AECL-7615
ATOMIC ENERGY |S9 L'ENERGIE ATOMIQUE OF CANADA UMITED Vj&J DU CANADA LIMITEE
REACTOR PHYSICS AND ECONOMIC ASPECTS OF THE CANDU REACTOR SYSTEM
Aspects techniques et 6conomiques de la filter© CANDU
J. GRIFFITHS
A series of lectures presented in Indonesia 1981 November 9-13
Chalk River Nuclear Laboratories Laboratoires nucleaires de Chalk River
Chalk River, Ontario
February 1983 fevrier ATOMIC ENERGY OF CANADA LIMITED
REACTOR PHYSICS AND ECONOMIC ASPECTS OF THE CANDU* REACTOR SYSTEM
Compiled by
J. Griffiths
A series of lectures presented in Indonesia 1981 November 9 - November 13
*CANDU is a registered trademark of Atomic Energy of Canada Limited
Chalk River Nuclear Laboratories Chalk River, Ontario, KOJ 1JO 1983 February
AECL-7615 L'ENERGIE ATOMIQUE DU CANADA, LIMITEE
Aspects techniques et économiques de la filiere CANDU
compilés par
J. Griffiths
Série de conférences données en Indonésie du 9 au 13 novembre 1981
Le présent document renferme les textes de base des neuf conférences données à l'Ecole nationale de physique de YogJakarta, Indonésie, du 9 au 13 novembre 1981. On y trouve l'histoire du développement de la filière CANDU ainsi qu'une description assez détaillée du réacteur CANDU de 600 MWe. Ces textes décrivent les méthodes de calcul de la physique des réacteurs et ils comparent les paramètres calculés â ceux mesurés dans les réacteurs de recherche et de puissance. Ils examinent les aspects économiques de la filière CANDU dans le contexte d'Ontario Hydro et ils font des comparaisons avec Tes centrales nucléaires 3 eau légère et les centrales alimentées par des combustibles fossiles. Certains aspects techniques et économiques des réacteurs CANDU alimentés en uranium faiblement enrichi ou en thorium sont passés en revue. Enfin, le programme de R & D touchant les cycles de combustible avancés est brièvement décrit. Ces conférences s'appuyaient sur les données alors disponibles. Le statut des études relatives aux cycles de combustible avancés change constamment. Par ailleurs une partie des données fournies dans ce domaine n'avait qu'un caractère illustratif.
Laboratoires nucléaires de Chalk River Chalk River, Ontario, KOJ 1J0 Février 1983 AECL-7615 ATOMIC ENERGY OF CANADA LIMITED
REACTOR PHYSICS AND ECONOMIC ASPECTS OF THE CANDU TtEACTOR SYSTEM
Compiled by
J. Griffiths
A series of lectures presented In Indonesia 1981 November 9 - November 13
ABSTRACT
This document contains the material presented In nine lectures at the National School of Physics at Yogjakarta, Indonesia, 1981 November 9 to November 13. A history of the development of the CANDU system Is given along with a fairly detailed description of the 600 MW(e) CANDU reactor. Reactor physics calculation methods are described as well as comparisons between calculated reactor physics parameters and those measured in research and power reactors. An examination of the economics of CANDU in the Ontario Hydro system, and a comparison between fossil fuelled and light water reactors is presented. Some physics, economics and resource aspects are given for both low enriched uranium and thorium fuelled CANDU reactors. Finally the R&D program in Advanced Fuel Cycles is briefly described.
These lectures were based on material available at the time. Perspective on advanced fuel cycles is continuously changing and some of the data in this area was presented for illustrative purposes only.
Chalk River Nuclear Laboratories Chalk River, Ontario, KOJ 1J0 1983 February
AECL-7615 TABLE OF CONTENTS
PAGE
A. THE CANDU-HHW NATURAL (7RANIUM REACTOR 1
A.I DEFINITION AND HISTORY 1
A.2 PHYSICAL DESCRIPTION 7
A.3 PHYSICS ASPECTS OF THE NATURAL URANIUM CANDU 10
Tables 57
Figures 64
B. THEORETICAL LATTICE PHYSICS 97
B.I INTRODUCTION 97
B.2 THE PHYSICS OF SOLACE 97
B.3 COMPARISON OF LATREP WITH EXPERIMENT 123
Table 124
Figures 125
C. CANDU ECONOMICS 132
C.I INTRODUCTION 132
C.2 COST CRITERIA 132
Tables 150
Figures 154
D. THE LOW ENRICHED URANIUM FUEL CYCLE 159
D.I SOME EFFECTS OF ENRICHMENT , 159
D.2 AXIAL POWER DISTRIBUTIONS 161
D.3 REACTIVITY COEFFICIENTS 162
D.4 IMPACT ON ECONOMICS AND RESOURCE UTILIZATION 162
Tables 167
Figures 172
E. THORIUM FUEL CYCLES IN CANDU 184
E • 1 INTRODUCTION 184
E.2 SOME FEATURES OF THORIUM FUEL CYCLES 185 TABLE OF CONTENTS (cont'd)
PAGE
E.3 SOME REACTOR PHYSICS PROBLEMS IN THE THORIUM CYCLE 187
E.4 REACTOR PHYSICS ASPECTS 193
Tables 196 Figures 200
F. RESOURCE UTILIZATION AND ECONOMICS WITH THE THORIUM
SYSTEM 212
F.I THORIUM CYCLES IN CANDU REACTORS 212
F.2 BURNUP ESTIMATES 215
F.3 RESULTS 222
Table F-2 223
F.4 CONCLUSIONS 224
Tables 229
Figures 232
G. THE AECL ADVANCED FUEL CYCLE PROGRAM 248
G.I INTRODUCTION 248
G.2 FUEL DEVELOPMENT 248
G.3 FUEL REPROCESSING 250 G.4 REACTOR PHYSICS 253
G.5 ASSESSMENT 261
G.6 SUMMARY 263
Tables 264
Figures 266
BIBLIOGRAPHY 268
ACKNOWLEDGEMENTS 269 A. THE CANDU-PHW NATURAL URANIUM REACTOR
A.I DEFINITION AND HISTORY
A CANDU* reactor is a heavy-water-moderated power reactor of the general type developed in Canada by Atomic Energy of Canada Limited (AECL).
It is somewhat an accident of history that Canada focussed all her nuclear power development effort on heavy water reactors.
The Canadian Nuclear Program had its origins almost 40 years ago when Canada, in collaboration with the U.K., undertook the develop- ment of heavy water moderated reactors. The result of this research was the design, construction and operation of a series of heavy water moderated reactors which established Atomic Energy of Canada Limited (AECL) as the world authority on scientific and technical knowledge in this field.
The first reactor to go critical outside of the United States was a heavy water reactor called ZEEP (Zero Energy Experimental Pile). This reactor, which went into service in 1945, is located at the Chalk River Nuclear Laboratories in Canada and provided basic experimental physics data for the development of CANDU.
In 1947, NRX (National Research Experimental), the highest flux reactor in the world at that time, went into service at Chalk River. This reactor was moderated with heavy water as was its higher power successor, NRU (National Research Universal), which became operational in 1957. ZEEP, NRX and NRU provided the fundamental
*CANDU signifies a Canadian design using jleuterium oxide (heavy water) as moderator and uranium as fuel. - 2 -
physics, chemistry, metallurgy, and engineering knowledge for the CANDU program. Other research reactors, PTR (Pool Test Reactor) and ZED-2 (Z-2, second zero energy reactor), have also contributed valuable data and information as indicated in Figure A-l.
In the early 1950's, a Canadian nuclear power program was being seriously considered by AECL. By 1954, a Nuclear Power branch was established at Chalk River. They considered the British Gas Cooled Reactor and the American Light Water Reactor, but concluded that the Heavy Water Moderated Reactor was most suitable for Canada. The system chosen exploited the merits of heavy water as a moderating material. The use of pressure tubes in place of pressure vessels for the coolant matched our national manufacturing capability and the use of natural uranium as fuel allowed the direct use of Canada's indigenous resources of uranium.
The Nuclear Power Branch laid down the conceptual design for NPD (Nuclear Power Demonstration), 22 MW(e), which has been in operation since 1962 and continues to generate not only electricity but knowledge for the CANDU program as also is indicated in Figure A-l.
The group also engineered the basic concept for a larger 200 tIW(e) plant which evolved into the Douglas Point Station that began operating in 1967. Pickering A, KANUPP in Pakistan and RAPP 1 in India were committed on the basis of design and development lessons learned from NPD and Douglas Point, and these all first produced electrical energy in 1971/72.
From Pickering A, 4 x 515 MW(e), we went on to the larger Bruce A Station, 4 x 740 MW(e), and the even larger Darlington Station, 4 x 881 MW(e), to meet the demands of our most populous Province, Ontario. The confidence we gained from the success of these evolutionary steps has led to the duplication both of the Pickering Station (Pickering B) and of the Bruce Station (Bruce B). It is - 3 -
expected that the Darlington Station also will be duplicated when the demand requires.
From the outstandingly successful Pickering A units we evolved the 600 MW(e) Gentiily 2 unit for the Province of Quebec and duplicated this for the Province of New Brunswick at Pt. Lepreau.
This unit was adopted as the standard export design which would most likely suit the requirements of our potential customers. The viability of this decision has been borne out by the sales of similar units to Argentina for Cordoba, to Korea for Wolsung and to Romania for Cernavoda.
Canadian export experience has extended over a wide range of scopes of supply from virtually total plant turnkey operations in India, Pakistan and Korea, through Nuclear Steam Plant supply in Argentina to licensing, engineering and procurement services in Romania.
At Cordoba in Argentina, our first 600 MW(e) export order, we contracted to supply the Nuclear Steam Plant but our contract became unworkable, largely as a result of unpredicted inflation in the client country, and this was severely detrimental both to our client and to ourselves. After renegotiation of our contract, with Argentine contractors performing a major role in the project, the work is proceeding smoothly and the plant is expected to be. in service by 1982.
At Wolsung-1 in Korea, our second 600 MW(e) export order, we contracted essentially a turnkey operation with minimal participa- tion by our client. The productivity of the Korean workers has been phenomenal and we expect that plant also to come into service in 1982.
Table A-l lists all CANDli power reactors in operation, under construction or committed in Canada and around the world. CANDU has often been referred to as a physicist's reactor, due to the scientific elegance of the concept and its dogged adherence to fundamental principles. This reference is used by some in a derogatory way, but the current economic competitiveness of CANDU with other types of power reactors ensures that it has appeal in the hard commercial world as well as merely to physicists. The excellent performance of the CANDU system is demonstrated in Table A-2, which shows the cumulative load factor of the world's twelve best reactors up to the end of September 1980, and in Figure A.2, which shows the load factor for all major reactor types. It is interesting to speculate as to whether this illustrates a general principle. In our complex modern world it is often difficult, if not impossible, to anticipate at an early stage in the development of a project just how well the final product will compete. By observing fundamental physical principles (e.g., neutron economy) the chances of eventual success can perhaps be maximized.
Before describing in more detail a typical CANDU reactor system, a few of the important distinguishing characteristics and features will be highlighted. Figure A.3.
Neutron Economy
While the CANDU concept can be adapted to enriched fuelling, the original target was a system which could be operated competitively using natural uranium as the fuel. In order to achieve this objective, a great deal of attention had to be paid to neutron economy in the choice and use of core materials and in the mode of operation.
Pressure Tubes
The CANDU concept is based on the use of pressure tubes rather than a pressure vessel. - 5 -
The reactor consists of an array of pressure tu'jes., generally arranged in a square lattice, which pass from end to end through a large cylindrical tank. The reactor fuel resides inside the pressure tubes and the coolant is pumped through then to cool the fuel. The fact that the coolant is at high pressure gives rise to the term pressure tube.
While it would be possible to use a pressure vessel concept similar to that used for a PWR, the vessel required would be considerably larger, because the required volume of D2O moderator is much larger tlian the volume of H20 moderator in a PWR. In fact, early development of the CANDU concept was based on a pressure vessel approach. Since the limit on size of pressure vessel that could i. = been manufactured in Canada at the time would have placed a limit on unit capacity, and since the development of & low neutron cross section zirconium alloy had proceeded to the point where this material could be used for pressure tubes, the pressure tube concept was developed and adopted.
Experimental evidence indicates that the pressure tubes will leak before they break since their thickness is much less than the critical crack length.
Separate Moderator
The heavy water moderator is held in the large cylindrical tank, called the calandria, that surrounds the pressure tubes.
Because the coolant, and hence the pressure tubes, mvt operate at high temperature in a power reactor, and because it is desirable to operate the moderator at low temperature, the pressure tubes must be insulated from the moderator. This is done by introducing a second tube surrounding the pressure tube but separated from it by a stagnant gas space. This second tube, the calandria tube, is sealed - 6 -
at both ends to the calandria end plates or tube sheets, thereby completing the moderator containment. With this arrangement the fuel and coolant is completely separated from the moderator, permitting a choice of coolants.
Leaking pressure tubes can be detected by monitoring Che moisture content and pressure of the gas in the sp=>ce between the pressure and calandria tubes.
There are several advantages of operating the moderator at low temperature, as follows:
(a) the calandria can operate at atmospheric pressure, avoiding the need for a heavy, high-pressure vessel.
(b) the cold moderator can act as a valuable heat sink under certain accident conditions,
(c) since the moderator is cold it cannot add to the energy to be controlled under accidert conditions, and
(d) the overall neutron economy is better.
On-Power Fuelling
The pressure tube approach lends itself to on-power refuelling, since the fuel residing in individual pressure tubes can be changed without affecting the other pressure tubes or the fuel in them. Or-power refuelling maximizes the attainable burnup, since only a nominal excess reactivity is required at equilibrium. The reactivity gained by the continuous replacement of old fuel by new compensates for the general decrease in reactivity with fuel irradiation. - 7 -
A. 2 PHYSICAL DESCRIPTION
In order to appreciate the reactor physics of CANDUs it is helpful to have a physical picture of the system. A fairly complete description of the 600 MW(e) CANDU-PHW will be given.
The CANDU-PHW is the pressurized heavy trater cooled version of CANDU. To date this version has always employed a horizontal reactor core orientation (i.e., the pressure tube axes are horizontal). Vertical orientation has been studied but no Incentive to switch has been identified. For completeness a schematic arrangement of the whole system is shown in Figure A.4.
Pressurized heavy water coolant at e pressure of 9.99 MPa ( 1449 psi) and a temperature of 266°C is supplied to each fuel ch nel via an individual pipe, called a feeder pipe. As the coolant passes through the fuel channel it picks up heat from the fuel and leaves the channel at a temperature of about 310°C. It is coaveyed to the outlet header via the outlet feeder pipe. From the outlet header, the coolant flows through the steam generators where it is cooled back to 266°C, its heat being given up to produce steam at about 4.69 MPa (660 psi) which is fed to the turbine. The coolant then enters the primary pumps which deliver it to the inlet header and thence to the inlet feeder pipes. There are two circuits with 190 fuel channels in each circuit. Each circuit contains two pumps, two inlet headers and two outlet headers in a figure eight arrangement. The flow through the channels is bidirectional and the feeders are sized such that the coolant flow to each channel is proportional to the channel power. One advantage of this figure eight scheme is that in the event of a primary pump failure the coolant flow is maintained at about 70% of the normal value.
A separate auxiliary circuit is employed to circulate the heavy water moderator through external heat exchangers. A more detailed - 8 -
view of the major reactor components is shown In Figure A.5. There are 380 fuel channels spaced on an 28.575 cm square lattice.
All the reactivity and control mechanisms penetrate the reactor from the top of the calandria as shown in Figure A.6. In the CANDU reactor the primary method of controlling reactivity in the long term is through the on-load refuelling capability. This eliminates the need to design reactivity control devices having large reactivity worth.
The light water zone control absorbers are the primary reactivity control devices. They act in unison for bulk reactivity control and differentially for flux tilt control. Six tubes within the reactor core contain fourteen compartments into which light water is introduced. The difference in reactivity between all compartments full and empty is 7.5 mk and the maximum rate of change of reactivity is 0.14 mk/s.
Four mechanical control absorbers are provided. They can be driven in or out at variable speeds or dropped under gravity. These absorbers are normally out of core and are used to supplement the zone controllers. The reactivity worth of these absorbers is about 10 mk. These absorbers along with the zone controllers are capable of shutting down the reactor.
To extend the range of the reactor regulating system in the positive direction beyond that available with the zone control system the reactor is designed to operate with a group of absorber rods fully inserted. This system of twenty-one stainless steel rods is called the adjuster rod system. This system is designed to have sufficient reactivity to compensate for the increase in the Xel35 absorption that occurs within thirty minutes following a reactor shutdown. Since the adjuster rods are nornally fully inserted the distribution of absorbing material amongst the rods is chosen to flatten the - 9 -
power distribution. The twenty-one adjusters are grouped in seven banks not all having the same number 01 rods. The banks are chosen so that the reactivity worth of any one bank does not exceed the range of the zone control system. The reactivity worth of the twenty-one rods is about 15 mk and the maximum rate of change is about O.J. mk/s.
Twenty-eight spring assisted gravity fall rods, similar to the mechanical control absorbers are provided as the primary protection shutdown system. They have a total reactivity worth of about 80 ink. They are augmented by a moderator poison system which injects at high speed a gadolinium solution into the calandria.
Figure A. 7 is an illustration of one of the 380 fuel channels. The approximate length of the pressure tube is 6.3 ID and the minimum internal diameter 103 mm. Fuel for the reactor, Figure A.8, is in the form of fuel bundl3S.
Each bundle is 49.5 cm (19.5 inches) long and consists of 37 elements containing compacted and sintered pellets of natural UO2 in a thin Zircaloy sheath. The elements are about 13 mm in diameter and are attached mechanically at their ends to form a bundle about 102 mm in diameter. A small space is maintained between the elements by spacers attached to the element cladding. The bundle is about 92 wt.% UO2 and 8 wt.% Zr. The structural material accounts for only 0.7% of the thermal neutron cross section of the bundle, to give a fuel assembly that is highly efficient in its use of neutrons.
Loading of new fuel into the reactor and removing spent fuel is carried out on-power by two co-ordinated fuelling machines (one at each end of the reactor) controlled from the station control centre. Fuel bundles are moved in opposite directions in adjacent channels, referred to as bidirectional fuelling, to achieve an axially symmetric neutron flux distribution. The coolant flows in the same direction as the fuel movement on refuelling and hence flows in - 10 -
opposite directions in adjacent: channels. There are twelve fuel bundles per channel and any ewn number of bundles can be replaced during one refuelling visit.
