant pump. Since,the reactor coolant system of the Saxton reactor contains no valves, complete blockage of the reactor coolant piping is considered incredible and such an accident is not analyted. Onh failure of the coolant pump is analy:cd and such failure could be caused by loss of electrical power or by mechanical failure of the coolant pump motor or of the coupling between the motor and the pump. As was assumed in the accident snalysis section of the Final Safeguards Report for Saxton, mechanical failure due to sudden seizure of the pump rotor is not considered credible. In the event of a loss of flov accident, the temperature of the primary coolant in the core vill increase because of scram circuit delays which allow the reactor to keep operating at full power vM k flow conntdovn ic occurrins. Such delays are critical because the DNB ratios are decreasing during this delay period and scram should be completed before possible DNB occurs. If the heat VI-17
generation of the core is not termirated rapidly enough to prevent Dh'B, fuel and clad temperatures may rise to excessive values and clad damage and failure could result. The first seven seconds of the flow coastdown curve used in the accident analysis is shown in Figure VI-C-1. This curve is the result of measured data taken for the Saxton coolant pump with the reactor at a power level of 14 MJt. The 1ol. loving additional assumptions vere use in the evaluation'cf this accident: a) Reactor power level is at 103% of nominal power settire because of calorimetric error (100% = 21.6 Mat). b) Maximumfuelpowerdensityis16.0kv/ft. c) Scram delay times of (1) Flow decay to set point and error in set point - 0 7 cee (2) Instramentation delay - 0.2 see (3) Rod motion in regions of small effectiveness - 0.6 see Totel - 15 sec d) A scram insertion vorth of 25 b k/k at a linear rate over a 0 9 see period. e) A maximum expected absolute value of the negative fuel temperature (Doppler) coefficient: -1.15 ;t 10~0bk/k/F' f) A minimum expected absolute value of the negative moderator temperature coefficient: -2 7 x 10' dk/k/F' 6 g) lov flow scram set point is 2.2 x 10 lb/h. The results of the analysis are presented in Figure VI-C-2. This figure shows the transient behavior of the Q and b H-DNB mtios (as calculated by the W-2 Correlation) during the loss of flow accident. The bH-DNB ratio is the critical parameter in this analysis and it reaches a minimum value of 1.89 at 2.2 seconds. The bH-DNB mtio is { VI-18
decreasedt at higher pressures and thus a maximum steady state cochnt pressure of 2050 psia was 'used in the bH-DNBcalculation. If a prestiure of 1950 psia is assumed, then the minimum bH-DIG is 1 92 at 2.2 seconds. Tha Q-DIG mtio is also calcuhted for the trencient and reaches a minimum of 2.08 in 1.6 seconds. The Q-IES rotio is decreased at lover pressures and thus a minimum steady state coolant pressure of 1950 peia is assumed for the Q-DNB calculation. From the above analysis, it sny be concluded that there is an ample mar 6 n of safety to prevent clad damaEe as a result of a i loss of flow accident. Cled damage could only result from coin-cident failure of the low flow scram circuitry and operator error. f VI-19 n ~ .-.- _,,- ~-.,- -,w<r,.... ---,+-----,a .r.-.m e.v,-- 4 ra,- .-n rn, ~,, - ,m, n.