A.3 PHYSICS ASPECTS OF THE NATURAL URANIUM OANDU
Critical Size and Reactivity
Measurements of the buckling of cold clean lattices have been made for a variety of fuel geometries in the ZED-2 reactor. A typical set for 28-eleraent fuel of the type that Is used in the Pickering reactors is shown in Figure A.9 as a function of hexagonal pitch fcj two coolants. The buckling is seen to exhibit a maximum of about
2.8 m~2 at a hexagona? pitch of about 30.7 cm which corresponds to the 28.575 cm square pitch in Pickering. At all pitches the lattice having voided coolant has the greatest buckling indicating that the coolant void reactivity coefficient ±r, positive. The optimum buckling for the 600 MW(e) CANDU which was described previously is about 3.3 m~2. This means that a bare critical cylinder with minimum volume would have a radius of about 1.6 m and a height of about 3.0 m giving a volume of 25 m^.
The actual reactor units are, of course, not simple bare cylinders. The 6-m-long fuel rods are enclosed in the calandria which has a radius of 4m over most of its length. (There is a small step decrease in calandria radius at both ends.) The 380 fuel channels on a 28.575 cm square lattice pitch gives an effective core radius of 3.143 m. The space between this radius and the calandria is occupied by heavy water to form a radial reflector. The geometric
buckling of th.'s arrangement is 0.65 m~2. por a cylinder of length 6 m, this buckling corresponds to a volume of about 290 m3 compared to the actual volume of about 190 tn^. - 11 -
This extra effective volume provides about 10.3% (Sk/k) in excess reactivity above cold, clean critical. This excess reactivity is allocated approximately as in Table A-3.
During the period 1960-1969, ZED-2 was used mainly for studies of lattices in the range of interest of the Canadian power reactor program. Lattices were issembled at a variety of lattice spacings and with air, D2O, H2O and HB-40 coolants using the following fuel assemblies:
- 7-element UO2 - 19-element UO2 - 28-element UO2 - 19-element U-metal - 7-elemsnt UC - 19-element TI1O2-UO2 - 31-element simulated irradiated fuel.
Since that time further lattice experiments have been done with rods having other geometries. The most notable example is a series of experiments carried out with 37-eleraent rods of UO2, U3Si, UC and U-metal made to identical geometries. The measurement techniques have improved with time so that the later experiments have been done with substitution techniques using as few as seven test rods, rather than with full-scale lattices.
Measurements were made in these lattices that give information on the critical size of the lattice in any geometry, the ratio of fissions in fertile isotopes (caused by high energy neutrons) to fissions in fissile isotopes, the ratio of fissile material produced (by neutron capture in fertile isotopes) to fissile material destroyed (by fission or neutron capture), and the shape of the neutron energy spectrum. Measurements were also made of the spatial - 12 -
distribution of neutron flux within channels, lattice cells and the lattice.
For full lattices, bucklings are obtained by a flux mapping technique using manganese activation measurements.
The homogeneous two-group diffusion equations for the fundamental
radial distributions of the thermal neutron flux flux >th(r) = A'JQ(Ar) + C'ygr) (1) f(z) = 4>(ZO> cos a(Z - zQ) (3) where A', C = amplitude coefficients S, S1 = fast-thermal coupling coefficients which depend on the core properties 2 A = radial buckling a = axial buckling = -TZ— rl ex H = extrapolated height - 13 - and 32 = i- + i— + 2a2 + X2 (4) L Ls The total geomttrlcal buckling for a finite cylindrical reactor is B2 = a2 + X2 (5) For a neutron detector whose activation cross section can be expressed as Stot«l where E . and Zc are activation cross sections averaged over the thermal and fast neutron flux distributions, respectively, the total activation Act(r.z) is r a + Act(r.z) = Substitution of (1) and (2) in (7) and neglecting the axial variation gives the radial variation of the induced activity Act'(r) as Act'(r) = (Zth+SZf)A'Jo(Xr) + (Eth-S'If)CIo(er) (8) Equation (8) can be expressed in the general form Act'(r) = A Jo(\r) + C IQ(8r) (9) - 14 where the magnitude and sign of C will be dependent on the detector parameters and the core properties. Equations(3) and (9){ valid for homogeneous systems, can be applied to a heterogeneous reactor if the neutron flux is measured at identical positions in each cell so that the macroscopic distribution is not distorted by the microscopic distribution. This implies separability of macroscopic and microscopic flux variations. Total bucklings are determined by measuring neutron flux distributions throughout the reactor core and fitting the measured radial and axial distributions to equations (9) and (3), respectively. Generally,points near the reactor boundaries are omitted in the fitting process. Because of the expense of providing whole cores of fuel it has become the practice to determine bucklings by means of substitution experiments. In this type of experiment the central few channels, typically seven, of a driver core are progressively replaced by the test fuel and the changes in critical height are used to estimate the buckling of a whole core of the test fuel. Two-grocp two-dimensional heterogeneous calculations are made using the MICRETE code. In this code the fuel rod is treated as a line sink of thermal neutrons. The spatial distribution of thermal neutrons due to the presence of this sink is determined by diffusion theory,. The absorption of thermal neutrons in the fuel rod results in a line source of fast neutrons whose spatial distribution is also calculated by diffusion theory. The loss of fast neutrons leads to a production of thermal neutrons in the moderator. Resonance absorption in the rod is assumed to be proportional to the.fast flux at some radius from the rod axis to make some account for the finite size of the fuel rod. The constant of proportionality is obtained by assuming a flat macroscopic flux and taking the resonance - 15 - absorption equal to (1-p), where p is the resonance escape probability, times the fast neutron absorption in the cell. The decrease by resonance absorption of the number of neutrons reaching the thermal group is taken into account by introducing a uniform negative source of thermal neutrons from the rod axis to some radius. The buckling of the driver core is obtained by flux mapping. MICRETE is forced to fit these measurements of buckling and the critical height by adjusting the extrapolation lengths, and the resonance es-ape probability(p). The core with the test rods is then set up with these extrapolation lengths, the driver fuel having the previously determined value of p, and p for the test fuel is adjusted until the measured critical height is reproduced. Finally, a lattice composed entirely of test rods is simulated, using the value of p previously determined, and the reactor size adjusted to achieve criticality. The buckling of the test lattice is then obtained from the calculated flux shape. The method has been tested with several D2O cooled lattices in which the bucklings of the test and driver lattice differed by as much as 7.75 m~2. Seven rod substitutions are found, for these lattices, to produce bucklings for the test fuel which agree with the flux map measurements to within +0.2 m~2. The discrepancy is larger, up to +0.5 m~2)when other coolants are used in the test lattice, which may indicate the onset of spectrum mismatch problems. Using only one test rod rather than seven results in larger discrepancies, Figure A.10 is a plot of the discrepancy for one and seven rod substitutions as a function of the moderating ratio. During the commissioning of the Bruce reactors, experiments were performed to verify the critical size and reactivity prediction. - 16 - Approach to Initial Critlcality The initial fuel loading of a Bruce reactor consisted of 5808 natural UO2 and 432 depleted fuel bundles. Prior to start-up, the reactor was in a guaranteed shutdown state. To ensure this state, a solution of B2O3 was added to the moderator so that the reactor was subcritical by at least 130 ink. The approach to initial criticality was done by removal of boron from the moderator system using ion exchange columns. During the approach to critical the neutron flux was monitored using sensitive neutron detectors. The estimated time of criticality was determined by plotting the reciprocal count rate against purification time. The boron concentration was also constantly measured since a prediction of this parameter at criticality had been made. The variation of reciprocal count rate with boron concentration is shown in Figure A. 11. At criticality, the measured boron concentration was 8»0 +_ 0.5 mg B/kg D20. The uncertainty includes estimated error in the sampling and chemical measurement techniques. The predicted critical boron concentration was 8.3 mg B/kg D£0 which agrees very well with the measurement. Calibration of the Light Water Zone Control System The light water zone control system is designed to perform two main functions, namely, (1) to pruvide short-tarm reactivity control to maintain reactor power at demanded level during normal operation, and (2) to control spatial power distribution by suppressing - 17 - regional overpower transients associated with reactivity perturbations. The system consists of fourteen in-core platinum self-powered neutron detectors and fourteen independently variable zone controllers. There are six vertical Zircaloy • tubes running through the reactor, with four of the tubes divided into two compartments and two into three for a total of fourteen compartments. During the low power commissioning, the system operated with no spatial control. The calibration consisted of initially raising all the zone levels to ^90% full by withdrawal of boron from the moderator. When the zone levels were stable, an accurately weighed amount of B2O3 solution was carefully added to the moderator system. To maintain the reactor power, the regulating system compensated for the added boron by lowering the average zone level. Thus the change in average zone level was equal to the reactivity worth of the added boron. This procedure was repeated until the average zone level reached ^10% full. The results of the calibration are shown in Figure A.12. By extrapolation to the 0% and 100% level, the total worth of the zone +0 S system is measured to be 5.4 f.*1 ink The calculated worth of 5.9 mk is in reasonable agreement with the measured worth. Calibration of the Mechanical Control Absorber Rods The purpose of the four mechanical control absorbers is to provide negative reactivity foe rapid power reductions, beyond the capability of the light water zone controllers. These absorber rods - 18 - are made of stainless steel sandwiched with cadmium. Insertion and withdrawal from the core is through vertical Zircaloy guide tubes located interstitially between channels. The reactivity worths of these absorbers were determined indirectly by noting the change in average light water zone level when an individual absorber or bank of absorbers were inserted in the core. The results are given in Table A-5 and show good agreement with predictions. Calibration of Shutoff Rods The purpose of the shutoff rods is to shut down the reactor rapidly when certain trip parameters are exceeded due to a fault condition. The structure of the shutoff rods is identical to the control absorbers. The individual rod worth was determined using the same procedure used with the mechanical control absorber rods. The total reactivity depth of the shutoff rods was also measured to determine the overall worth when all rods were inserted. The procedure required inserting all the rods while maintaining the reactor critical by boron withdrawal from the moderator system. After all the rods were inserted, individual rods were withdrawn while adding accurately measured batches of boron to the moderator. The reactivity worth of the total amount of boron added to the moderator required to withdraw all the rods is equivalent to the reactivity depth of the shutoff rods. During this calibration, shutdown capability was provided by the second shutdown system (Liquid Injection System). The results are shown in Table A-5 and show agreement with predictions to within 5%. - 19 - Calibration of the Booster Rods The booster system is designed to provide positive reactivity to compensate for xenon build-up after a power reduction or to compensate for reactivity loss due to burnup as a result of unavailability of the on-power fuelling system. Each booster rod contains uranium which is highly enriched with U235• There are 16 booster rods arranged throughout the reactor core. As with the control absorbers, and the shutoff rods, the individual booster rods were calibrated against the zone controllers. The total reactivity worth of the boosters was determined by individually inserting all the rods while keeping the reactor critical by adding accurately weighed amounts of boron in the moderator system. When all the booster rods were inserted and the average zone level matched the initial average zone level, the total reactivity worth of the booster rods is equivalent :.o the worth of the boron added to the moderator. Typical results are given in Table A-5. Agreement is within 5%. Neutron Balance In order to get a reasonable burnup with natural uranium fuel, great care must be taken not to waste neutrons. A 0.1% increase in available reactivity would allow a burnup increase of about 125 MW.d/t. The approximate neutron balance in an equilibrium core is summarized in table A-4. It is immediately evident that a very significant - 20 - fraction of the power comes from fissions in the fissile plutonium isotopes. Fuel Burnup Figure A.13 is an estimated curve of reactivity change as a function of uniform burnup for a Pickering lattice. Since there is nominally 3.0% Sk/k in reactivity available to compensate for burnup effects, we can see from this figure that, if the burnup were uniform throughout the core, the excess reactivity could sustain a burnup of ^4500 MW.d/*-TJ However, since the reactors ar > fuelled almost continuously on-power, there is always a range of burnups in the fuel in the reactor - running from fresh fuel to fuel ready to be discharged. The fresher fuel helps to compensate for the extra neutron requirements of the fuel near discharge burnup. It turns out that a very good estimate of discharge burnup can be obtained by integrating the curve in Figure A.13 to a point such that the area under the curve is zero - as shown in Figure A.14. This point gives an estimate of the average discharge burnup. Of course the achievable discharge burnup depends on the fuel management scheme used and this topic will be discussed in some detail later. Very good measurements of fuel burnup in the Pickering units are made by keeping account of the fuel consumption as a function of total cumulative heat produced. The current burnup estimates from up-to-date measurements of this type are: for Pickering Units 1 and 2 - 175 MW.Wkg for Pickering Units 3 and 4 - 190 MW.h/kg - 21 - A fairly sensitive check on the accuracy of burnup calculations can be made if the composition of discharged fuel can be determined. Eight fuel bundles discharged from the 25 MW(e) CANDU, NPD, have been analyzed for nuclide content. The average burnup of the bundles ranged from about 980 MW.d/t to 10800 MW.d/t. NPD fuel consists of a bundle of 19 fuel pins arranged in three rings containing 1, 6 and 12 fuel pins. Mass spectrometric techniques were used to determine the relative urani'un isotope concentrations and the relative plutonium isotope concentration while the relative amounts of plutonium to uranium were obtained by isotopic dilution measurements. The fuel composition in terms of nuclide ratios was determined for each ring of fuel pins and for the bundle average. The measured fuel composition in the form of nuclide ratios have been compared with calculations made using the AECL lattice cell and fuel depletion code LATREP. Figure A.15 is a comparison of measured and calculated bundle average nuclide ratios plotted against the burnup indicator (1-ot) where a = R(B)/R(O) and R(B) is the ratio of the number of TI235 atoms to U238 atoms at burnup B Table A-6 indicates the mean error in the nuclide ratios over the eight measurements and the standard deviation about this mean both in units of experimental error for each ring of fuel pins and the bundle average. The calculations are normalized to the measured bundle average ratio of U235 to U238 atoms. The calculations agree very well with experiment with the exception of the Pu240/Pu239 ratio for the two low burnup results. Table A-7 contains the percentage difference between calculated and experimental bundle - 22 - average plutonium to uranium ratios and again good agreement is demonstrated. Reactivity Coefficients The most important reactivity coefficients in CANDUs are associated with temperatures (fuel, coolant and moderator) and coolant density. The fuel temperature coefficient is strongly negative for fresh fuel due mainly to Doppler broadening of the U238 resonances as the fuel temperature increases. The broadening increases the effective neutron capture resonance integral. A typical fresh fuel temperature coefficient would be n>k/°C. As fuel burnup proceeds, and plutonium builds up, this coefficient becomes somewhat less negative. Pu239 has a capture and fission resonance at about 0.3 eV. An increase in the fuel temperature causes the thermal spectrum to harden, i.e. become less thermal, which along with the broadening of the predominantly fission resonance causes an increase in the number of fissions and hence the reactivity. A typical value of the fuel temperature coefficient for equilibrium fuel is -0.0043 mk/°C while at discharge the value is -0.0005 mk/°C. Increase? in coolant temperature result in lower coolant density and a higher effective neutron temperature. The effects are complicated but some of the parameters influenced are i) The U238 resonance integral is decreased, which tends to decrease resonance absorption ii) There is less scatter during slowing down, which tends to increase the resonance absorption iil) The increased neutron temperature increases the number of fissions in Pu239. - 23 - The overall effect is a coefficient that is strongly burnup dependent varying from strongly negative for fresh fuel to almost zero for equilibrium fuel. Figure A.16 shows the reactivity change in Pickering units as a function of heat transport system temperature for fresh and equilibrium fuel. The moderator temperature coefficient is a balance among several effects. An increase in moderator temperature decreases the moderator density, resulting in small changes in moderator absorption and moderating effect, and also raises the neutron temperature. The net result is a coefficient with some burnup dependence varying in the range -0.098 mk/°C to +0.048 mk/°C for fresh to equilibrium fuel. This coefficient is not very significant from the point of view of control or safety since changes in moderator temperature occur slowly. The overall effect of all these coefficients is, in Pickering, a negative power coefficient at all fuel burnups. The coolant density reactivity coefficient, the void effect, is concerned with the hypothetical instantaneous loss of all coolant in a unit. The coefficient is a complex interaction of at least five effects : 1) The U238 resonance integral decreases, which causes an increase in reactivity. 2) There is less slowing down and consequently neutrons have a better chance of being captured in a U238 resonance, which causes a reactivity decrease. 3) The neutron temperature increases, which if Pu239 is present causes a reactivity increase. 4) The neutron absorption in the coolant decreases,causing an increase in reactivity. 5) Ttv? leakage from the core changes, probably causing a decrease in reactivi y. - 24 - The net effect in Pickering is an increase in reactivity of about 1% when the coolant is voided. Kinetics Parameters In a small reactor the time dependence of the neutron flux, following a small change in reactivity from critical, can be well approximated by assuming that the spatial flux shape is independent of time. The time dependence of the flux level can then be calculated using standard formulae. The data on which this time behaviour depends are the change in reactivity, the mean neutron lifetime in the core, and the yields and half-lives of delayed and photoneutron precursors. In heavy water reactors, in addition to the six delayed neutron groups (corresponding to six precursors), nine photoneutron groups have been identified. While these photoneutron groups have relatively low yields, seven of them have precursors with half-lives longer than the longest half-life of the delayed neutron precursors. Experiments in ZED-2 and ZEFP have shown that good agreement can be obtained with experimental transient behaviour using all fifteen delayed and photoneutron groups. Good agreement for times less than about 100 seconds can be obtained by collapsing the photoneutron groups into the delayed neutron group. Doing this gives the following effective yields and decay constants for the six delayed neutron groups: oup Yield Decay Constant (s ) 1 0.0003808 0.0007363 2 0.001524 0.03174 3 0.001392 0.1177 4 0.003278 0.3139 5 0.001077 1.402 6 0.000252 3.919 0.0079038 - 25 - Tn large reactors the assumption of constant spatial flux shape after a small reactivity change from critical is no longer valid, due to the loose coupling between widely separated regions of the core. The problem faced was how to get relevant data from a small reactor. The solution adopted was to do experiments in ZED-2 with split cores. One hundred and twelve natural uranium metal rods on a 15 cm square pitch formed the core and a six lattice pitch split core configuration was used, Figure A.17. The core segment could be further decoupled by the insertion of a row of H2O filled aluminum tubes. Flux transients were initiated by the freefall insertion of a full length cadmium bearing aluminum rod into a guide tube located interstitially in the lattice. The location was chosen to induce a maximum first azimuthal flux tilt. A second guide tube was located in the other core segment at a similar position to maintain pre-drop flux symmetry. Detailed static flux measurements were carried out before commencing the dynamic experiments, Figures A.18, A.19. The experiments were simulated with a multi-dimensional kinetics code which employs the improved quasistatic method. In this method the space time dependent flux is factored into an amplitude function which is only time dependent and a space function which is only weakly dependent upon time. The time dependent multi-group diffusion equation is [-MfFp]«|)(r,E,t) + Sa[())(r,E,f)] = ± ^ cf)(r,E,T) where M is the removal and scattering operator, Fp the prompt fission source operator and S a the delayed neutron source. The total flux is then factorized into The factorization is defined by the arbitrarily imposed normalization Here ^* is the adjoint flux. With this condition the amplitude equation for 6(t) reduces to the point kinetics equation m. a^p wt> k where the quantities P(t), $(t), etc. must be derived by suitable averaging with the time dependent shape function ip(r,E,t). Upon substitution of the factorized total flux into the diffusion equation the shape equation takes the form [-MfFp]*(r,E,t) v l t 6(t) " 9t •"'»"»-" The derivative r— l^(r,E,t) is replaced by the first order backward at difference operator 9 g-r- i|j(r,E,t) = [i|i(r,E,t) - tf)(r,E,t-At)]/At which is valid when ip(r,E,t) changes slowly compared to 9(t). The advantages of the method appear in the great stability of the numerical solution at large time steps and in the small number of shape calculations necessary to achieve acceptable accuracy in cases - 27 - where flux shape variations are much slower than amplitude variations. Typical results are shown in Figure A.20. The curves show the ratio of the flux detector signals in each core segment as a function of time for two cases, with and without the H2O absorber curtain. Generally good agreement is obtained, particularly in the early part of the transient. Discrepancies are attributed primarily to the treatment of the delayed photo- neutrons. In Bruce, shutdown transients following the drop of 30 shutoff rods and 28 shutoff rods were obtained by recording the output of flux detectors (ion chambers) located at various positions. A typical set of results is shown in Figure A.21. The dynamic experiments were simulated with a two-group three- dimensional finite-difference code which uses the Improved Quasistatic (IQS) method. Results of this simulation are also shown in Figure A.21. Since out-of-core measurements showed some scatter in the rate of rod drop, two transients were calculated for each case. The FAST transient was based on the fastest drop measured in the laboratory tests and the SLOW transient was calculated with a slower rod drop characteristic, to study the sensitivity of neutron transients to possible scatter in the rod drop rates. The measurements fall close to the FAST calculated transient, in accordance with the fact that most rods were found to fall at the corresponding speed. Figure A.22 shows the dynamic reactivity calculated for the 30 rod drop case. The static reactivity of the rods is also given. The dynamic reactivity at full insertion is more negative than the static. This is because of the retardation of flux shape brought about by delayed neutron holdup, together with neutronic decoupling. This retardation enhances the importance of the absorber rods and their reactivity. - 28 - Definition of Fuel Management The term "fuel management" is used in reactor physics to indicate those aspects of fuel loading, whether related to physics, engineering or economic decisions, which are associated with fuel utilization and fuelling system performance. Fuel management fits within the framework of the overall fuel cycle, and the design limits imposed on the reactor. A more precise designation is "in-core fuel management" since other aspects of fuel management such as fuel procurement, fabrication and disposal are peripheral to the work of the reactor physicist. It is in this sense that the term "fuel management" will be used here. Fuel management plays a central role in the design and subsequent operation of a CANDU reactor. As an integral part of the design process, it defines the requirements for several reactor parameters, for example, fuel and channel design, fuel handling system, control margins, etc. During reactor operation, fuel management defines the fuelling strategy and the fuel replacement pattern. Information generated by fuel management is also used for station fuel accounting. Objectives of Fuel Management The primary objective of fuel management is to determine fuel loading and fuel replacement strategies which will result in minimum total unit energy cost while operating the reactor in a safe and reliable fashion. Within this context, the specific objectives of CANDU fuel management are as follows: - 29 - a) The reactor must be kept critical and at full power. On-power fuelling is the means of maintaining criticality. If the fuelling rate is inadequate, the reactor would eventually have to be derated, resulting in large cost penalties. b) The core power distribution must be controlled to satisfy safety and operational limits on fuel and channel powers. c) The fuel burnup is to be maximized within the operational constraints to minimize fuel cost. d) Fuel defects are to be avoided or minimized. This minimizes replacement fuel costs and radiological occupational hazards. e) Fuel handling capability must be optimized. This minimises capital, operating and maintenance costs. Periods During Operating Life From the point of view of fuel management the operating life of a CANDU reactor can be separated into three periods. The first period is from first criticality until onset of fuelling. It is of limited duration, usually about 100 to 150 full power days.* The reactor is Initially loaded with fresh fuel. Consequently, there is considerable excess reactivity which is compensated by adding boron poison to the moderator. This period does not have a major impact on the determination of the total unit energy cost. Fuel management calculations are required to assess the effect of the initial fuel loading on the subsequent power operation. * A "full power day" is the energy generated by a reactor when operating at rated power for 24 hours. - 30 - When the excess reactivity In the core falls to a small value, fuelling begins to maintain the reactor critical. During this transitional or "pre-equllibrium" period, the reactor gradually approaches the final or "equilibrium" state. The equilibrium condition In a CAKDU reactor is reached after approximately 400 to 500 full power days. It is characterized by a relatively unchanging core configuration in which the macroscopic or global power and burnup distributions do not vary significantly with time. The burnup of the discharged fuel and the fuelling rate of new fuel become essentially constant. Although the macroscopic power distribution Is relatively unchang- ing, the microscopic or local power distribution Is constantly changing as fuel burns up and is replaced. The equilibrium period covers about 95 percent of the reactor life. It is by far the most important period in determining the total unit energy cost. Most fuel management studies are therefore performed for this core configuration. On-Power Bidirectional Fuelling A CANDU reactor is fuelled while the reactor is operating at full power. After the excess reactivity of the initial fuel load has decreased to a small value, several fuelling operations are carried out every day on various channels In the core. Only a very small fraction of the core is changed at one fuelling operation, so that fuelling is essentially continuous. This avoids the need for a fuelling shutdown, making it unnecessary to tailor the fuelling schedule to the utility power grid's system requirements. - 31 - Fuelling is the primary method of controlling the long-term core reactivity, and maintaining the desired power distribution. By adjusting the overall fuelling rate we can control the core reactivity. By fuelling one region of the core more frequently than another we can control the power distribution - a lower fuelling rate leads to an Increased irradiation and reduces the region's relative power. An important consequence of this feature is that the reactor regulating system has only to compensate for changes in reactivity due to power maneuvering, fuelling of one channel at a time, and short-term xenon transients. Therefore, the reactivity control margin can be maintained at a small value resulting in good neutron economy and contributing to low fuelling costs. Another important feature of CANDU fuelling is the bidirectional axial fuelling. Alternate fuel channels are fuelled in opposite directions by inserting new fuel in one end and removing irradiated fuel at the other end. This results in a symmetric axial flux and power distribution and an approximately constant burnup for all discharged fuel bundles in a radial burnup region. Normally only a portion of the fuel in a channel is replaced in a single fuelling operation. The remaining fuel is shifted along the channel. During its life in the reactor each fuel bundle moves along the channel in a series of steps. Selection of Axial Fuel Management Scheme One of the primary requirements of fuel management studies for CANDU reactors is the determination of the optimum axial fuel management scheme. For the once-through fuelling normally used in CANDU-PHW, this requires determining the number of bundles which are replaced with fresh ones or "shifted" every time a channel is fuelled. This is referred to as the "bundle shift scheme". - 32 - The selection of the bundle shift scheme is usually done by comparing a number of possible schemes and determining the one which minimizes total unit energy cost for the reactor while satisfying the appropriate limits. This comparison is done during the design phase to select the best scheme for reference design and initial fuelling. It is also, at times, repeated during the operating life of a plant to estimate whether different conditions, such as availability of improved fuel design or better knowledge of certain parameters acquired during commissioning and operation warrant a change. Since the CANDU reactors are fuelled on power, a change in the bundle shifting scheme can easily be accomplished with a change in the operating procedure, requiring little lead time, and without shutting down the reactor. Basis for Comparing Axial Fuelling Schemes In order to arrive at the axial scheme with the minimum total unit energy cost, a series of parameters have to be calculated and compared for each scheme. The parameters which are normally required are: a) discharge burnup b) maximum channel power and channel power peaking factor c) fuelling machine usage d) fuel performance and bundle power Many of these parameters are interrelated and therefore they cannot be considered separately. Nevertheless, some remarks about each will be made to give an appreciation of their importance. - 33 - Fuel Burnup UsualLy, bundle shifting schemes involving a small number of bundles, 2 or 4, give the highest burnup since they approximate more closely the ideal "continuous" fuelling. However, the loss of burnup associated with an 8 bundle shift is small while 10 or 12 bundle shifts produce a significant burnup penalty. Maximum Channel Power and CPPF Usually the time averaged maximum channel power is not significantly affected by the axial fuel management scheme. The variations in channel power with respect to the "nominal" distribution, however, are strongly dependent ur ^n the axial fuel management scheme. The Reactor Overpower Protection (ROP) detectors are calibrated, on a frequent basis, to reflect the maximum ratio of instantaneous to nominal channel power or CPPF. In order to maximize the margin to trip it is important that the CPPF be kept at a minimum value. A fuel shifting scheme involving a small number of bundles shifted may be preferable since it causes small variation in the channel power distribution and therefore a small CPPF. For instance, the (CPPF-1) obtained from ai eight bundle shift scheme is about twice that obtained from a four bundle shift scheme. Current CANDU reactors are designed so that the ratio of the limiting channel power to the nominal channel power is higher for the channels in the outer part of the core. Thus, higher CPPF can be tolerated in the outer region than in the inner. As a consequence, mixed fuel shifting schemes involving, for instance, A bundle shift in the inner region and 8 bundle shift in the outer core region are often considered and used. - 34 - Fuelling Machine Usage In order to minimize the load, and the maintenance costs on the fuelling machine, the number of fuelling operations should be kept to a minimum. From this point of view, the more bundles thai are loaded per cycle, the fewer visits the fuelling machine has to make to the core. Fuel shifting schemes involving large numbers of bundles are therefore desirable. There is an upper limit on the capability of the fuel handling system so that it may not be possible to fuel the entire core with a fuelling scheme involving a small number of bundles per shift, for example, a four bundle shift. As discussed previously, it is possible to fuel outer region channels with eight bundle shifts and inner region channels with four bundle shift. The fuel handling system capacity then determines the size of the region which can be fuelled with four bundle shifts. Fuel Performance Another factor considered in selecting the axial fuelling scheme is the fuel performance, specifically, the probability of fuel defects. If irradiated fuel bundles are moved along the channel from a low power position to a high power position, the sudden increase in power (power ramp) may cause the bundle to fail. Fuel Scheduling During Approach to Equilibrium and Equilibrium Introduction During approach to equilibrium and equilibrium operation of a reactor the fuel scheduling or fuel replacement pattern has to be determined. For the CANDU reactors, fuel scheduling requires the determination of: i) the bundle shift scheme, and ii) the sequence of channels to be fuelled. - 35 - The determination of the bundle shift scheme is mainly a design problem. Normally, it is specified before the reactor is commissioned, although operating experience may lead to changing the originally specified scheme. The determination of the sequence of channels to be fuelled is an operating more than a design problem. While designing a reactor, it would be possible to specify the detailed schedule of fuel movements, giving the time of each move in the lifetime of the reactor in terms of integrated reactor power. Such a detailed fuel schedule Is of limited value for the following reasons: a) It might be unusable because no allowance can be made in this type of approach for unusual, random or unforeseen occurrences such as temporary unavailability of the fuelling machines, fuel failures, or action of the reactor regulating system. b) It would likely be inefficient, since it would not take into account inaccuracies in the calculations, and experience gained during commissioning and operation of the reactor. Detailed simulation of reactor operation is performed to provide the operator with some broad guidance in the form of general fuelling rules. These rules should be framed so as to allow the operator to make a decision as to the action to take under a large range of conditions using the information which he has available on the state of the reactor. In this section, the general rules used to select the channels to be fuelled during reactor operation will be examined. How these rules are used in computer programs for reactor simulation will be examined and some of the characteristics of fuel management during approach to equilibrium and equilibrium will be reviewed. Some data obtained from the operation of the Bruce NGS A reactors will be used for this purpose. - 36 - Rules for Channel Selection The selection of the channels to be fuelled is usually made on the basis of the following general guidelines: 1. Priority in fuelling is given to channels with the highest burnup. This is an obvious requirement. To minimize the total unit energy cost it is desirable to maximize the burnup obtainable from the fuel. Therefore, for any reactor configur- ation, the high burnup channels are the ones to fuel, where possible. 2. The power distribution must be controlled to approximate the reference power distribution in order to limit overpowers and minimize CPPF. Underpowered zones are fuelled preferentially and overpowered zones are avoided. 3. The power distribution has to be kept symmetrical. Distortions in the power distribution would increase the load on the reactor regulating system and increase the probability of reactor trips by reducing the margin available at the ROP detectors. In order to maintain a symmetrical power distribution, two factors have to be considered: a) Axial distortions can be minimized by fuelling an equal number of channels having opposite fuelling direction at about the same time. As we have seen, in CANDU reactors, adjacent channels have opposite fuelling and coolant flow direction. In order to avoid end-to-end tilts it is desirable that an approximately equal amount of fresh fuel is present at each end of the core at all times. - 37 - b) Radial and azimuthal distortion can be avoided by fuelling an equal number of channels in each zone controller region. CANDU reactors are divided for control purposes into seven radial and azimuthal zones, each containing two axial light water controllers and a set of measuring devices. Fuelling an equal number of channels in each controller zone helps to keep the level of light water in each controller compartment close to the average for the reactor and to minimize radial and azimuthal distortions in the power distribution. If distortions are already present, low zones are preferentially fuelled to raise the power. 