D. F%XIM'JM EnUfKICICAL ACCIDDE The maximum hypothetical accident for the Saxton reactor is defined in Section 508 of the Saxton Final Safeguards Report. The assumption and condition for the accident are as follows: 1. Instantaneous release of the primary coolant along with the associated heat sources and sinks as described in Section 506 of the Final Safeguards Report. 2. Failure of the engineered safeguards to prevent core damage by recovering the core with borated water. 3 complete core meltdown and the associated fission product releases as calculated in Section 603 of the Saxton Final Safeguards Report. The energy release associated with this accident is assumed to cause an instantaneous pressurization of the vapor container. The resultant pressure is approximately 30 poig. The vapor container pressure transient for this accident is represented in Figure VI-D-1 taken from the Saxton Final Safeguards Report. The initial pressure peak falls rapidly to approximately 10 pois followed by a slight rise and than a gradual reduction and leveling off at approximately 5 psig. The rise following the initial peak assumes that decay heat is released from the core into the containment. Complete core meltdown is also assumed to occur with this accident because of malfunction of the engineered safeguards designed to prevent core damage. Because of the substantini fraction of the Saxton second core that is clad with zirconium, the consequences of a metal-vater reaction must be considered in conjunction vith the mximum hypothetical accident as described above. The plutonium region vill concict of a r.Aximum of nine standard Saxton assemblies of 72 rod spaces per assembly. Two of these spaces are nomally left empty to accommodate in-core instrumentation. In addition, ( VI-20
not all of the plutonium fueled rods vill be clad with Zircaloy as is indicated in Table IV-2 in Section IV. However, as a conservative upper limit of nine 72-rod assemblies clad with Zircaloy are assumed for this analysis. Using the nominal Zircaloy clad dimersions of Table II-2, and assuming a total of 648 rods, the total mass of Zircaloy in the core can be calculated as follows: n-2 2 .648 x ~T x (0 3911 - 0 3kh5 ) x 39 051 3 410lb/ft 1728 M = 161.6 lbs or 1 77 lb moles 14 tat,1 For the Zircaloy-vater reaction, the energy release is assumcd to be approximately 250,000 Btu /lb mole of Zr reacted due directly to the reaction. Additional energy may be obtained from the recombinatien of the ltrdrogen gas liberated during the reaction. This recombination energy is approximate 4 213,000 Btu /lb mole of Zr reacted. Thus, the maximum total energy release from such a reaction is approximately 820,000 Btu for 1 77 lb moles. If this amount of energy were added to the steam-air mixture of the vapor container that vould exist following the hypothetical accident, the pressure vould be raised by approximately 2 psi. Such a postulated instantaneous release is not physically possible because the core vould require some time to heat up and finally react following the loss of coolant. Because of the rapid decrease of the containment pressure following this accident, the added pressure rise due to the metal-vater reution could not cause the containment design pressure to be exceeded during the hypothetical. accident. The hydrogen recombination reaction can be violent if it occurs in a confined space and the concentration is at the explosive mixture lic.. The total amount of hydrogen released from the 100% metal-vater reaction l l VI-21 5. 9 L