4. Maximum separation Is maintained between channels fuelled at about the same time. Concentrations of freshly fuelled channels would create "hot spots" In the power distribution increasing the probability of fuel defects, and increase the CPPF, thereby reducing the margin to ROP trip. 5. High reactivity gain in the channels fuelled is desirable. Sufficient reactivity is normally maintained by adjusting the fuelling rate to compensate for the reactivity loss due to fuel burnup. When thia fuelling rate cannot be maintained, it is necessary to select channels which will produce high reactivity gain upon fuelling. These are the high burnup channels with high neutron importance, usually channels in the innermost part of the core. Obviously these guidelines do not provide a unique selection and are often conflicting. The process of channel selection for fuelling is then performed by compromise and trying to achieve a satisfactory overall balance. The procedure used to select the channels to be fuelled in an operating station will be examined later on. - 38 - Simulation Methods In the computer simulation the selection of the channels to be fuelled can be made "manually" or "automatically". In the "manual" selection mode, the user specifies in input to the program the channels to be fuelled. The manual selection is usually performed in the following steps: a) The user decides how many channels should be fuelled in the next irradiation srop, typically 5 or 10 full power days, in order to maintain the desired reactivity margin. This could be done on the basis of previous experience or with the aid of the reactivity gain per channel fuelled calculated using a simple one group perturbation theory method. b) The codes produce, from the last calculation, a list of channels in decreasing order of burnup, the channel bundle power and irradiation distribution. On the basis of these data, the channels which best fit the above-mentioned guidelines are marked for fuelling in the next simulation step. The process may require repetition of some step's if the results are not satisfactory and the conclusions depend to some extent on the "experience" of the user. The manual method is very flexible since different rules can be used easily. It is, however, very time consuming and its use is reserved for detailed studies when the parameters of the reactor are well established. - 39 - The AUTOFUEL Method To eliminate this repetitive "manual" selection of channels to be fuelled, an "automatic" procedure has been developed. In the "automatic" mode, the guidelines discussed above are translated into a series of logic steps that are programmed into the computer, and result in a selection of a set of channels at every time step. In performing a simulation with this method, the channel selection is done in the following steps: a) The reactor core is divided into burnup regions. There are typically two burnup regions in current CANDU designs. For each burnup region a minimum acceptable discharge burnup is specified. b) The burnup regions are further subdivided into blocks of channels. Each channel block is assigned a target or reference power, usually obtained from a time-averaged calculation. c) The allowable maximum bundle and channel powers, and fuelling rate ar» specified. Also a linear reactivity rundown rate is given in input, together with the excess reactivity from the previous flux calculation. All other data needed, such as reactor geometry, bundle and channel power distribution, are available from other parts of the calculation. For each time step the sequence of operations is: a) For each channel, the expected reactivity gain from fuelling is evaluated using a one-group perturbation theory method. The required reactivity Insertion to maintain the desired margin over the next period of simulation is also evaluated. - 40 - b) Each pair of symmetric channel blocks is tested for excessive difference in power or tilts. If the tilt is in excess of a specified limit, the block with lower power is marked for fuelling. c) The power of each symmetric pair of blocks is calculated and a list of the power for each pair of blocks is produced in order of decreasing percent variation from the reference block power. d) For each block selected in b) and, starting from the top of list produced in c), for each pair of symmetric blocks a channel is chosen for testing on the basis of the highest burnup in the block. Before the channel is accepted for fuelling, it is tested for: i) fuelling direction - if too many channels have already been fuelled in that direction the channel is rejected. This is done to maintain axial symmetry. ii) surrounding channels are checked for high bundle and channel power, or for recent refuelling. This is dona avoid creating "hot spots" in the power distribution. If the channel does not satisfy the above two criteria, it is rejected and the channel with the second highest burnup in the block is tested. e) The process is repeated until enough channels have been found to give the desired reactivity insertion over the next period of simulation. f) At the end of the selection process, the channels are fuelled and the calculation continues. - 41 - This method usually gives maximum channel and bundle powers which are slightly higher than obtained with the "manual" selection. Obviously, In using "manual" selection the same rules can be used with more flexibility. However, comparisons of results obtained from AUTOFUEL with actual data for the Pickering reactors give a rather satisfactory agreement. Initial Fuel Loading and Approach to Equilibrium Having discussed the criteria used in deciding the fuel scheduling for the equilibrium core, some of the characteristics of the initial core and the transient leading to equilibrium operation will be exauined. The Fresh Core Most CANDU designs have two burnup regions when at equilibrium. This differential burnup is used in combination with absorber rods to flatten the power distribution. At the time of first start-up, a CANDU core consists entirely of fresh fuel. In order to be able to operate the reactor at full power without exceeding the target values on maximum channel and bundle power, additional flux flattening is often required. Bundles having a concentration of U235 lower than natural uranium are loaded in some positions of the inner core channels. These "depleted" bundles have lower reactivity than natural UO2 bundles, and hence tend to reduce the flux and power in the inner region. They are removed from the core during the course of normal fuelling and replaced with natural UO2 bundles. By that time, there is sufficient differential burnup to flatten the flux distribution, and special bundles are no longer required. - 42 - The 850 MWe CANDU reactors are loaded at start-up with two "depleted" bundles in each of the inner burnup region channels. In the case of the Darlington reactors, the U235 concentration of these depleted bundles will be 0.57 percent by weight. Initial Transient to Onset of Fuelling The initial reactivity transient is Illustrated in Figure A.,23. At start-up, the core has a considerable excess reactivity which is compensated by adding soluble boron poison to the moderator. As the fuel burns up, plutonium is produced so that the core excess reactivity initially increases up to about 50 full power days. From this time the concentration of fissionable isotopes In the fuel decreases, as does the reactivity. By approximately 110 full power days, the core excess reactivity falls to zero, and some fuelling must be done to keep the reactor operating. The maximum channel power versus time is presented in Figure A.24. The trend is similar to that of the reactivity. A peak Is reached at approximately 50 to 60 full power days. Then the maximum channel power decreases until onset of fuelling. By this time the channel power distribution is overflattened and some fuelling must be done to remove the depleted bundles from the core inner region and bring the channel power distribution to equilibrium shape. Initial Fuelling Up to Equilibrium Fuelling usually begins shortly before the boron concentration In the moderator falls to zero. The fuel scheduling is done using procedures similar to the ones followed for the equilibrium core. The bundle shifting scheme which will be used for equilibrium fuelling is also used in this period while the channels to be fuelled are selected on the basis of the general guidelines discussed previously. These guidelines, however, must be applied with some caution since fuel management during approach to equilibrium presents some special problems. - 43 - At the onset of fuelling the inner core region has the highest turnup and the lowest power relative to the equilibrium power distribution. If we wish to maximize the discharge burnup and bring the power of the inner channels rapidly up to the equilibrium value, we should fuel preferentially in this region. Only some channels, however, can be fuelled in the inner region before the power rises to the equilibrium value. After that, some outer region channels must be fuelled in order to keep the reactor critical. In the outer region, channel burnup decreases with increasing radius so that fuelling tends to proceed generally from the inside toward the outside. To avoid "hot spots" we have to maintain separation between channels which are fuelled at approximately the same time. Therefore, only a few channels are fuelled in each ring in the first cycle. Channels missed will be fuelled on subsequent cycles until the burnup in each ring reaches an approximately uniform distribution between fresh and discharge. In practice, for every period a mixture of inner and outer channels are fuelled with the aim of maintaining a symmetric power distribution and a fuelling rate high enough to compensate for the reactivity loss due to fuel burnup. This procedure implies that the channel with the highest burnup is not always the one which is fuelled. In going from onset of fuelling to equilibrium, the average burnup of the discharged fuel bundles increases approximately linearly until the constant equilibrium burnup is achieved. Usually, during this period the fuelling rate is significantly higher than at equilibrium and may present considerable variations around the average, if care is not taken to maintain a balance of inner and outer channel fuellings in a given time period. Fuelling outer channels results in very small reactivity gains since these channels have low neutronic importance and also low burnup. - 44 - In the approach to equilibrium period, many of the discharged fuel bundles have low burnup since they come from outer channels or end positions in the chaunel. These bundles can be recycled into other channels to acquire additional burnup. This would result in a better utilization of the fuel. However, it would complicate considerably the fuel scheduling and increase the fuelling machine usage at a time when the fuelling rate is already abnormally high. Recycling of fuel bundles has not been used in the nuclear generating stations operated by Ontario Hydro because the operational problems outweigh the advantages. Different economic conditions, however, can make the recycling attractive. Usually a detailed simulation of the approach to equilibrium phase is performed just prior to commissioning the reactor. This is useful because of *"he special problems associated with fuel scheduling during this period and also to provide the operator with some general guidance on the reactor response before actual experience is available. Approach to Equilibrium for the Bruce A Reactors The Bruce A Nuclear Generating Station consists of four identical 740 MWe reactors which were placed in service between 1976 and 1979. Figure A.25 illustrates the burnup of the discharge bundles as a function of integrated reactor power. The data obtained from the "manual" simulation are plotted with those obtained during operation as calculated in the routine "monthly" run for the first two reactor units placed in service. - 45 - The -ipreement is quite good. Note that the channel fuelling sequence is different in the three cases. The burnup increases in linear fashion until an approximately constant value is reached. At this point the core is considered at equilibrium. Review of Operating Experier.ee Introduction The excellent performance of the Pickering reactors is a measure of the successful operation of CANDU reactors. These reactors have achieved a high capacity factor and very low fuelling cost. Approximately 30 reactor-years of commercial experience have now been accumulated with the four Pickering reactors. The average capacity factor of all units since their in-service date is very close to 80 percent. In 1978 the total unit energy cost at Pickering was 10.1 m$/(kW.h). This compares favourably with the fuelling cost alone of 13.7 m$/(kW.h) for the most efficient coal-fired station in the Ontario Hydro system. Meeting the objectives of efficient fuel management during both design and operation has made a significant contribution to this achievement. To illustrate how this performance is achieved with the methods and procedures for fuel management discussed in the previous sections, the operating experience at the Pickering and Bruce Nuclear Generating Stations will be reviewed. In doing this one typical unit from each station will be considered to highlight the lessons learned and problems of general Interest encountered. - 46 - Fuel Management at Pickering Initial Operation Pickering Unit 1 went critical on February 25, 1971. The Pickering core was designed to have a radially uniform discharge burnup. Adjuster rods provide radial and axi^l flux flattening. The first fuel charge consisted entirely of natural UO2 bundles. Because the adjusters provided sufficient flattening, no "depleted" fuel bundles were required. Fuelling began at approximately 5 TW\h (^130 FPD) when the excess reactivity in the core was reduced to approximately 5 x 10-3 6k/k. A uniform 8 bundle shift scheme was selected as being the most suitable on the basis of the following economic factors: a) fuel make-up cost, that is the cost of fuel bundles inserted in the core per unit energy output; b) fualling machine operating and maintenance cost; and c) costs associated with fuel defects. As that time the only known cause of fuel failure associated with fuelling was large increases in fuel rating of highly irradiated bundles after fuel shifting. No special consideration was given to minimizing the CPPF. The Pickering reactors use out-of-core instrumentation rather than in-core detectors for overpower protec- tion. Out-of-core instrumentation is not very sensitive to local flux perturbations. In addition, the Pickering reactors have pre- specified limits on the coolant temperature rise across each channel. If these limits are approached, because of an overpower, the reactor power is reduced. By late 1971, a large increase in the iodine-131 concentration in the heat transport system indicated the presence of fuel failures. The resulting investigation revealed two reasons for the fuel defects: - 47 - a) excessive variations in bundle power due to adjuster rod manoeuvering; and b) high incremental bundle powers due to the eight bundle shifting scheme. The first problem was eliminated by re-analyzing the adjuster rod sequencing and associated reactor power levels, with an imposed arbitrary limit of 15% on bundle power variations. The new adjuster rod withdrawal sequence considerably reduced bundle power variations. The second cause of fuel defects had two components: a) large permanent increase in fuel rating when bundles in position 1 are moved to position 9 by an 8 bundle shift, and b) a short exposure (about 15 minutes) of bundles in position 1 and 2 to high powers at the centre of the channel during the fuel movement. To remedy the problem two steps were taken. Fuel management simula- tions were carried oat to compare the economics of 8, 10 and 12 bundle shifting in light of the increased cost associated with fuel defects. The 10 bundle shift scheme was found to give a small burnup penalty when compared to the 8 bundle shifting scheme, while eliminating permanent increases in bundle power due to fuel re- arrangement within the channel. This scheme was, therefore, adopted for all the high power channels in the core. - 48 - The exposure of some bundles to high fluxes in the centre of the channel during fuel movement was shortened to about 5 minutes by a change in the sequence of operation of the fuelling machine. These steps were very effective. Of the approximately 48,000 bundles fuelled to the end of 1974, 101 had developed defects. Of these, 88 are attributed to the above effects. Excluding these the defect rate is 0.027 percent. Some of the lessons learned in the area of fuel scheduling from this period of operation of the Pickering reactors, can be summarized as follows: a) It is valuable to be able to simulate reactor operation accurately and in a timely manner. The availability of individual bundle power histories from the simulations enabled a prompt identification of the defective bundle location and provided the data to understand the underlying causes of fuel defects. b) It is valuable to be able to compare different bundle shifting schemes on short notice, taking into account changing operating requirements or situations Unforeseen during the design phase. c) The flexibility of the CANDU fuel scheduling allowed the station to adapt to new schemes and incorporate changing operating requirements without having to shut down the reactors. d) It was important to develop a fuel design more tolerant of the variations in power which can be expected from movement of the reactivity mechanisms or from the fuel shifting scheme itself. - 49 - Subsequent Experience In 1972, tests indicated that the deposition of a thin graphite layer on the inner surface of the fuel sheath would make the fuel more tolerant to power variations. This fuel, designated 'CANLUB1, became the standard design and bundles of this type were used to fuel the reactor starting in 1974. The experience gained with this fuel design allowed the relaxation of the 15% limit on short-term power variation. The fuel shifting scheme, however, was not changed. Since then the fuel management at Pickering has been very successful when measured against the objectives described in the introduction of the first section. Maximum Bundle Power History Figure A. 26 shows the maximum bundle power as a function of integrated reactor energy for Pickering Unit 1. The shaded band is a 440% variation about the nominal reference bundle power of 640 kW, the upper end of the band being the target power limit of 705 kW. The original objective of the fuel scheduling was to maintain the maximum bundle power close to the nominal 640 kW value. The data show that, in general, this was achieved. In particular, if we examine the period after 1974 O40 TW.h) with the unit having reached maturity, the variation in maximum bundle power is very small. A slow trend towards a lower target maximum bundle power is also evident from about 50 TW.h. In the last 4 years the maximum bundle power has been varying around av. average value of approximately 600 kW with variations of the order of 6% or less. - 50 - Maximum Channel Power History The history of the maximum channel power is given in Figure A.27. The reference design value of 5.5 MW has been maintained, with a few exceptions, to within +10%. As with the maximum bundle power, the band of variation around the nominal value decreased as the unit reached maturity. Burnup and Fuel Consumed Figure A.28 shows the core excess reactivity as a function of reactor integrated power. After the initial reactivity transient due to fresh fuel had decayed, the excess reactivity in the