is 3 54 lb moles of K2 gas r 7 1 lb of H. The dissolved hydrogen 2 in the reactor coolant in specified as a maximum of 90 cc at STP for each hilogram of reactor coolant. At ambient conditions, the primary coolant system volume contains approximacely 21,100 lb (9,600 kg) of water which would yield a maximum of 0.865 M3 of gas or approximately 77 6 gms at STP which is insignificant vhen compared to the amount released oy the metal-vater reaction. Some additiomi bydrogen is also available from the pressurizer steam bubble. Under normal operating conditions the maximum hydrogen content of the pressurizer steambubbleisapproximately500cc/kgofsteamcondensedto atmospheric conditions which yie]ds a hydrogen mas of approximately 7 6 gms. 3 The vapor container has a free volume of approximately 141,000 ft vhich would hold about 11,400 lbs of air at STP which is approximately 393 lb moles of air. Thus the ratio of lb moles of air in the contain-ment to lb molen of hydrogen released is approximately 110.so that the average concentration of hydrogen in the vapor container vould be about 1% by volume. Thishydrogenconcentrationisapproximately1/hofthe l lover flame limit and 1/7 of the lower limit of mixtures that could l explode. The rapid dispersal characteristics of hydrogen gas plus the large amount of turbulence that vill be associated with the rapidly cooling steam-air mixture in the containment vill most assuredly prevent any concentrated pcckets of hydrogen. The fission product activities that might be released to the containment as a result of the maximum hypothetical accident have also been re-evaluated. The inclusion of a partial core of plutonium does not appreciably alter the radiation sources used for the calculation of off-site doses in the Final Safeguards Report for Core I. As a result, the off-site direct whole body and inhalation doses for the maximum hypothetical accident vill be the same as presented in Section 600 of the Saxton Final Safeguards Report. I VI-22 i +
10' FUEL TEMPEMTURE COETa'ILIENT a -l.1 x 10-5 af.y ~ 10-5 HDDEMMR TEMPEMWRE Q5IFFICIENT = 0 3 x 10* &/*F 4 10 10*7 D 5 F2 10-8 -- 3 10*9 10"# -- .~. 10 20-leo 60 50 100 TIME, BE00NDS' COLD STAIN INCIMNT, 2 5 x 10~ Ak/8EC INBERTION 1 ,R. GURE VI - B - 1 ?~ em
__., _.__ _.~.___ _. _ _ _.. _. _. _ _ _ _ _.. _. _ _ _ _. _ _ _ _ _ _ FWL TD07.RATURE 00D7ICuffT = -1,1 x 10~5 &/*F 1 10 MODERAIDR TDG'ERATURE COEFFICIErf = 0 3 x 10 M/*F 0 2 10 10*1 = 10-2 10-3 I 4 9 i .10 10~0 I I 8 I 8 I I I 8 . J00 110 120 130-12 150 TIME,: SECORDB cold STARTUP INCI!WFF,' 2 5 x 10-k n/SEC IEEENTION FIGURE VI - B - 2 + 9 7g ww ev t' y7 g p yyg-ap*v9.--+9 y ; y e eQ-gywy 4.w.y, ggg. p-+ys' py.e7-u-1'w'sfe-r-WT-er--TMMT-yp---a r3 mm ge em w ftD wTw iy 3 &4y-w-+1r42e.wg-hmp.teisert c: rr-v W w G
RfEL TIMITRAnfE COETTICIINT = -1.1 x 10" &/*T M3DERA10R TEMPERARIE COETTICIENT = 0 3 x 10~ M/*F 1200 1000 ? 800 .a D 600 200 400 150 h - CIAD I Y. 8 100g 200 N. f -- _ WA'5R. U 1 i 0 8 I I 50 100 110 120 130 -- lic TIME, SECONDS COLD STARTUP INCIIENT, 2 5 x 16 &/SEC INSERTION FIGURE VI4-3 l l 1-
.. ~.... -. -... - - i FUEL TD07.RMVRE COEFFICIEC = -1.1 x 10*b Ar./*F MOIERMUR RMPERATURE C4 EFFICIENT.' O.3 x 10~ h/*F 120 i 100 / OO 60 [D (D kO 20 0 i I I 1 e i ~ ~ ~ i 100 110 120 130 140 150 TIME, seconds 001D STAJm7 INCIME,- 2 5 x 10~ &/8EC INSERTION ~ FIGURE VI-B h me + x. 4 i.
10 WEL TDFEPATURE CCDTICIENT =.1,0 x 10 tA/*F MODERATOR TTMPERAWRE COE7FICIENT = 2 7 x 10' tA/'T 10-5,_ 10 10 10'O D m O H b e 10 9 10 10 = 2 10'U O 20 M 60 60 100 i TIME, SECONDB BOT BTARIUP INCIIENT, 2 5 x 10 ok/SEC IEEERTION FIGURE VI-B-5 i 5)
0 10 -- FUEL TIMPEMTURE 00EF17CIENT =.l.0 x 10'N &/ *F MOIERAlt)R TEMPERATURE COETTICIENT = -2 7 x 10'4 Ak/*F ~ 10~1 10-2 y E B g 10~3 10" 2 10-5 i i i i i 100 110 120 TIME, SE03NIE HOT STAP5UP INCIIENT, 2 5 x 10~ ak/5EC IEERIION FIGURE VI-B-6 a
FUEL TDG'EMWIE COD 71CIENT =.1.0 x,10-5 gj.F 750 - MDIEMER'TEMPEMWHE 00 EFFICIENT = 2.7 x 10' Ak/*F W 700 I650 FUEL a 600 N U M - 550 8 550 g - c1An J ~ _, 500 1 1 8-525 95 100 105 110 US 120 125 _ TIME, 8E00128 HOT STARTUP INCIIENT, 2 5 x lo" &/SEC INSEREION FIGURE VI-B-T 4
i a -5 FUEL 'IEMPEMWRE COEFFICIENT = -1.0 v 10 gjy M)IEMTOR TEMPERTURE CO1FFICIENT = ~2,7 x 10* &/*F 20 10 p 4 0 1 I I I t 95 100 105 110 115 120 125 TIME, SECONDS HOT IfrARIUP INCIIENT, 2 5 x 10' A/SEC IIEiHTION FIGURE VI-B-8 ~ i 1 l t 4 ,.s - - - - +. ,s.,w. ,.n-,,,n .,,a,,ew-, ,.,,,.,,w,, e,p.,.y+ ,,p.,v.7,p g-w,w,,p y r. ....we
1 SCRAM INITIATED FUEL TEMPERAWRE COEFTICIENT =.1.0 x 10-5 3 77 120 MOIEhAMR TEMPERAWHE COKFFICITNT =.2 7 x 10 h/F' d 100 NUCIKAR FIUX HOT 8 POT [ HEAT FIUX 60 4-R + PRIMARY C00IANT g PRES 8URE . om g 7 R 10 20 A - aon - l ~ SCRAM COMPIETED 0 I i 0 5 10 15 20 TIME, SECONDS CONTINUOUS ROD WITHDRAWAL 3 (REACTIVITY INSERTION RATE = 2 5 x 10' &/see) FIGURS VI-B-9 (
i l l [ i 2000 N N m 1500 H d L7 m 1000 w h to h e i a 500 l I 0 I I i. 1 I t O 20 40 60 80 TIME, SECONDS l l i 1 STEAM BEAK ACCIIENT FIGURE VI-B-10 t ,/ I l 3-W2 w w "'em A Vw'w 9 9 s' ? -e 'yy4wi+
4-.A. A. A. s .J a. __da -.u Mas.aem Ap a 800 x H !? 600 M2 sa 400 b 2x LQ 'tA l 1 \\ l Q 140 tj h 100 I I 2 2 on .Hy 60 d g ~ x ~ o-140 Ob m m i i i i g 100 nim xy 60 M $3I =mbb 20 h* 0 I t h t 1 l o 20. 40 60 80 TIME, SECONDS S22AM EREAK ACCIDENT ( - FIGURE '/I-P-11 / c .y s-. .._.....m.,
N 600 m
- 8 b9 I
~ gg 500 g4 koo MhM 300 i '~ P 600 d b $; 50 s e EE" 300 8, y x 300 { i 1 s P 600 hE 500 E M E h 1.00 e d o i { l t i 600 NY m 500 =Mg I 400 3 8 300 do. i i I o go 60 60 TI}E, SECOE6 STEAM BREAK ACCIENT ( yIgunE VI-B-12 1
TOTAL BANK WORTH, ALL RODS Otfr 16 A 12 \\ h '\\ H t; 8 N'B g ~ \\ 4 L %,4 0 I I 0 a i 2000 1000 0 BOBON CONCENTFATION, FPM FIGURE VI-B-13 STEAM BREAK ACCIDENT, REACTIVITY INSERIION REQUIREMENTS A - Reactivity required in c ptrol rods to maintain a shutdown unrgiu ol' at least 0 5% Ok/k if one rod sticks upon scram. B - Reactivity required in egntrol rods to unintain a shutdown targin of at least, 0 5% Ok/k if no rods stirA upon scram. I h
100 FFIMARY 0001AVI TIDW COASTDOhd FOLli/41NG IOGS 0F PUMP FVWER l l 80 i 60 $L g l l ko. _. l t I i l eo l e o i I I I l i 1 0 1 2 3 4 5 6 7 l TDet, SECONE i FIGURS VI-C-1 .. ~.