core was maintained close to zero. Moderator poison as a means of "storing" reactivity has been used very rarely at Pickering. Figure A.29 shows the average monthly discharge burnup as obtained by the simulation. Figure A.30 shows the fuel added versus reactor heat. Also plotted is the line of "ideal fuel added" assuming a burnup of 175 MW.h/kg. The data indicate that the burnup is in the range 170-175 MW.h/kg. Fuel Defect Performance The performance of the fuel at Pickering after the initial problems were solved has been extremely good. The total number of defective bundles, including suspected ones, was 112 up to the end of June 1978. This gives a defect rate for the 4 units of 0.12%. Fuel Management at Bruce The Bruce units were the first CANDU reactors to incorporate a regional overpower (ROP) system for overpower protection. Early fuel management studies indicated that controlling and minimizing - 51 - the CPPF was imperative to maintain an adequate operating margin at the ROP In-core detectors. The ROP detectors are calibrated on a frequent basis to reflect the CPPF existing in the core. Large variations in the CPPF would increase the frequency of time consuming calibrations and the probability of spurious activation of the shutdown systems. The fuelling schemes used at Pickering (8 or 10 bundle shift) would produce unacceptably high CPPF in Bruce A. On the other hand schemes involving a small number of bundles, 2 or 4, tend to increase the fuelling machine usage and cause relatively high increases in bundle power during shifting, leading to higher probability of fuel defects. From a number of studies performed during the final design phase, it was concluded that the most suitable scheme was a mixed 4 and 8 bundle shifting. The channels In the inner part of the core are fuelled with a four bundle shift, while the outer channels are fuelled with an eight bundle shift. The studies showed that this scheme has the following character- istics: a) It yields a discharge burnup comparable to that obtainable from a uniform 2 or 4 bundle shifting and slightly higher than that of an 8 or 10 bundle shifting scheme. b) The value of the CPPF minus one is approximately half that of an 8 or 10 bundle shift. c) The probability of fuel defects due to fuelling was acceptably low. CANLUB fuel, which is more tolerant to power variations, is the reference design. - 52 - A uniform 4 bundle shift would have been preferable from the point of view of minimizing CPPF. This scheme, however, would have Increased excessively the fuelling machine usage. Fuel management at Bruce has been generally successful as we shall see later by reviewing the operating data. Some operational difficulties, associated with fuelling, have been encountered with the reactor regulating system and the ROP system. Reactor Regulating System The reactor regulating system employs in-core, self-powered detectors to provide an estimate of the power in 14 zones (regions) of the core which are controlled by the 14 light water zone control- lers. The flux detectors are continuously calibrated to estimates of the thermal power in each zone. The calibration factors are derived from 3 or 4 fully instrumented channels (inlet and outlet flow, inlet temperature and temperature rise measurements) located in each zone. Ideally, the fuel scheduling provides the means of controlling the power distribution and, hence, of maintaining an approximately uniform level distribution in the controllers. At Bruce, however, Individual controllers have displayed a tendency to drift to extreme level over a period of ^tsilling. This anomalous drift has adversely affected, at times, the ability to maintain the desired fuel scheduling. Analysis has shown that the drift is directly related to local burnup dependent flux variations which are not completely calibrated out of the In-core flux detector signals. A fuelling operation causes a redistribution of the flux and power distribution in the core. The perturbation consists of both a global tilt component and localized flux peaking component. Since - 53 - the number of controllers is limited, only the global tilts can be effectively controlled. Localized flux peaks due to the fine structure in the Irradiation distribution cannot be controlled. If these are not calibrated out of the detector signals, an inappropriate control action might be taken which can cause the controllers to drift to extreme values. The problem is presently being handled by performing periodic manual adjustments of the pre-calibrated detector signals and spatial power distribution setpoints. Efforts Te currently underway to modify the spatial control algorithm to make it less sensitive to these local flux variations. Another difficulty encountered is related to over-response of individual zone controller level changes when one of the fully Instrumented channels is fuelled. Since an average of a few measured channel powers In a zone is used to provide an estimate of the average power in the zone, the local power peaking in a freshly fuelled channel may result in an apparent higher zone average power. The zone controller is called upon to reduce the estimated power to the desired setpoint and does so by filling to a higher level than is required. This type of calibration, which is also used at Pickering, is only a temporary solution to the problem. The Bruce reactors are currently operated at full electrical output which corresponds to approximately 88% of full thermal power. Once the additional power is required to provide steam for a heavy water plant on the site, some channels will be boiling at the outlet end. The power estimate from temperature measurements will, therefore, be unreliable. A more permanent solution is being sought through modification of the spatial control algorithm. - 54 - These problems have indicated the need for a representation of the control system response in fuel management programs, which are often used to predict the fuelling pattern during operation. The Regional Overpower System The occasional difficulties in maintaining a reasonable distribution of zone controller levels have also caused additional problems in maintaining an adequate margin at the ROP detectors. Since the desired fuelling pattern could not always be maintained, relatively high CPPF's were occasionally encountered. Moreover, as mentioned previously, particular care had to be taken in fuelling channels close to ROP detectors. The observed changes in detector reading following fuelling close to a detector are in the range 1.5 to 6 percent. Of this change, up to 2 percent has, in some cases, been attributed to control system action. In general, those changes in detector signals exceed the changes in channel powers, thereby resulting in an effective reduction of the operating margin. The net result of fuelling on the ROP system detectors and regulating system response has been to place an abnormal burden on operating personnel to ensure that fuelling does not lead to spurious overpower trips of the reactor. This has resulted in small temporary deratings and loss of production due to inadequate operating margin in the ROP system.. Some improvements are expected to be obtained through proposed modifications of the spatial control algorithm. The benefits will be derived through an improved ability to maintain a better distribution of the level in the zone controllers and an attendant improvement in the control of the CPPF. - 55 - Maximum Bundle Power History Maximum bundle power as a function of integrated energy for Bruce Unit 1 is presented in Figure A.31. It can be seen that after the onset of fuelling the maximum bundle power Increased fairly rapidly to a value of approximately 800 kW +10%. The variations in Bruce are somehow larger than Pickering due to the fact that the reactor has not yet reached maturity and also to the fact that no special attempt was made to strictly control the bundle power. The operational margin on maximum bundle power is larger in Bruce than it is in Pickering. Maximum Channel Power History Because of the relatively narrow margin on the ROP system and the need to minimize the CPPF, good control was kept on the maximum channel power. As can be seen from Figure A.32, the target value of 6.0 MW has been maintained, with a few exceptions, within a very small band. Burnup and Fuel Consumed Figure A.33 shows the fuel usage, burnup, and core excess reactivity as function of reactor integrated power. The upper portion of the figure shows the core excess reactivity. After the initial reactivity transient due to fresh fuel had decayed, the excess reactivity in the core was maintained close to 5 x 10~3 6k/k. This excess reactivity is held In soluble boron in the moderator to provide additional "shim" reactivity when fuelling machines are unavailable. The Bruce reactor does not have adjuster rods for shim purposes. Power shaping Is provided principally by burnup flattening. - 56 - The lower portion of the figure shows the average monthly discharge burnup as obtained by the simulation, and the fuel added versus reactor heat curve. Also plotted is the line of "ideal fuel added", for a burnup of 195 MW.h/kg. With boron in the moderator, the actual fuel added data indicate a burnup of about 175 MW.h/kg. - 57 - TABLE A-l CANDU POWER REACTORS POWER DATE OF MW TOTAL 18,208 MW(e) - 58 - TABLE A-2 CUMULATIVE LOAD FACTORS FOR REACTORS OVER 500 MW(e) TO END OF SEPTEMBER 1980 STATION CUMULATIVE LOAD TYPE FACTOR % Bruce-3 82.0 CANDU Stade-1 81.2 PWR PickerIng-2 80.9 CANDU Pickering-1 80.3 CANDU Point Beach-2 77.4 PWR Pickering-4 77.3 CANDU Pickering-3 75.4 CANDU Prairie Island-2 75.2 PWR Calvert Cliffs-2 74.7 PWR Connecticut Yankee 74.6 PWR Bruce-4 73.5 CANDU Bruce-1 73.0 CANDU - 59 - TABLE A-3 ALLOCATION OF EXCESS REACTIVITY Excess Reactivity Above Allocation Cold, Clean, Critical (X k/k) ADJUSTER RODS FULLY IN 1.5 MISCELLANEOUS STRUCTURAL 0.8 TEMPERATURE EFFECTS (a) 2.0 XENON AND SATURATING FISSION 3.6 PRODUCTS CONTROL MARGIN 0.2 BURNUP 2.2 10.3 (a) This is the allowance which must be made for going to operating temperatures and pressures in a fresh core. As the irradiation proceeds, the temperature effects decrease, and consequently more reactivity is available to increase burnup. - 60 - TABLE A-4 APPROXIMATE NEUTRON BALANCE IN EQUILIBRIUM CORE Leakage 0.0253 Non fuel absorptions 0.0724 - sheaths 0.0084 - coolant 0.0009 - pressure tube 0.0196 - calandria tube 0.0086 - moderator 0.0174 - structural material, control and adjuster rods 0.0175 0.0724 Fuel absorptions 0.9023 - U235 0.2445 - U238 0.3343 resonance 0.0930 non resonance 0.2413 - Pu239 0.2314 - Pu240 0.0207 - .Pu241 0.0119 - Pu242 0.0001 - Saturated fission product (Xe, Sm, etc.) 0.0320 - Other fission products 0.0274 0.9023 Total 1.0000 - 61 - TABLE A-5 CALIBRATION OF REACTIVITY DEVICES Reactivity Worth* (mk) Reactivity Device Measured Calculated % Deviation Mechanical Control Rods CA3 1.46 1.49 +2.0 CA4 1.57 1.54 -2.0 CA3 + CA4 2.53 2.59 +2.4 CA1 + CA2 + CA3 + CA4 4.93 4.96 +0.6 Shutoff Rods SA20 1.49 1.46 -2.0 28 Rods** 3.30 3.45 +4.5 30 Rods** 2.65 2.75 +3.8 Booster Rods BA5 0.91 0.90 -1.0 16 Booster Rods** 3.25 3.42 +4.0 k - 1 * mk = — x 1 000, where k = effective multiplication factor ** measured in units of mg B/kg D2o added to the moderator Table Mean Differences (Calculation—Experiment) and Standard Deviations About the Mean in Units of Experimental Errors 24oPu/339Pu J3fl 2 3=u/ u 2 39 Pu/338U 2 •! 0 Pu/333Pu excluding the 34 1 Pu/2 39Pu 342 Pu/2 39Pu first 2 bundles Centre 0.75±0 43 -3. 14±2 .12 0 75±2. 43 2. 12±0. 52 0.24±1 .02 1.02±2.20 Single Pin Middle- I 0. 17±0 19 -0. 72±1 .64 -0 43±3. 00 1.25x0 7 4 -0. 07±0 .97 22il.02 Six Pins Outer 17:10 .17 1.97±0 ..66 — J> .42±3, 40 -1. 59±1 4C -2 95±1 .72 94iO.55 Twelve pins -c 1 Bundle : 0 86±0 .73 -2 . 35±3 ?7 -0 55±l .13 -i 99^1 .•?7 -0 58±0.46 Average i - 63 - TABLE A-7 PERCENTAGE DIFFERENCE BETWEEN CALCULATED AND EXPERIMENTAL BUNDLE AVERAGE PLUTONIUM TO URANIUM RATIOS Nominal Burnup 980 1500 3250 5000 6500 7900 9100 10800 MW.d/t % difference +0.46 •M).43 -0.43 +0.24 +1.31 +0.86 +0.41 +0.24 800 DARLINGTON Length and location of rectangles denotes 7(10 - construction and commissioning period. Arrows denote flow of information. 600 POWERS 500 REACTORS MW(e) 250 200 I ON RESEARCH nEACTORS MW(th) General reseaich, devoit4irneiu and design information. 1945 1950 195S 1965 1970 197!> 1980 1985 1990 YEARS A.I GENEALOGY CF CANDU REACTORS 20 1973 1974 1976 1977 1978 1979 198P (Courtesy of Nuclear Engineering lni6rnational,i9U0 December) FIGURE A.2 COMPARISON OF THE PERFORMANCE OF FOUR TYPES OF NUCLEAR REACTORS FIELD WELD END GAS ANNULUS (Between Lines) a- i FIGUHK A-3 REACTOR CORE SCHEMATIC - 67 - 1/5 (/> IX LU O a. o 2 O z o a.ill S5 O - 68 - CA(_Ai\,Ofii« - SIDE TueesHEET •, 7. CALANDR.A VAUU AAL.. CA'.flNDRIA TUBES 18. \*OOER-TOH EXPANSION TO H EMBEDMENT RfNG ?9. CUBTAIIS. SHIELDING SLABS FUELLING MACHINE - SIDE TUSeSHEET 20, PRESSURE RELIEF PPES FN3 SMrgLO i-ATTiCE TUBES 21. RUPTURE OlSC tND SHIELD COOLING PIPES 22. REACTIVITV CONT^CL UNJT t^." INLET-OUTLET STRAINER 23. VIEWING POUT STEEL 8ALL SHIELOING 34. SHUTQFF UNIT END FITTINGS 25. AQJUSTEH UNI" FEEDER PI-'ES 26. CONTROL ABSORBPa UNIT MODERATOR INLET 28. VERTICAL FLUX DETECTOR LMN i "ORlZONTAL FLUX DETECTOR UNIT 29. LIQUID INJECTION SHUTDOWN ION CHAMBER 30 9ALL FILLING PIPE FIGURE A«5 REACTOR ASSEMBLY - 69 - 1 i 1 i i [ I 1 1 i 1 i 1 i ' ' ' i—l- f- --' |l [;.TT .A r IT > 1 [ I 1 (Ml i 2 3 4 i fi 7 a c in 11 2 11 14 IS IF 1 if 1 7 > 5 19 2C • VERTICAL FLUX DETECTOR (26) g SOLIi, CONTROL ABSORBER (4, (S) ADJUSTED (21) Q LIQUID ZONE CONTROLLER (6; (^ SHUTOFF ROD (28) HORIZONTAL FLUX DETECTOR (7) ZONE CONTROL ABSORBERS ZONE CONTROL DETECTORS VIEW OF REACTOR FACE FIGURE A-b REACTIVITY MECHANISM LAYOUT FIXED END OF CHANNEL o 1 1 CHANNEL CLOSURE 2 CLOSURE SEAL INSERT 3 FEEDER COUPLING 4 LINER TUBE 5 END FITTING BODY 6 END FITTING BEARING 7 TUBE SPACER 8 FUEL BUNDLE 9 PRESSURE TUBE 10 CALANDRIA TUBE 11 CALANDRIA SIDE TUBE SHEET 12 END SHIELD LATT1CF TUBr 13 SHIELD PLUG 14 END SHIELD SHIELDING BALLS 15 FIIF.LLING MACHIN- SIDE TUBE SHEET 16 CHANNEL ANMDLU3 BELLOWS 17 CHANNEL FV.-ITIOMING ASSEMBI Y FUEL CHAMNSL FND VIEW INSIUL PRESSURE TUBK ZIRCALOY SCARING PADS ZIRCALOY FUEL SHEATH ZIRCAIOY EMDCAP ZIRCALOY END SUPPORT PLATE URANIUM DIOXIDE PELLETS 6 CANL.UB GRAPHITE INTERLAYER 7 INTER ELEMENT SPACERS 8 PRESSURE TUBE FIGURE A-8 3/F.LEMENT FUEL BUNDLR EXPT O,O COOLANT: O Ain COOLANT: M ID K/ i .j I 3.0- 2 0 2U HEXAGONAL PITCH (cml "iGURE A-9 28-ELEMENT URAN.y.M OXICt L-4T7ICH, MA7E3IAL BUCKUNG VERSUS PITCH - 73 - i i I i i i r REFERENCE LATTICE i £ 9 U0 (He) CD 2 2 Q_ ROD - X 19 UO, ID C\l m 7 U02(D20) X i LU 7 RODS o 03 9 U02(C(8H22) I I I I I 2 Z FIGURE A. 10 2 The discrepancy between MICRETE values of B and those from flux mapping as a m function of Z = [Mass of 11/(Volume of D.O + 3.75 Volume C1QH.,jl / Mass of I lo 2.1 Test U/(Volume of DO + 3.75 Volume C1QHOO)]D f for Reference Lattice I. Points indicated by x are for one rod; those by 0 are for seven rods. - 74 - FIGURE A.1! BORON CONCENTRATION vs BF-3 DETECTOR RECIPROCAL COUNT RATE o 0.4- a. z o o o or a. 0.2« C.1- 80R0M CONCENTRATION img fi/kg 020) - 75 - FIGURE A.tt CALIBRATION OF AVERAGE ZONE LEVFL \ \ \ J \ a a \ z o 'S g» \ a] \ TTI 0 10 20 30 40 SO 60 70 80 90 100 AVERAGE ZONE LSVEL (%) - 76 - Q _ 1 -2 > h- o -3 LU -4 CC -5 -6 -7 -8 0 2000 4000 60C0 8000 BURNUP Mwd/V FIGURE A. 13= REACTIVITY VS BURNUP NATURAL URANIUM - 77 - O < LU ct: 2000 4000 8000 8000 BURNUP Mwd/V FIGURE A. I4-- APPROXIMATE ESTIMATE OF DISCHARGE BURNUP - 78 - BUNDLE AVERAGE NUCLIDE RATIOS COMPARISON OF CANADIAN CALCULATIONS AND EXPERIMENT 1.0 FIGURE A.15 239Pu/238U 0.9 240Pu/239Pu 0.8 241?u/239Pu 242Pu/239Pu 0.7 - 0.6 CD 0.5 0.4 I o o 0.3,— 0.2 3. 1 — 0 0.1 0.2 0.3 04 0.5 C.5 0.7 0. - 79 - n/m EX I T 101/ 150 200 I I _| ,, in r'"- 7 EXPERIMENT X UNIT 3 IN;T!A'_ STARTUP UMiT 4 INIT'AL jr,»R7ilo UNIT 2 STARTUP WITM t"f{fiy»'EO HjE'. :SULTS T F 100 150 2000 250 HEAT TRANSPORT SVSTEV TEMPERATUflE I°C) FIGUREAI6REACTSVITY CHANGE DUE TO HEAT TRANSPORT SYSTEM TEMPERATURE CHANGE IN THE PICKERING REACTORS TH"'. ?lv COMPARED WITH EXPERiMEMT - 80 - ZED-2 REACTOR - EXPERIMENTAL ARRANGFMEN1 O O O O \ OOOOOOOO OOOOOOOOOO ooooo ^ o o o o oooooooooooo oooooooooooo oooooooooooo oooooooooooo o o o o o o o o o o OOOOOOOOOO OOOOOOOO o o o o O \-,:URAL 'JRA.NiJM ,,E7fiL U'EfP) Fu.L RO'J? • h.O - F -LELV ALjf-;ii;jM :"UB AlR-fiLLED GUI LH IUB.I FIGURE A.17 Core Conf ig "o^ Troir" ;r-m Measurements- - 81 - ROC DROP 10 + EXPERIMENT CERKIW (m) FIG-A.18 ComDarisr.-. of Measured anc calculated - or r\ Lt-I \ A.19 Meor nd Ci' cu la ted R^iJ ["'•'.•r-iii i M < I 1 I 1 I I I i I I • : I i I i 2t. LClORb 2 -.ND 4 2.0 TILT x RUN 2 U- O 1 I I i i ; i I : i i .1 1.0 10.0 50.0 ELAPSED TIME (SECONDS) FIGURE A.10 . Comparison of Measured and Calculated Flux Tilts, With and Without yji Absorber Curtain. i 1 1 1 1 r ION CHAMBER IC5 COORDINATES (0.0,448.9,396.5) + VISICORDER DATA I 00 0 .6 .8 1.0 1.2 1.8 2.0 TIME (SECONDS) FIGURE A.2I- SDS I TEST - 30 RODS (UNIT - I) 0 -10 -20 STATIC REACTIVITY -30 00 LP O < I -40 -50 -60 .8 1.0 1.2 1.4 .6 .8 20 TIME (SECONDS) FIGURE A.22: DYNAMIC REACTIVITY CURVES - 30 RODS FIGURE A.23 EXCESS REACTIVITV VERSUS TIME - 87 - e.5« 3 LU o i u 3 5 i3 0- I 50 ,00 r a._ HGA2R DA"', F'GURE A.if MAXIMUM CHANNEL PCMT^ VERSUS TIME • • 20C- • • ..* ' 1QD MWh 'We |3U LVIVfll, ^j •• •• 1 5 K " • • *\ * CD UJ O X < o S oo ° ioo-4 oo *** • STUDY • UNIT 1 ^4 * 50- * UNIT 2 K 1 1 1 1 < i 1 1? 16 20 24 32 36 REACTOR HEAT (TWh) FIGURE 0-25 BRUCE A DISCHARGE BURIMUP 800 TARGET( I MIT MOMINAL MAXIMUM I 00 350 REACTOR HEAT(TWh) FIGURE A«A6 PICKERING UNIT 1, MAXIMUM BUNDLE POWER NOMINAL o i 3000 v,0 REACTOR HEAT, TWh FIGURE A.A7 PICKERING UNIT 1, MAXIMUM CHANNEL POWER Ul V, UJ o UJ 110 REACTOR HEAT. T\Vh FIGURE A.i? PICKERING UNIT 1, CORE EXCESS REACTIVITY !70 H WERAGE DIS'-HA -GF "UhUUP AVERAGb IVCORE BURNUP I 10 20 I 40 50 60 70 ao r 90 mo no REACTOR HEAT, TWh FIGURF. A.29 PICKERING UNIT 1, BURNUP - 93 - 400 5 IT 300- as a a < 200 130- •00 RGUflE A.5O PICKERING UNIT 1. FUEL USAGE I 10 20 REACTOR HEATTWh FIGURE A3I BRUCE UNIT 1. MAXIMUM BUNDLE POWER - 95 - - 96 - (MOV 78 OCT 78 SEPT 78 AUG 78 - ,^ EXCESS REACTIVITY / Q) AVERAGE OiSCHARGE 8URNU,'- ) AVERAGE IN-CORE BLJRNUP (-20 GUIDELINE (195 MWh/kg'J) I ^ lo 13 20 22 24 26 28 30 32 34 36 REACTOR HEAT, TWh FIGURE A.33 BRUCE UNIT 1 FUELLING PERFORMANCE - 97 - B. THEORETICAL LATTICE PHYSICS B.I INTRODUCTION Lattice physics addresses the problem of predicting the distribution of neutrons, as a function of space and energy, in a lattice cell. In a CANDU reactor a lattice cell, Figure B.I, consists of the fuel, coolant, pressure tube and calandria tube for one fuel channel along with its share of the moderator. Predicting the neutron spectrum in a lattice cell, therefore, involves all the physics of a reactor. AECL has developed a family of lattice codes which are called LATREP. At any time there exists a production version of the code and a development version. The physics of a CANDU lattice using the methods utilized in the present development version of LATREP which is referred to as SOLACE will be examined. B.2 THE PHYSICS OF SOLACE SOLACE divides -he neutron energy spectrum in the range 0 to 10 MeV Into 46 discrete energy groups. Of these groups 45 cover the range from 10 MeV to 0.625 eV and the remaining group deals with the thermal neutrons at averages below 0.625 eV. Besides calculating flux distributions and reaction rates the code will perform a fuel depletion calculation for each distinct ring of fuel pins in a fuel bundle. The code organization is shown In Figure B.2. Neutron Flux Distribution The generally complex geometry of the fuel bundle is converted to a set of concentric annul!. A jjt collision probability method is employed to calculate the neutron distribution throughout this set of annul!. - 98 - Multiple collision probabilities (W1^) are used to relate the neutron currents (J*) entering or leaving an annulus to the neutron source within the annulus. The resulting set of equations for each annulus can then be solved for the currents when an appropriate boundary condition is specified. The multiple collision probabilities are derived from first flight collision probabilities Pxv. Here x represents the mechanism by which a neutron appears in the annulus and y represents the mechanism by which it disappears from the annulus. There are three possibilities which are listed in Table B-l. The annular geometry restricts the number of first flight collision probabilities to eight. Some examples are: Pov : the probability that a neutron entering an annulus through the outer surface is absorbed within the annulus. pi° : the probability that a neutron entering an annulus through the inner surface leaves the annulus through the outer surface. The eight first flight collision probabilities can be derived, via reciprocity and conservation arguments, from the two probabilities Pvi and Pv°. Multiple collis' n probabilities are related to the first flight probabilities via the number of secondary neutrons produced per collision (T) where in general rt - 99 - v = number of neutrons per fission | = macroscopic fission cross section I = macroscopic ingroup scattering cross section s ' = lacroscoplc crc-s section for the n,2n reaction 2 = macroscopic total cross section 'or example io iv V iV VV VO W = pi° + p Tp ° + p Tp TP With the notation of Figure B.3 we can then relate the currents and sources for the nth annulus by J~ =0 Wvl+ J~Woi + J+ W11 Jn-1 QnWn Jnn + Jn-lWn which can be transformed Into J+ = A J~ ,+ B n n n-1 n J + ~n = Cn+ .,J1 n + Dn+ .1. - 100 - where C = (w"+ W01 C A ) n n n n+ A11 n D » WO±(C ...B + D .,) + WViQ n n n+1 n n+1 n n For the outermost region n=*N JN = CN+1 JN which defines the boundary condition. For example D N+r° gives a reflective boundary condition. Working from the boundary condition the An> Bn, Cn, Dn can be obtained and thence the currents J and J . The neutron collis- n n ion rate in the annulus in terms of the currents and source is w + iv — ov C = Q W + J ,W + J W n xn n n-1 n n n which must be equax to E Where £ is the macroscopic total cross section 0 'Is the neutron flux in the annulus. n ww vv +. .