I FUIL TDFEMTURE CCRTTICIErf = -1.15 x lo-5 6gjgfy. MERAIQL TEMPERATURE CCEFFICIEYt =-2 7 x lo Am/a/r-k- Scram Completion 3 Q-IMB RTIo (195o PBIA) scram Initiation i 2 - es-Ins mTro (soso Pau) D } I e l i l i l e i o 1 2 3 4 5 Tint, SE00NM IDS 8 0F FIDW ACCIIENT DiB RATICS YEMW TIE FIGURE VI-C-2 1 )
INSTANTANEOUS RELE ASE OF M AIN COOL ANT 30' CC:A AllA1ENT VESSEL PRf'SSURE VS =- TIME 25 - ll } 20 ?9 y 1 S 16 C cr i E g / 10 J 4 \\ s N N m 5 4 0 O 500 1000 1500-T IME.sec. FIGURE VI D 1 ( )
VII. SAFLTY C0!!OIDEPATIO!jS, A. .",;T:TICATION FOR INCLUSION OF 9 x 9 ASSDOLIES OF VIBFATIONELY COMPACTED TUEL IN SAXTON ILUf0NIUM PROGRIM INaluation of Defect Fotential Technicial feasibility of the vibrations 1 co:cpaction process has been der.?nstrated by satisfactory in-pile performance of a number of test samples and by irradiation of bulk quantities of fuel rods in the TRTE.(1' # 3' While a number of defects oc::urred during the early stages of these tests, the causes vere identified in almost every case and the net result has been increased confidence in the use of vibrationally corepncted fuels. This increased confidence ic reflected in the choice of such fuel for the EINR program. Under this program, 1,2% zircaloy clad fuel rods containing vibrationally compacted UO2' l.$ v/o PuO vill De irradiated t exposures of 15,000 to 20,000 WD/T. 2 Thirty tvo of the thirty-three defects which occurred in PRTR have been attributed to fluoride contamination, excess moisture content, and traces of oil introduced by faulty powder attrition apparatus.( 5,6,7,8,9) Tne defects occurred in both vibrationally co:cpacted and svaged Puo 'UU fuel rods. Th'e only defect which has not been explained to 2 2 date occurred in a svaged rod. Investigation of this defect is continuing. The ittpurities cited are now being controlled and the results have been considered in developing cpecifications for the Saxton Plutonium Program. Since control of these impurities was initiated, one hundred and fifty seven fuel rods containing vibrationally compacted Pu0 -UO fuel have been 2 2 irradiatedinPRTRtoexposuresofover1,000 MWD /t(peakat1600 MWD /T). Thee exposures are significant since all defects of vibrationally compacted rodspreparedunderoldtechnologyoccurredatless-than400 MWD /T. Also, 560 rods from this latter group are still in pile and have attained exposures of 6,000 MWD /T vithout a defect. I VII-1 %q -_.--______A___m_.-____
Chloride, moisture, bsd velds, and poor spacer design (resulting in fretting) were the enin causes of the five defects which occurred at 'll ravannah River during the early stages of the program. Two defects which occurred in svaged rods ( have not been explained to date and investigation is continuing. Since corrective measures vere taken, they have had no defects in vibrational 3y compacted rods. The rode have been irradiated to exposures of 9,000 MWD /T. Vnile these rods contained only UO, there is n rent, n t suspect different behavior 2 for 1%O -UO2 "i*t"#*** p fvaluation of Water Logging and Washout potential Daring the early stages of development of loose powder fuels, fuet vashout and waterlogging in the event of a defect vere considered ne possible perforrannee limitations. These conjectures vere based on (1) preliminary results of purposely defected fuel rods containing low density fuel (less than 85% T.D.) which vae Lore susceptible to vashout, and (2) the reporting of a possible waterlogging failure of a svaged UO2 fuel rod by Savannah River (details of this experiment are classified). (II' Apprehensions were greatly relieved when the former investigators reported the results of an unintentional defect in a svaged fuel rod containing UD at 66 T.D.(19) A hhough the longitudinal split was 1 5 inches long 2 and the reactor allcwed to run at full power for 15 hours after the defect was detected, only a small amount of UO2(amaximum f 10% of the fuel in the area of the split) was eroded out of the rod. Apprehensions were further relieved when out-of-pile tests at the General Electric Laboratories indicated that vibrationally compacted, svaged and rolled UO fuels had sufficient erosion resistance to prevent substantial losses 3 to the coolant. More recent results, however, have shown clearly that the potentials of these problems were greatly exaggerated during the early stages of development. No significant fuel vashout and no waterlogging failures ( VII-2
vere observed for the 33 defects with occurred in TRTE during the last two years.(2,5,6,7,0 Although cladding losses occurred in some cases, no severe reactor operating difficulties were reported. In some cases, the reactor underwent several pressura and power cycles after the defects were detected yet no vaterlogging failures occurred. I The Hanford PRTE resulto are confirmed by the experience at the Savannah River laboratories. They have never observed erosion of UO2 I#0* vibratory competed or svaged fuel rods although they experienced seven defects including the previously nentioned ossible waterlogging failure which had a 9 inch erack. (10,11,12,13,17,18 The erosion resistance of loose powder fuels results from high density peking coupled with in pile sintering. Evidence exists which indicates ) that in-pile sintering occurs at temperatures as low as 300'C. Other evidence of enhanced sintering in a radiation field has also been report-3' ed. These results may be explained by the following mechanism: (1) localized high temperatures resulting in increased rates for all sintering mechanisms (2) increased bulk and surface diffusion rates resulting from increased vacancy concentrations (3) enhancement of the vaporization-condensation mechanism through recoil processes. ?vo waterlogging - vashout type failures reported in 1962 should be mentioned since they caused some unnecessary anxiety. Neither failure can be considered applicable to the present situation. The first case reported was failure of a svaged, Mgo-PuO fuelrdexposedfor8 MWD /T 2 in the FRTR. Failure resulted from interaction between the Mg0 and water resulting in swelling and loosening of the fuel compact. No such reaction occurs in Pu0 'J0 fuels. The initial defect apparently resulted 2 2 from fluoride contamination and release of absorbed water from the Mg0. The second case reported vas. failure of a ovaged UO fuel r d undergoing 2 transient tests in SPERT.( The failure occurred during a 7 5 sec-period power excursion test in which fuel-temperatures rote by 300*C within ( VII-3 l l l
0.02 sec. It is certain that even a vaterlogged rod containing pellets voald burst under these conditions. The initial defects apparently resulted from broken epoxy resin seals used to insert eleven thermocouples into the center regions of the rod. The rod underwent several power excursion tests and remained in the reactor water for two days prior to the last test. Puel vashout occurred because of the large opening,12 inches long and up to 0 5 inches vide, and because the fuel was not in-pile long enough to sinter. The two cases cited cannot be employed to evaluate failure probabilities in loose powder fuelo. Similarly one cannot employ the waterlogging failure reported by the Bettis Laboratories to evaluate failure I potential in pelletized fuel. The failure in this case was attributed to low pellet density (60% T.D.) and reaction of water with uranium carbide contaminant in the U0 pellet. The initial defect was intentional. 2 SUWGRY i 1. Based on technology already developed at the national laboratories, no defects in vibrationally compacted fuel rods are anticipated. 2. The results of the national laboratory experiments shov that no significant fuel vashout and no vaterlogging results from defects in rods containing loose powder fuel. d 4 ( VII-4 mmm
P. OTEEATICN WITH DETECTITE WEL I As indicated in the previous section, there is a considerable amount of experience to indicate that clad failure under nornal Saxton operating conditions is highly improbable. In addition, the available experience and infornation demonstrates that even under thee fonditions of clad failure and exposure of the fuel to the coolant that fuel washout is also highly leprobable. However, even in the event of clad failure while the plutonium core is in Saxton, there would be no danger to the plant personnel or undue hazard to the public. Unless caused by some violent transient, that is much less probable than the accidents postulated and analyzed in sections VI-A and VI-B, clad failure vill be a gradual phenomena that would first nanifest itself by an increase in the activity level of the reactor coolant. This increate would be due to the release of the Bap activity of the failed rod into the coolant. If such an increase occurred, utilitation of the alpha detection equipment dascribed in Section V-B in conjunction with the nornal Saxton primary coolant sampling techniques vould be able to determine { the extent, if any, of the plutonium contamination of the coolant. If no l plutonium is detected and the ^ coolant activity level is below the limit in the technical specifications, the continued normal operation of the reactor is possible. Continuing alpha analysis of the coolant during subsequent operation vould detect any plutonium contamination increase. Significant plutonium con
- amination of the coolant voul'i indicate gross clad failure and associated fuel vashout n:.d would require that reactor be shut down and procedures instituted to clean up the system and locate and remove the failed fuel.