+ ..lv . .-. ov So The sources Q for energy group g are defined by g-1 g-1 n g fnTn ^ s n G=l G=l where Xg is the fraction of the mission neutrons having energies within group g. Non Thermal Neutron Cross Sections A distinction is made between the principal fertile material, U238 or thorium, and the other materials. Cross sections for non fertile materials are calculated from a library of microscopic cross sections. For resonance cross sections the infinite dilution cross section is used. Additionally the data for the 1 eV resonance in Pu2^0 is Doppler broadened. Cross sections in the resonance region for the principal fertile isotope are derived from effective resonance integrals. Assuming for now that we know the effactive resonance integral (R) then the neutron capture due Ko that resonance integral is A where with If the flux is $ T. • where £e ±$ the effective absorption cross section and consequently I = -H- 6 In SOLACE the unperturbed flux in a particular average group is calculated and used to obtain the number of absorptions in the fertile material. To calculate the effective cross section it is now necessary to obtain the perturbed flux and the assumption is made that the absorptions A are equivalent to a reduction in the source. The flux distribution is therefore recalculated with reduced sources to give the perturbed flux and the effective cross section of the fertile material is obtained for later use. There remains the problem of the calculation of the resonance integral. The data library contains information which relates the effective resonance integral to the ratio of the effective fuel surface to the fuel weight and to the fuel temperature. This information is generated by very fine group slowing down calculations for a single fuel pin. The effective resonance integral within each of the coarser SOLACE energy groups is then fitted to the equation R = A+BS/M and also to R = A 4~ + B' % = A' + B1 -| $ M M s - 103 - where The Doppler broadening is represented by R(T) = R(20) 1+.0KC+D | + E«/T> < T~ 293.16) where T is the fuel temperature - The fitted parameters A, B, B1, C, D and E are obtained for each SOLACE energy group and for Th232 and U238 in the form of metal, oxide, carbide and silicide and form part of the SOLACE data library. The problem of calculating the resonance integral has now been shifted to the problem of defining the effective surface. The effective surface is usually related to the physical surface (Sp) by a Dancoff factor (d) S = Sp(l-d) For arbitrarily placed fuel rods, Figure B.4, the Dancoff factor for rod A is Ki,(x) -r - 104 - whore the other fuel rods are assumed to be black to neutrons. Here T is the optical length measured from the surface cf cod A to its next interception by a fuel rod. The trapezoidal rule is used to integrate this equation anrf tiie double integral written in summation form is 20 20 = Ki (T )+Ki (T ' WT I I [ 3 ljk 3 2Jk>J where Tj^ is the optical path defined by (y,a) and T2 is that defined by (y, CC+TT). The Dancoff correction is evaluated for a representative fuel pin in each ring of fuel pins and the effective surface for ring i is the s - Pi Ft The equation R = A+B f-M — i« used to calculate the cluster average \ t resonance integral with St equal to the sum of the effective surfaces and Mt the sum of the weights for each distinct ring of fuel pins. The second form of the relationship between effective resonance integral and effective surface is used to define the effective resonance integral for each ring of fuel pins. but note that the ratio -— is unknown. - 105 - The cluster average resonance integral can be defined to be n 1 r and then A+B |^ = A' £- + B' jp t s t giving We now assume that the fJux distribution is to a good approximation flat - 106 - end therefore VS S S We have therefore defined a resonance integral for each ring of fuel pins taking accurate account of the geometry and materials preaant in the lattice cell under the assumption of an approximately flat radial flux distribution. Thermal Neutron Cross Sections The neutrons in a CANDU reactor are well therraalized and hence the thermal spectrum can be weJl characterized by the Westcott convention. Because the thermal group extends only to 0.625 eV a modified Westcott convention is employed. Two parameters are required to characterize the spectrum, they are epithermal index r and the effective neutron temperature Tn. The neutron density spectrum is assumed to be n(E) = N(l-f;F (E)+NfF (E) in e where N is the total neutron density 2 1 E c is the Maxwellian distribution —-TJ-^-T - 107 - F is the epithermal 1/E distribution -J+ A is a joining function. The Westcott S^ function is used 4.75kT 5 l-.26/{l+(E/16.4kTn) } for which M«3.681 and f is the fraction of neutrons in the 1/E distribution, f is related to the epithermal index r by '••/*' The effective neutron cross section is .62625 .62.625 a = ] n(E)v(E) (E)dE / | n(E)v(E)dE o o The Westcott g function is defined to be 8 = FT" f Fm(E)v(E)a(E)dE 0 0 J o where a0 is the cross section at velocity Vo=2200 m/s. We define a similar function .625 ge " —|- f Fe(E)V(E)a(E)dE - 108 - Both g and ge are functions of the neutron temperature and depend upon the material. The library of nuclear data contains power series fits to g and ge as a function of Tn for each material for each reaction type. Another function dependent only upon Tn ±s also needed VoiiT Fe(E)V(E)dE f o After some algebra one finds {g+f(ge-g)} Because of the importance of the Pu239 and Pv:241 absorptions corrections are made to the cross sections to account for shielding effects in their low lying resonances- These effects reduce the plutonium cross sections and also the cross sections of any other isotopes in the fuel pin. We proceed as follows. The presence of the flux depression due to the resonances modifies the neutron density distribution 1 n(E) == JN(l-f)Fm(E) + NfFe(E)j(f>p(E) From a first order narrow resonance approximation 2 )p(E) = (1+x ) - 109 - 2(E-ER) where and where E is the energy at the resonance peak K F is the width of the resonance a is the effective potential scattering cross section per pluton'um atom a is the cr /ss section at the resonance peak In calculating , Pu241 peak cross section n ,-, where _ „.. "—r ~ Or 0.47 Pu239 peak cross section *• if we now write j> (E) = 1-* then n*(E) = n(E)(l-4>R(E)) where n*(E) is the modified neutron density. - 110 - The reaction rate is then .625 R = f n(E)(l-cf) (E))a(E)V(E)dE .625 E2 R* = n(E)a(e)v(E)dE - n(E)<|) (E)a(E)V(E)dE J J K D o E-L and the neutron flux is .625 Here Ej and E2 represent the range of energies over which the resonance flux depression is assumed to have effect. In the calculation of (g-g*)(l-f) + f(g-g*) 6 6 3^ TTT - Ill - where g* = Q~ j CT(E)Fm(E)V(E)<|)R(E)dE El g* = ~- j a(E)fe(e)v(E)(|)R(E)dE o o ^ f2 Fm(E)v(E) When the generally complex bundle geometry is converted to annuli those annuli which contain fuel pins have cross sections weighted by the thermal flux shape calculated for a single fuel pin. Thermal Neutron Spectrum The thermal neutron spectrum is characterized by two parameters, the epithermal index and the neutron temperature, both of which have to be calculated. We will start with the epithermal index, but remember that the calculation involves iteration. Above 0.625 eV the Maxwellian part of the spectrum is negligible and the neutron flux can be written as - 112 - Q is obtained by fitting the calculated flux for the lower groups of the epithermal calculation to this equation. If NvQ and xn are known from a previous iteration the calculation of the epithermal index can be completed. The calculation of the neutron temperature is the only part of SOLACE which is essentially empirical. It is, therefore, Lhe weakest part o:~ the code; however, for the lattices of interest we have no indication that any discrepancies are due to this calculation. The average neutron temperature in the fuel is defined to be: TNF where T , T , T , T are the physical temperatures of the moderator, coolant, fuel and structural materials respectively. AT , accounts for the spectral hardening in the moderator due to the slowing c'own of fast neutrons AT accounts for the spectral hardening due to absorption of r n neutrons in the fuel. The three remaining terras represent re-t.hermalization effects by the hot materials in the cell. The first term (TC~TM) is the increase in neutron temperature caused by hot coolant. This is the effect of raising the temperature of all materials in the cell up to the coolant temperature. The other two terms represent the additional effect of the higher temperatures in the fuel and sheaths. The moderator hardening is evaluated by an empirical relation derived from the ZED-2 lattice measurements: T = 6 56 + 664 01 r MH ~ * ' M - 113 - The model adopted gives fuel hardening effects in terms of effective temperature changes in a Maxwellian spectrum. It is assumed that the hardened spectrum in the fuel can be described in terms of two overlapping Maxwellians: (i) at the temperature of the incoming spectrum T = (T + AT ) K° LV UM MET (ii) at a temperature TJJ such that [AT ) TH " I T /H where H is the hardened or re-thermalized spectrum U Is the unhardened spectrum. It is further assumed that the 1/v absorption of one neutron can be described as the transfer of one neutron from the Maxwellian with T(j to that with Tg, with an overall reduction in total neutrons by one. The average neutron temperature rise due to hardening is then given by U AT u = (1-K, . a E ) " " ^ ." - T Fh 1 s ^u H i.e. ATFh - 114 - where $„ = neutron flux in hardened Maxwellian (j> = neutron flux in unhardened Maxwellian a = effective fuel bundle radius Z = average scattering cross section within "a" s K = fitted parameter An expression for — is derived assuming that the incurrent of unhardened neutrons at the effective fuel bundle radius a is •Tin» an^ the olackness of the fuel is B. In the definition of Z, unhardened neutrons are lost either by abt;orption ov_ by re-thermalization, characterized by an absorption cross section £aU and re-thermalization cross section £R. In that case 2 „(£ + Zj,}n& = total absorption of unhardened neutr>ns in the u aTuT K bundle 2 Total absorption of unhardened neutrons in a fuel bundle can be expressed in terms of incurrents and blackness as 2ira Hence: Jin B - 115 - Therefore: 2 Jin B The source of re-thermalized neutrons, %,in a hardened spectrum is given b/: 2 SR = wa These neutrons make collisions be.fore they are absorbed. The number of neutrons which are re-therm-jlized by the first collision is: SH -SS P (a TH) H m C TH and the number re-thermalized in the second collision is: £ is the total re-thermalization cross section TI: £S H is the total scattering cross section TH 3n SH P (a £_,.) is the first flight collision probability for a cylinder of radius a and crots section £_.,. in Then, the total number of re-thermalized neutrons which make successive collisions and are absorbed is: - 116 - SH if PC<3 V [1 + if Z IT a SH E — P (a Z ETH C TH Therefore: «S£ P (a ZTH C and substituting for SH: P (a Z ) C TH =§S P (a ZTH C If we consider the loss of unhardened neutrons by absorption (cross section ZaU)) rather than re-thermalization, then there is some feed back from the fictitious hardened spectrum by re-thermalization with a cross section K2ZR. Where ZR is the normal re-thernal- ization cross section and K2 is the fitted parameter. In that case: - 117 - ZSH and Z ~ P (a TH C , ZTH~K2ZR , „ , 1 = (a 2 ) lm TH Re-therraalization cross sections were derived from experiments on re- thermalization in metals. The remaining two parameters, K.. and { •=- \ , were derived using 1 (T )H methods based on unpublished work. Neutron temperatures were calculated for non-moderating rods of various aZs and aZa, by considering absorption probabilities of incident neutrons at a given energy, and integrating over the incident Maxwellian. A neutron temperature was derived which gave the same absorption in a one group / AT ) theory. \ —> was determined by fitting the equation for AT_,, to ( T J rl F n the calculations for ^s=0; Kj was then adjusted to give the best fit to the curves for various aZ . s K2 was derived from ZEEP measurements of neutron temperatures in CANDU type fuel bundles for various coolants. The values are: T H = 0-5 = 0.04 Kl K2 = 0.25 a is defined as the ratio of the change in average neutron temperature in a rod due to a uniform change in physical temperature of the rod, to the change In physical temperature (at constant moderator temperature). The coefficients otp and as are defined - 118 - in a similar way. A two Max:?ellian model is used, with the temperature of the upper one taken as the physical temperature of the hot material in which re-thermalization occurs. For a, the lower temperature is TM> and the upper temperature is the bulk coolant temperature Tc. The hardening effect of the whole fuel assembly uniformly raised to the coolant temperature is: *H TC + *U TM (T - T ) = — - - - - T Kl C V where ^ Va V 1" £R and £7 are the "re-thermalization" and total cross sections for the homogenized rod, averaged over materials at the upper temperature Tc> i.e., R V Rod (i running over all the materials in the rod) I V. E ± i Ri As the effect of all the coolant must be included, "a" is usually taken as the radius of the inner surface of the pressure tube. - 119 - The other two coefficients are calculated in an exactly similar manner, by replacing tc by Tp for «j? and by Ts for ag. Examination of ZED-II experimental data indicates that the increase in neutron temperature in the fuel region above that in the moderator varies rarabollcally with radius. Let a be the radius of the fuel region A be the excess neutron temperature at radius r A be the excess neutron temperature at the pressure tube o then A = A (l-Sr2/a2) where 3=1-1 /(.3425a+.52666) The average temperature difference in the fuel is then 2 2 J| A° (l-gr /a )rdr A r a rdr o = AQ(1-3/2) A is, however, T, -T -AT _ which has just described r NTF in wT H r 2r 1-572 For any region i in the system, having inner and outer radii and vi, the average temperature rise is - 120 - 2 2 (1-gr /aZ)rdr (1-3/2) r± rdr (1-3/2) The neutron temperature in region i within the fuel region is then Ti = [VATnH + AT ) Between the fuel region and the pressure tube the temperature rise is found to vary with r as substituting for A gives 3 and similar algebra results in the neutron temperature in region i between the fuel region and pressure tube being T. = | T +AT + AT + a(T -T ) + ct (T -T ) + a (T -T ) I i \_m m uH FH en FFc ss c _| x 2(l-3)a/ (l-g - 121 - Within the moderator the neutron temperature is assumed constant and so T = T + AT 1 m mH Ne roi: Leakage and Boundary Conditions The Benoist formulation for the diffusion coefficient in direction K is used D i 3 where V is the volume of annulus i <(>. is the neutron flux annulus i A. is the total mean free path in annulus j P . is the probability that a neutron generation annulus i will suffer its first collision in annulus j; k indicates the axial or radial direction The P^j^ are calculated in Lerms of the usual Bickley functions. Axial leakage is taken into account by increasing the absorption cross 2 2 section in each annulus by D B where B is the axial buckling. z z z Radial leakage is accounted for by the boundary condition. Recall that the currents crossing the boundary are related by JN = S+l JN - 122 - ^N+l Is set equal to one then the net outcurrent is JN JN The net outcurrent is, however, £ D. B r and therefore DN+1 = - [ D . B^.V. The presence of tj^ in this equation implies an iterative calculation. The first pass through the code assumes that Dfj+i=O, i.e. a reflective boundary condition. Fuel Depletion The reaction rates are summed for each isotope in each fuel region and are used in the depletion equations. Special treatment is given to Pu240. Due to the presence of the large 1 eV resonance the effective Pu240 cross section is strongly dependent upon the amount of Pu240 present. This is accounted for by separating the reaction rate into two parts resonance and non resonance so that where N^Q ±S the Pu240 concentration, a is estimated from the actual reaction rate over the resonance and the reaction rate that would occur if the resonance did not cause a flux depression. The code is capable of handling a wide selection of burnup chains, Figure B.5. The user specifies which nuclides are to be included in the calculation. - 123 - Fission products are treated in one of two ways. Either equilibrium amounts of Xe , Sm , and Rh with the remaining absorption represented by three pseudo fission products; or the detailed fission product chains shown in Figure B.6 are used with one pseudo fission product representing the remaining absorption. Summary The LATREP family of codes are relatively fast, 2 to 10 seconds of CDC 6600 CPU time depending upon the version of the code, per lattice calculation. Empirical correlations are kept to a minimum, and the effects of resonance shielding In important isotopes, burnup and neutron leakage are treated in a rigorous manner. B.3 COMPARISON OF LATREP WITH EXPERIMENT The LATREP codes are checked against a series of zero energy experiments and measurements at several burnups of the composition of fuel irradiated in the NPD reactor. A summary of the mean error and its standard deviation, taken over the lattice pitch, for a series of natural UO2 fuels with different coolants is given for six parameters in Table B-2. The calculated results are generally within two experimental errors of the experimental values with a few very obvious exceptions. The calculated neutron temperature, as expected, agrees least well of all the parameters. However, its effect is small. Figure B.7 compares calculated and experimental nuclide ratios for eight NPD fuel bundles over a range of burnup from about 980 MW.d/t to about 10,800 MW.d/t. The results are plotted against 1-a . n (N5/N8)t Where a = (N5/N8)o i.e. the ratio of the U235 to U238 concentrations at time t to that at time zero. The agreement is good. - 124 - TABLE B-l Mechanisms for neutron production or loss x,y mechanism neutrons are generated in the annulus by the source or are lost by absorption neutrons enter or leave the annulus via the inner surface neutrons enter or leave the annulus via the outer surface - 125 - LJSL. ELEMENT DiO COOLANT PHHSSURS TUBS AIR GAP CA_-\NCf?IA TUBE 020 M00ERA70R PiGURE B.I LATTICE CELL FOR 37-SLE.ME.N7 FUEL - 126 - Input Setup Initial conditions Cross sections Collision Probabilities Fluxes iterate All groups above 0.625 eV Leakage and Boundary Conditions | Thermal Source iterate r & T Leakage and Boundary Conditions iterate Thermal group Cross Section Collision Probabilities Fluxes Fission Source Distribution Edit Burnup Fuel Cross Sections Depletion Figure B.2 Code Organization - 127 - ANNULUS n FIGURE B.3 COLLISION PROBABILITY NOMENCLATURE - 128 - FIGURE B.4: CALCULATION OF DANCOFF CORRECTION FOR ROD A KEY FISSILE (n,f) £57 , (n.y) REACTION WITH CALCULATED FISSION i PRODUCTS AND YIELDS /3 - DECAY a - DECAY ) ISOMETRIC TRANSITION * - (n.2n> REACTION INETANTENOUS DECAY T.UICTRON CAPTURE N (UNLESS OVERRIDEN) I THERMALLY FISSILE (n.fl I to I \ 83 PO V*PO»«/PO*» /ho*^ Po«5/ FIGURE B.5: ACTINIDE CHAINS KEY LU m 49 4B o § 47 i t—' UJ 46 O 1 45 ui 5 -I 44 o 43 42 41 -J40 FIGURE B.6: EXPLICITLY DETAILED FISSION-PRODUCT CHAINS - 131 - BUNDLE AVERAGE NUCLIDE RATIOS COMPARISON OF CANADIAN CALCULATIONS AND EXPERIMENT FIGURE B.7 239Pu/238U 0.9| 24QPu/239Pu 0.8 241Pu/23SPu 242Pu/233Pu U.I *, n.5f- 4?i=?rrr3zZ- 1 I i... L I 0 01 0.2 0.3 0 4 0.5 0.6 0.7 0. - 132 - C. CANDU ECONOMICS C.I INTRODUCTION The purpose of this lecture is to: - discuss the cost of producing electricity from CANDU-PHW nuclear generating units, - present actual cost experience of CANDU units in Ontario Hydro, - compare CANDU cost experience with fossil (coal) experience in Ontario Hydro, - present projected CANDU and fossil cost data in Ontario Hydro, and - present cost estimate comparisons of CANDU and Light Water Reactors (LWR) in Ontario Hydro. C.2 COST CRITERIA The Cost Object of Ontario Hydro is to produce and deliver electricity at the lowest long-term cost to Ontario customers, while satisfying the other important objectives: - Worker Safety - Public Safety - Environmental Protection - Reliability If a comparison is made between two alternative types of generation, the degree to which all of those objectives are satisfied should be considered. The load of Ontario Hydro (as with most electrical utilities) varies with time. Loads peak in the daytime Monday to Friday when factories are busy and society is active. In Ontario, the loads are higher during the winter when temperatures are low. - 133 - The most economical generating system for Ontario Hydro is a mix of hydraulic generation, fossil generation, and nuclear generation. The majority of available economic hydraulic resources in Ontario has been developed. New loads must be met by alternative resources of which nuclear and coal are the primary options for the balance of this century. The Ontario Hydro system Load Factor is typically 68% (the ratio of average annual power to peak annual power). Fossil-fired generation is most economical for peak load requirements because of its lower capital and OM&A costs*. Nuclear generation is most economical for base-load application because its higher capital and OM&A costs are more than offset by the very low fuelling costs. All costs presented will be for base-load generation within the Ontario Hydro system. Cost evaluations for generation commitment decisions made by Ontario Hydro are very complex, utilizing present value techniques, uncertainty analyses, load forecasts, reliability assessments, environmental impacts, etc., that are beyond the scope of this lecture. The Total Unit Energy Cost (TUEC) method is a simple and accurate indicator of the relative economics for base-load application, and is used here. *OM&A: Operating, Maintenance, and Administrative Costs. - 13A - Total Unit Energy Cost (TUEC) The cost of producing electricity from generating stations involves the following cost classifications: - the research and development of generation concepts, - the cost of building the stations, - the cost of operating and maintaining the stations, - the cost of fuelling the stations, - the cost associated with disposal of the stations at the end of their useful life, and - overhead costs to support the above cost classifications. In addition, the cost of producing electricity must also consider: - the method employed for financing and amortizing the investments, - the interest rates applicable to the above classifications, - the lifetime assumed for the facilities, - the reliability of the stations to produce electricity, and - the policies that are adopted concerning source of supply, taxes, regulations, etc. The Total Unit Energy Cost (TUEC) is defined as the total annual cost of producing electricity (dollars) divided by the total annual electricity energy produced (kilowatt-hours). TUEC = Total Annual Cost Total Annual Electricity Produced - 135 - The research required to develop the generation concepts and the cost of station disposal have been excluded. These exclusions are not expected to have a serious effect on the absolute costs and relative costs of the generation alternatives, in the long-term for a major program. The four cost components for the CANDU-PHW concept are: 1. Annual Interest and Depreciation on the Capital Cost 2. Annual Operation, Maintenance, and Administration Cost 3. Annual Fuelling Cost 4. Annual Heavy Water Upkeep Cost The three cost components for the Light Water Reactor (LWR) concept and coal-fired stations are: 1. Annual Interest and Depreciation on the Capital Cost 2. Annual Operation, Maintenance, and Administration Cost 3. Annual Fuelling Cost The computation of the Annual Interest and Depreciation Cost depends upon four factors: 1. The Initial Capital Cost and the Capital Modifications Cost 2. The Interest Rate 3. The Lifetime of the Station 4. The Method of Amortization of the Initial Capital Cost and the Capital Modifications Costs. - 136 - The Initial Capital Cost Includes: 1. The Design and Engineering Cost 2. The Construction Cost 3. The Commissioning Cost 4. The Permanent In-Reactor Fuel Charge 5. Tha Heavy Water Inventory 6. Overheads 7. Accumulated Compound Interest During Construction The Initial Capital Cost includes the Permanent In-Reactor Fuel Charge (one-half of the Initial Fuel Charge) and the Heavy Water Inventory. The Initial "Dry" Capital Cost is identical to the Initial Capital Cost except that the Permanent In-Reactor Fuel Charge, the Heavy Water Inventory and Commissioning are included. The Annual Operation, Maintenance, and Administration Cost includes: 1. Labour 2. Materials 3. Purchased Services 4. Interest on Operating and Maintenance Inventories 5. Overheads (including taxes) The Annual Fuelling Cost includes: 1. Fuel (quantity and price) 2. Interest on Inventory 3. Transportation 4. Overheads - 137 - The Annual Heavy Water Upkeep Cost comprises two basic factors: 1. The cost of replacing any heavy water lost during operation. 2. The cost of upgrading any heavy water which becomes downgraded during operation (diluted with ordinary water). The total Unit Energy Cost (TUEC) is the sum of the Unit Energy Cost (UEC) for each of the cost components. As an example, the Fuelling Unit Energy Cost is as follows: Fuelling Unit Energy Cost = Fuelling Annual Cost Total Annual Electricity Energy Produced The Total Unit Energy Cost is very dependent on the Capacity Factor* achieved. The Total Annual Electricity Energy used to determine TUEC may be either the gross or net electricity produced. Ontario Hydro prefers to use net energy - TUEC (net). However, for some utilities only the gross production is published and TUEC (gross) is determined. j,_ . Actual Energy Produced *Capacity Factor = —-—-z—_ _^ .—— r Perfect Production for any specified period. - 138 - The Specific Capital Cost is the Total Initial Capital Cost ($) divided by the Net Capacity (kW) and is expressed in dollars per kilowatt. All costs are Canadian dollars unless otherwise noted. 2. Ontario Hydro Cost Comparison — CANDU-PHW Versus Fossil (Coal) The cost comparison between CANDU-PHW units and alternative sources of generation will depend upon many factors which are particular to the electrical utility making the comparison. Nuclear fuel cost tends to be independent of the distance between the uranium source and the generating station because transport cost of nuclear fuel is small. In the case of coal, the transport cost is low if the generating unit is near the coal mine, but can be very high if the coal has to be transported a great distance. Tl following data illustrates that the CANDU-PHW is very competitive within Ontario Hydro where economic hydroelectric resources have been almost fully developed and where coal must be transported a minimum of 800 kilometres. There are other locations in Canada in which coal-fired generation is cheaper than CANDU-PHW where the generating unit is near the mine. The high current and projected cost of oil and gas makes their use uneconomical for base- load generation in Ontario Hydro. More specifically, the Ontario Hydro Pickering Nuclear Generating Station-A (NGS-A) will be compared with the Ontario Hydro Lambton Thermal Generating Station (TGS). The Pickering NGS-A comprises four 515 MW (net) nuclear units of the CANDU-PHW type. The Lambton - 139 - TGS comprises four 495 MW (net) units which burn coal. Both stations were built at the same time, both are of modern design, and both stations are fully operational with good performance records. For the year 1980, Pickering NGS-A had a Net Capacity Factor of 82.6%. Table C-l illustrates the Unit Energy Costs (UEC) of these two stations. The following should be noted from Table Crl: - the coal-fired capital cost is much lower than the nuclear capital cost, - the coal-fired OM&A cost is lower than the nuclear OM&A cost, - the nuclear fuelling cost is very much lower than the coal-fired fuelling cost, - the heavy water upkeep cost, which applies only to the nuclear, is only a small percentage (about 4%) of the Total Unit Energy Cost, and - for base-load application, Pickering NGS-A had approximately one- half the Total Unit Energy Cost of Lambton in 1980. During the period up to 1980, in-service capital modifications have been made to both Pickering NGS-A and Lambton TGS. These modifications are amortized on a remaining lifetime basis and are included in the comparison. It is expected that further capital modifications will be required from time to time to replace major components and meet new requirements. - 140 - For example, the pressure tubes at Pickering NGS-A may have to be replaced and Lambton may have to be retro-fitted with SO2 scrubbers to meet acid rain requirements. Figure C.I presents the Pickering versus Lambton Unit Energy Costs (assuming Lambton operated at the same high capacity factors as Pickering) for each year 1975 to 1980 inclusive. This graph clearly shows the steadily increasing cost advantage of the nuclear plant due to che continuing Inflation of coal costs. This is an example of the "inflation-proof" characteristics of the CANDU-PHW. This graph also shows that the cost of Heavy Water Upkeep is only a small component (about 4%) of the Total Unit Energy Cost. The base load cost (TUEC) of the Pickering NGS-A has been consistently well below the cost of the Lambton TGS (coal-fired). This cost advantage is expected to increase as fossil fuels become more expensive. 3. CANDU Costs Versus Time The actual or estimated TUEC for in-service stations and stations under construction will vary with time due to a variety of reasons including: - escalation of labour and material costs, - changes in interest rates, - escalation of fuel costs, - 141 - - changes in design and operating requirements, - changes in operating performance, and - competence and maturity of workforces (design, manufacturing, construction, operation). The Pickering NGS-A and the Lambton TGS were built in the late 1960s and placed in-service in the early 1970s. During the 1970s, high inflation caused Capital, OM&A, and Fuelling Costs to be driven rapidly upwards. As a result, new coal-fired generating stations such as Nanticoke TGS (8 x 490 MW net) and new nuclear stations such as Bruce NGS-A (4 x 740 MW net) have higher capital costs. In addition, the TUEC of the in-service coal-fired station, Lambton TGS (4 x 495 MW) and the in-service nuclear station, Pickering NGS-A (4 x 515 MW), are rising due to inflation in OM&A and fuelling costs. The Specific Capital Cost of Bruce NGS-A compared with the Specific Capital Cost of Pickering NGS-A is affected by three major factors: - Bruce NGS-A has lower costs due to larger unit size, - Bruce NGS-A has higher costs due to new regulatory requirements, - Bruce NGS-A has much higher costs due to inflation of labour and materials. The result is that the Pickering NGS-A Specific Capital Cost was 362.4$/kW (net) and Bruce NGS-A was 662.5 $/kW (net). Pickering NGS-A came into service between 1971 and 1973, while Bruce NGS-A came into service between 1977 and 1979. - 142 - Table C-2 shows Bruce-A Unit Energy Costs in 1980, while Figure C.2 shows the lifetime trends (1977 to 1980). Table C3 presents the actual Initial Capital Cost of Pickering NGS-A and Bruce NGS-A together with the estimated costs of three nuclear stations under construction — Pickering NGS-B, Bruce NGS-B, and Darlington NGS-A. 4. Ontario Hydro Cost Projections The CANDU-PHW at Pickering NGS-A has demonstrated major cost advantage for base-loaded application in Ontario Hydro in the 1970s. The TUEC has been projected for CANDU-PHW and coal-fired stations for the period from 1980 to 2000. These projections exclude the possible retrofit of S02 scrubbers in coal-fired stations and exclude possible major retrofits in nuclear stations to meet new requirements. Figure C.3 compares actual TUEC for Pickering NGS-A with Lambton TGS (assuming base-load application) for the period up to 1980. It also compares the forecast TUEC for these stations for the period from 1980 to 2000. Figure C.4 displays forecast TUEC for base-load application of five stations currently in service: Coal-Fired - Lambton TGS (4 x 495 MW) - Nanticoke TGS (8 x 490 MW) CANDU-PHW - Pickering NGS-A (4 x 515 MW) - Bruce NGS-A (4 x 740 MW) Oil-Fired - Lennox TGS (4 x 495 MW) - 143 - These projections indicate: - that the base-load advantage of CANDU-PHW is expected to increase with time, and - the "inflation-proof" characteristic of CANDU-PHW. 5. Ontario liydro Cost Comparison — CANDU-PHW Versus Light Water Reactors (I.W3&) The Ontario Hydro nuclear program to date has been limited to experience with CANDU-PHW units- Ontario Hydro has exchanged cost information and operating performance with other utilities in the USA, Europe, and Asia. In particular, this information applied to alternative nuclear types — Light Water reactors (LWR) and Gas- Cooled Reactors (GCR). At the present time, the LWR is the only viable nuclear alternative to CANDU-PHW in Ontario Hydro. The LWR has two basic options — the Pressurized Water Reactor (PWR) and the Boiling Water Reactor (BWR). Because Ontario Hydro has had no design and operating experience with LWR, the cost comparisons between CANDU-PHW and LWR must be based upon the following: - comparison of costs reported by other utilities for LWR with Ontario Hydro costs for CANDU-PHW, and - estimates of CANDU-PHW and LWR assumed to tie built in Ontario under Canadian licensing requirements. - 144 - The judgments expressed below are based upon the following: - the detailed insight Ontario Hydro possesses on CANDU-PHW with regard to cost, - the detailed insight Ontario Hydro possesses on CANDU-PHW with regard to performance, - capital cost information on LWR units built in the USA, and extensive discussions with utilities in the USA, - detailed performance information on LWR units throughout the world, - interpolative judgment regarding expected LWR costs and performance of LWR in Ontario. CANDU-PHW Costs — Ontario Hydro The actual cost data and projected cost data for CANDU-PHW units in Ontario have been presented above. The eight commercial CANDU units have demonstrated a net Capability Factor of 77% (from first electricity production) and 79% from the In-Service Dates. Capital Cost Information - LWR - USA The actual or estimated initial Specific Dry Capital Costs for LWR units of 500 MW and greater built in the USA are shown in Figure C.5 based on information provided by a number of USA utilities, in USA dollars. - 145 - The actual or estimated initial Specific Dry Capital Costs for CAi\"l)U-PUW units in Ontario Hydro are also shown in Figure C.5 in Canadian dollars. The following observations may be made: (a) There is a wide scatter In the Specific Capital Cost data. (b) The CAND'J-PHW units have a simile*- cost and cost trend compared with LWR. LWR Performance - Actual The actual LWR lifetime performance has been reviewed and documented by many people. The lifetime average Gross Capacity Factor of PWR and BWR has been 57% and 55%,respectively. In the following comparisons, we have used the PWR performance of 57%. Ontario Hydro's Jjdgment Through examination, discussion, and stuc!y of available data, the following judgments have been made (a) Performance There is a wide range of PWR perforuance (Capacity Factor) with a world average of 57%. The PWR Capacity Factor performance is expected to further improve. The available data is based upon stations which usually have only one or two units per station. Four-unit stations (Ontario Hydro practice) are expected to have better performance due to improved diversity (spare parts, technical support, etc). - 146 - Ontario Hydro feels that they enjoy better than average project management and operator training. Ontario Hydro believe that If they had an extensive PWR program (same number of units in service, same number of units per station and in-service dates similar to the actual CANDU- PHW program) the expected PWR performance in Ontario Hydro would be 67%. This Is to be compared with a typical coal-fired capability of 74% in the USA and a demonstrated CANDU-PHW performance of 77% (since first electricity production). Both coal-fired and CANDU-PHW stations enjoy the advantage of on-power fuelling. The expected 10% superiority of CANDU-PHW assumes a judged 6% credit for on-power fuelling and a 4% credit for other concept advantages. Ontario Hydro's performance for coal-fired stations is similar to USA experience at 74% for 500 MW units. Ontario Hydro believe that they could achieve 77% in coal-fired stations if staffing levels and spare part diversity were increased. However, this is not economically justified for peak load application. (b) Capital Cost The above comparison of Ontario Hydro and US utility data (Figure C.5) Indicates no major Specific Dry Capital Cost difference between CANDU-PHW units built in Ontario and LWR units built in the USA. - 147 - Examination of the design requirements of the two concepts suggests there should be no major Dry Capital Cost differences for most facilities such as site, turbine-generator, cooling systems, instrumentation and controls, buildings and containment. Assuming Identical supply capability and manufacturing volume, the CANDU reactor with on-power fuelling should be less expensive due to the absence of enriched fuel, very demanding pressure vessel specifications as compared with pressure tubes, the need for in-core high pressure regulating and shutdown devices and the like. However, the cost comparisons which follow assume the Dry Capital Cost for CANDU-PHW and LWR in Ontario Hydro to be identical. (c) OM&A Costs The Capacity Factor achieved by Ontario Hydro has depended, in part, on maintaining around-the-clock maintenance staff at four-unit stations. This high staff level is economically warranted because of the high cost of burning coal whenever a CANDQ unit is shut down. Example — in the Bruce NGS-A, one- percent Capacity Factor is eq;- valent to the wages of about 100 people. Since PWR units also have a low fuelling cost compared with coal, around-the-clock maintenance would also be justified. There is no significant difference in optimum staff levels for four-unit CANDU-PHW and four-unit PWR In Ontario Hydro. Similarly, the total OM&A cost is expected to be the same. - 148 - (d) Fuel Examination of available USA data suggests that, for a large PWR program in Ontario, the Fuelling Unit Energy Cost for PWR would be 5.46 m$/(kW.h) or higher in 1980. This evaluation assumes the mining and refining costs of natural uranium are identical for CANDU and PWR. Enrichment cost is peculiar to the PWR units. Fabrication costs are particular to each design. Comparison - CANDU-PHW Versus PWR - Ontario Hydro Table C-4 is a cost comparison of CANDU-PHW and PWR in Ontario Hydro. It is based on actual cost and performance experience of CANDU units in Ontario and the estimated performance and cost of PWR as outlined above. Two sets of estimates are given for the PWR, corresponding to the world average capacity factor of 57% and the judgment of 67% for a major PWR program in Ontario Hydro. This comparison indicates that: - the TUEC for PWR operating in Ontario at the world average Capacity Factor (57%) would be approximately 37% higher than the TUEC of CANDU-PHW actually experienced. - the costs of Capital Heavy Water plus Heavy Water Upkeep for CANDU are more than offset by the higher Fuelling UEC of the enriched PWR fuel. - 149 - 7. Summary 1. Nuclear-Electric and Coal-Electric Generating Stations are the primary options in Ontario for new generating requirements for the period from 1980 to 2000. 