VII-5 ( ---.r~ m,,,n-o
RETERENCES 1. W. E. Foake, " Irradiation Alteration of Uranium Dioxide" Hanford IN 73072, March 1962. 2. " Quarterly Progress Report, Ceramics Research and Development Operation *, IN 81600 Janunty, February, March 1964. 3 J. J. Hawth, " vibration-Compacted Ceramic Puels" Nucleonics, Vol. 20, 9-1962, p. 50. h. D. F. Labcock. e t al., "An Evaluation of Heavy Water Moderated Power Eeactors", DP 830, March 1963 5 " Commercial Fabrication of Plutonium Puel", Hanford Laboratories Invitational Meetings, April 2-3, 1964, p. 15, 18. 6. " Quarterly Progress Report, Ceramics Research and Development" IN-76302, pp. k,5 7 Ibid, IN-7630L, pp. L.1, 4.20. 6. "Onelussified Research and Development Programs, Division Reactor Development", IN-80305, November 63 and IN-81651, April 64. 9 Private comunication, M. Treshly, S. Goldsmith, W. E. Roake. 10. Private comunication, A. S. Ferrara. 11. Heavy Water Moderated Power Reactors Progress Report, DP-830, March 63 12. Heavy Water Moderated Power Reactors Progress Report, DF-875, September-October 1963 , 13 Heavy Water Moderated Power Reactors Progress Report, DP-905, March-April 1964.
- 14. Heavy Vater Moderated Power Reactors Progreen Report, DP-895, January.
- February, 15 M. M, M111ho11en, A. R. Horn, J. L. Bates, "Hydriding in Purposely Defected Zircaloy-Clad Fuel Rods", IN 65465, 1961.
16. " Quarterly Progress Report, Fuel Development Operation" 1N-69085, pp. 5 31-5 39, July-September,1960. t 17 R. R. Hood and L. Isokoff, " Heavy Water Moderated Power Reactors Progress Report, US AEC Report DP-505, p. 38 March 1960. ( Yu-6
18. R. R. liood and L. Isoloff, Ibid, DP-525, pp.16-17, July 1%0. 19 M.1;. Millhollen, A. R. llorn, J. L. Intes, " Erosion Resistance of Svaged 00 Following an in-Reactor Fuel Rod Cladding Tailure", IN-73015,1%1. i 7 20. J. W. Lingarelter, E. A. Lees, R. J. Seely, GEAP-4020, 1962. 21. C. H. Spalaric, F. A. Compre111, M. Siegler, GEAP-3698. 22. "Qunrterly Progrecs Report, Ceramics Research and Development", HW-76304 pp. L.10, October-Iecember 1962. 23 E. A. Aitken, "S$ntering Characteristics in a Radiation Environment", ASTM Meeting, June 28-29, 1961. 21+. W.1:. Iarney, B. D. Wemple, Meta 11ography of Irradiated UO Containing 2 Fuel Elements", l'APL 1836,1958. 25 " Quarterly Progress Report, Fuel revelopment Operation", HV 76300 October-Icee:tber 1962. 26. J. E. lloughtaling, T. M. Quegley, A. H. Spana, " Calculation and Measurement of the Trancient Tet::perature in a Low-Enrichment UO Fuel Rod During Large 2 Power Excursions", IDO-16773,1962 27 J. D. Eichenberg et al., " Effects of Irradiation on Eulk UO '" US AEC Report 2 kAFD-183, Westinghouse Electric Company, October 1957 28. J. J. Havth, " Vibration-Compacted Ceramic Puels" Nucleonics, Vol. 20, September 1962, p. 50. 29 V. W. Storhoh, " Fabricating Plutonium for IWtter Performance", Nucleonics, Vol. 21, January 1963, p. 38. 30. United Strates-Farutm Joint Research and Development Program Report - EVRAEC 590, October 1962. 31. F. G. Davcon, " Plutonium Ac A Power Reactor Fuel", Hanford HW 75007, September 1962. VII-7 %v
VIII. CONCUJSIONS Based on the preceding report, it mny be concluded that the propoced partial plutonium can be safely installed and operated as Core II in the Saxton reactor under chemical shim operating conditions or non-chemical shim conditions. The operating history of Saxton to date plus the results of the che.tical shim experiment have clearly demonstrated the feasibility of such operation. The occident armlyses performed show that the results are no vorse than for the previously analyzed Sarton Core I and that the present Saxton control and safety systems are more than adequate for une with the partial plutonium Core II. VIII-1 . -.. -. -... - _ -.,}}