2. Coal-Electric Generating Stations are the best choice to meet peaking requirements. 3. Nuclear-Electric Generating Stations are the best choice to meet base-load requirements. 4. For base-load applications, the CANDU-PHW has a proven lower cost than Coal-Electric. In 1980, the Total Unit Energy Cost for Pickering NGS-A was 12.55 m$/(kW.h) compared with the same size, same vintage Lambton coal-fired station with a corresponding cost of 21.08 m$/(kW.h). 5. CANDU-PHW has inflation-proof characteristics due to its very low fuelling costs. 6. CANDU-PHW with on-power fuelling is expected to continue to enjoy high Capacity Factor performance. 7. The actual performance of CANDU-PHW units in Ontario is 77% Capacity Factor since first electricity production and 79% since the In-Service Dates. The actual average lifetime performance of PWR in the world is 57% since first electricity production. Ontario Hydro believe that 67% LWR Capacity Factor is achievable with a mature LWR program. 8. For Ontario conditions and requirements, the estimated Total Unit Energy Cost of LWR units as compared to CANDU-PHW experienced costs indicates the LWR to be 26% higher, assuming a 67% LWR performance and a 77% CANDU performance. - 150 - TABLE C-l PICKERING/LAMBTON COST COMPARISON - 1980 Pickering and Lambton Net Capacity Factor: 82»6%* UEC tn$/(kW.h)(net) Pickering NGS-A Lambton TGS Interest and Depreciation 6.01 1.94 Operation, Maintenance, and Administration 3.77 1.57 Fuelling 2.33 17.57 Heavy Water Upkeep 0.44 - Total Unit Energy Cost (Net) 12.55 21.08 Station Data Pickering Lambton Capacity (Maximum Continuous Rating) MW(e) net 4 x 515 4 x 495 In Service 1971 - 1973 1969 - 1970 Initial Capital Cost (M$ Canadian escalated) 746.5 257.0 Specific Capital Cost ($/kW) 362.4 129.8 Economic Lifetime (years) 30 30 Depreciation Method Straight Line Interest Rate (%) 9.9 9.9 ^Assumes Lambton also operated at base load with net capacity factor of 82.6%. Lambton actual 1980 net capacity factor was 56.6%. - 151 - TABLE C-2 Bruce NGS-A 1980 Costs Net Capacity Factor = 86.7% UEC m$/(kW.h) Interest and Depreciation 9.88 Operation, Maintenance and Administration 2.62 Fuelling 2.67 Heavy Water Upkeep 0.39 15.56 Station Data Capacity (Maximum Continuous Rating) MW(e) net 4 x 740 In Service 1977 - 1979 Original Capital Cost (M$ Canadian escalated) 1 961.1 Specific Capital Cost ($/kW) 662.5 Economic Lifetime (years) 30 Depreciation Method Straight Line Interest Rate (%) 9.9 - 152 - TABLE C-3 Nuclear Capital Poet Data (Net) Actual Initial Net Capital Specific Dry* Specific Dry* Capacity Cost Cost Capital Cost Capital Cost Year In Station MJ(e) M$ $/kW M$ $/kW Service Pickering NGS-A 2 060 746 .5 362.4 565.7 274.6 1971-1973 Bruce NGS-A 2 960 1 961.1 662.5 1 498.9 506.4 1977-1979 Estimated Initial Net Capital Specific Dry* Specific Dry* Capacity Cost Cost Capital Cost Capital Cost Year In Station $/kW M$ $/kW Service Pickering NGS-B 2 064 2 810.3 1 361.6 2 076.0 1 005.8 1983-1984 Bruce NGS-B 3 024 4 350.9 1 438.8 3 232.0 1 068.8 1983-1987 Darlington NGS 3 524 6 695.6 1 900.0 5 490.5 1 558.0 1988-1991 *Dry capital costs exclude heavy water, fuel, and commissioning. - 153 - TABLE C-4 ONTARIO HYDRO 1980 CANDU-PHW VERSUS PWR COSTS CANDU PWR Average PNGS-A High NCF NCF Station Size 2 060 2 060 2 060 (MW(e) net) Net Capacity Factor 77 67 57 (NCF %) Capital UEC Dry Capital 4.89 5.62 6.61 Commissioning 0.17 0.20 0.23 Fuel 0.07 0.35 0.41 Heavy Water 1.32 _ _ Capital UEC 6.45 6.71 7.25 OM&A UEC 4.04 4.64 5.46 Fuelling UEC 2.33 5.46 5.46 Heavy Water Upkeep UEC 0.47 _ Total UEC 13.29 16.81 18.17 Note: All UEC data in m$/(kW.h) 1980$. - 154 - FIGURE C.I TOTAL UNIT ENERGY COST COMPONENTS THERMAL VERSUS NUCLEAR (1975 - 1980) 22 P = Pickering (Nuclear) 4x515MW(e) 20 L = Lambton (Coai) 4 x 4S5 MW(e) IS 16 - 14 - 222 8 - 6 - 4 - 2 - 0 P L P L P L P L P L P L 1975 1976 1977 1978 1979 1980 LEGEND • I&O Heavy water upkeep Q O, M & A Fuel - 155 - FIGURE C.2 BRUCE NGS-A - LIFETIME TRENDS TOTAL UNIT ENERGY COST NET CAPACITY FACTOR Im S/kWh) NCF (%) 100 80 - 60 - 40 20 n 76 77 78 79 80 76 77 78 79 80 OPERATIONS, MAINTENANCE INTEREST & DEPRECIATION & ADMINISTRATION FUELLING UEC ImS/kWh) UEC !mS/kWhl UEC (mS/kWh) 76 77 78 79 80 76 77 78 79 80 76 77 78 79 80 FIGURE C.3 TOTAL UNIT ENERGY COST For Pickering and Lambton m$/kWh 100 BASIS. STRAIGHT LINE DEPRECIATION 90 ESCALATION FORECAST OF OCTOBER 1978 80 70 ACTUAL •«- -• FORECAST 60 80% ANNUAL CAPACITY FACTOR 10% INTEREST RATE 50 40 30 20 PICKERING 'A' (Nuclear) 10 1970 1975 1980 1985 1990 1995 2000 Figure C.4 PROJECTED TOTAL UNIT ENERGY COST For Major Operating Thermal Stations mS/kWI) 100 BASIS: STRAIGHT LINE DEPRECIATION ESCALATION FORECAST OF OCTOBER 1978 90 A COV- ANNUAL CAPACITY FACTOR 10% INTEREST RATE 80 i 70 I I-1 60 ^ I 50 40 H 30 20 r~" BRUCE 'A' (Nuclear] 10 PICKERING 'A' (Nuclear) I 1 1 1 1 1980 " 19'85 1995 2000 FIGURE C.5 IIMTERUTILITY COMPARISON - NUCLEAR PROJECTS SPECIFIC DRY CAPITAL COSTS Z5U0 • SINGLE UNIT PLANT —7~ O TWO UNIT PLANT A MULTIPLE UNIT PLANT .- X ONTARIO HYDRO PLANT 2000 c o O o X CMCRLir> .. a JGTC N 1500 n o ) C I] / O o o O 00 X I PICK ERING'B' < . • 81 UCE ;£^< o c 1000 w o -^ u O^ n ' "r rp 3 -tr —-•-- . s 1UCE D00 XB —- — a o PIC JG 'A a - ' y iJ If? ^ 0 1970 71 72 73 74 75 76 77 78 79 80 81 82 83 84 85 86 87 88 89 90 91 92 93 1994 MIDPOINT OF IN SERVICE DATES CANADIAN STATIONS IN CANADIAN DOLLARS OTHER STATIONS IN US DOLLARS s - 159 - D. THE LOW ENRICHED URANIUM FUEL CYCLE The use of natural uranium is the obvious route to follow due to the cheapness of the fuel. However, the amount of fissile material in natural uranium is not ideal. Figure D.I shows the energy derived per gram of U235 fed to a CAMDU-PHW on a once-through cycle as a function of the uranium enrichment and for several tails enrichment. The energy produced peaks at about 1.5% enrichment but the maxima is quite flat and enrichments of around IX are almost as good. If natural uranium had this sort of enrichment we could derive 30% to 40% more energy from a given amount of U235 using a once-through cycle CANDU-PHW. This is the incentive to investigate low enriched fuel cycles. The extra enrichment, however, has its drawbacks particularly in its effect on fuel management and power peaking. D.I SOME EFFECTS OF ENRICHMENT When a channel is refuelled the average fission cross section is increased which causes an increase in the heat generated in the channel above the normal average value. This refuelling ripple effect, of course, increases with the enrichment of the fuel assuming that the same number of fuel bundles are replaced. These effects are investigated with a supercell, two-group neutron diffusion calculation. This calculation has incorporated in it a thermal hydraulic simulation of the fuel channel so that changes in fuel temperature, coolant density and temperature along a channel are calculated and allowed for. The supercell is a group of channels of size nxn, using a square lattice pitch, where n is an odd number, usually seven. Reflective boundary conditions are assumed at the outside boundary of the supercell and the supercell power is held constant to simulate the zone control system. The choice of an odd number of channels along each side of the supercell - 160 - results In a latt. ^e with a central fuel channel. Equilibrium conditions, taking into account the fuelling scheme, are calculated and assigned to all but the central channel* The central channel has allocated in turn conditions appropriate to a freshly fuelled channel, an equilibrium channel and a channel about to be refuelled, so that fuelling ripple effects and power histories can be obtained. Besides this worsening of the refuelling ripple the increased enrichment increases the burnup of the fuel. The increase in the burnup results in a reduction of the amount of spent fuel to be disposed of and in a reduction in fuelling frequency or fuelling machine load. The worsened refuelling ripple results in decreased hydraulic margins and more severe bundle power histories leading to greater fuel failure probabilities. The reduction in fuelling machine load and worsened refuelling ripple can, to some extent, be traded off. If the fuelling frequency is increased to near the value obtained for natural uranium by reducing the number of fuel bundles loaded at each fuelling operation, the power ripple can be reduced. The number of bundles that can be loaded per fuelling operation without exceeding the fuelling machine load appropriate for natural uranium fuel is illustrated in Figure D.2 as a function of enrichment. Two bundle fuelling covers the range of enrichment which includes the maximum In the energy produced per gram of U235. A measure of the size of the refuelling ripple is obtained from the power of the maximum rated bundle. This is also a factor in the fuel failure probability. Figure D.3 shows the maximum bundle power normalized to that for a four bundle push through natural uranium fuelling scheme as a function of enrichment. The lowest enrichment for a given number of bundles loaded per fuelling operation gives the smallest refuelling ripple. It is evident that an enrichment of about 0.93% by weight - 161 - along with two bundle push through refuelling is the best from the viewpoint of refuelling ripple and energy extracted per gram of U235. D.2 AXIAL POWER DISTRIBUTIONS Axial power distributions have been calculated for two enrichments, 0.93% and 1.2%, as well as for natural uranium using the methods noted previously. The effect of the xenon override adjusters, which also contribute to the axial flattening of the natural uranium reactor has been included by distributing a uniform poison over the centre six bundles in the twelve bundle fuel channel. This represents the overall flux shaping role of the adjusters but minimizes local peaking effects due to the discrete absorber banks. In the following figures the axial distribution of linear heat rating is plotted for a freshly fuelled, equilibrium and about to be refuelled channel. The natural uranium result is for a four bundle push-through fuelling scheme while the 0.93% and 1.2% enriched cases used a two bundle push-through fuelling scheme. Figure D.4 is for the natural uranium fuel and indicates quite well the nicely flattened axial power distribution and the quite small fuelling ripple. Figure D.5 is for the 0.93% case and included in the figure are the equilibrium natural uranium distribution and the situation that would exist if there were no adjuster rods. The power distribution with the adjusters is quite distorted and shows a larger refuelling ripple than for the natural uranium. Note that the power distribution without adjusters shows no central dip. This implies that adjusters can only worsen the axial form factor even if their position in the core is changed. - 162 - Figure D.6 is similar to Figure D.4 but for the 1.2% enriched fuel. The refuelling ripple is now very pronounced and the power distribution is severely distorted. With no adjusters the power distribution still exhibits a dip in the centre of the channel. This implies that if the position of the absorbers was to be changed then the power distribution could perhaps be improved. An investigation of fuel failure probabilities indicates failures are likely with the 1.2% enriched fuel when the fuel in position 1 is moved through the power peak on its way to position 3 and that a power derating of about 10% would be needed to reduce the failure probability to acceptable levels. D.3 REACTIVITY COEFFICIENTS Table D-l summarizes the results of calculations for natural uranium, 0.93% and 1.2% enriched fuels. The differences are quite small and the overall effect is such that the effect on system controllability is extremely small. D.A IMPACT ON ECONOMICS AND RESOURCE UTILIZATION There is a clear advantage to be gained through the use of low enriched uranium fuel in terras of mined uranium requirements ever, when the uranium for the initial fuel charge is taken into account . The economics impact, due to both the fuelling and reactor capital charges, is less certain. This is due to the uncertainty in the cost of fabricated fuel and also due to the possible reactor channel power derating at the higher enrichment. The impact on both economics and resource utilization will be discussed for both siagle station and system studies, The base values of the economic parameters are given in Table D-2. - 163 - D.4.1 Single Station Studies I).4.1.1 Resource Utilization The 30 year lifetime mined uranium requirements, excluding processing losses for various enrichments, are given in Figure D.7. In this analysis equilibrium refuelling is assumed to start after 0.36 of a core lifetime. The core lifetime is defined as the minimum time Interval for all the core to be refuelled. It can be seen that the total mined uranium requirement is insensitive to enrichment over the range 0.85 to 1.4% having a value of about 1900 t u compared to 2600 t u for a natural uranium core, a saving of about 27%. 15.4.1.2 Economic Assessment All costs are quoted in 1978 dollars capitalized and discounted to the year of reactor in service. The costs that vary with enrichment are the equilibrium fuelling cost and the initial core inventory charge. Reactor derating would Increase the initial capital charge due mainly to the increased core uranium and D2O inventories associated with the larger core. The variation of fuelling cost with enrichment is shown in Figure D.8. Minimum costs occur near an enrichment of 1.0% assuming no derating. The total unit energy cost Is sensitive to many parameters. Some of these are given in Table D-3. - 164 - At an enrichment of 0.93% when no derating is required the improvement in the total unit energy cost over that for natural uranium is in the range 0.60 to 0.65 m$/(kW.h). This is the same for 1.2% enrichment; however, a 10% derating at this enrichment would lead to a reduced cost advantage of at best 0.15 m$/(kW.h). D.4.1.3 Summary of Single Station Studies Use of 0.93% enriched fuel which requires no derating would produce an improvement in both the total unit energy cost and the resource utilization. There is no Incentive to go to higher enrichments as neither the total unit energy cost nor the resource utilization show significant improvement and the possibility of derating would remove any advantage. D.4.2 System Studies System studies have been performed to compare the use of natural uranium and low enriched uranium fuel in an expanding Canadian nuclear power program. The base economic parameters are given in Table D-4. The assumed nuclear power growth rate is shown in Figure D.9. Here all new stations from 1995 on used low enriched uranium fuel. D.4.2.1 Resource Utilization It has been assumed that there is a delay of 3-1/4 years between the time that uranium is mined and the fuel bundle is fabricated. Figure D.10 shows annual mined uranium requirements for a 0.93% enriched fuel cycle and for a natural uranium fuel cycle. The initial increase in the annual demand for uranium above that for natural uranium is due to the increased core inventory for the enriched cycle. - 165 - Figure D.ll shows cumulative mined uranium requirements for a 0.93% enriched fuel cycle and for a natural uranium fuel cycle. There Is no reduction in cumulative uranium requirements before the year 2010 with the system growth scenario which has been used. The use of a 1.2% eririched fuel cycle would lead to a further 5% reduction in uranium demand, but the initial Increase in uranium demand would be larger. D.4.2.2 Economic Assessment Figure D.12 shows how the total unit energy cost for a 0.93% enriched fuel cycle compares with the cost for a natural uranium fuel cycle. Here the total unit energy cost is derived from a levelized cash flow and gives the average value over the life of a station. The maxima and minima in the curves are due to the assumed nuclear system growth rate over the first decade. The second peak for the enriched fuel cycle is due to the extra mined uranium needed to provide the initial core inventory. Eventually under a steady growth rate a differential of 0.6 to 0.7 m$/(kW.h) is seen to develop. This differential would increase slightly if 1.2% enriched fuel were to be used with no derating. D.4.2.3 Impact on Advanced Fuel Cycles We have seen that a 0.93% enriched fuel cycle decreases the cumulative mined uranium required by 20-30% but not until after the year 2020. If we accept as a convenient round figure 5x10^ MgU as the total of the Canadian uranium reserves we can see from Figure D.8 that the introduction of the low enriched fuel cycle would only extend these reserves by about 3 years. Thus, although the low enriched uranium cycle will reduce the demand for uranium it cannot have much influence on the perceived need for advanced fuel cycles which would reduce much more significantly the uranium demand. - 166 - These advanced fuel cycles would probably use plutonium thorium fuel. The plutonium production from the low enriched fuel cycles is shown in Tables D-5 and D-6. The plutonium has a lower content of fissile isotopes and there is less of it. However, each bundle of fuel contains more plutonlum which would reduce the reprocessing cost of plutonium. This could have an impact on the advanced fuel cycles. D.4.2.4 System Study Summary The improvement in the total unit energy cost of 0.5 to 0.7 m$/(kW.h) is small but reasonably well assured. Savings in total mined uranium would amount to about 20% but would not becon.2 apparent until after the year 2010. Even so the extension of the Canadian resource base would only amount to some 3 years. From a national viewpoint then there is little incentive to implement the low enriched fuel cycle. However, the picture is rather different for a utility for the uranium saved is w>rth a great deal of money. By the year 2030 we would have, on the scenario presented here, a uranium saving of some 50,000 tonnes worth at least one billion dollars. - 167 - TABLE D-l REACTIVITY COEFFICIENTS Enrichment Natur 1 , 0.93 1.20$ Burnup MW.d/t 0 3800 7500 0 7300 14600 0 10600 21200 Moderator temperature -.098 .048 .112 -.101 .065 .152 -.103 .066 .168 mk/°C Coolant temperature -17.2 11.8 24.9 -15.2 14.2 29.2 -13.0 13.8 31.4 mk/°C Coolant density -23.9 -18.0 -15.4 -25.3 -18.7 -13.6 -26.3 -19.5 -21.5 mk/(g/cm ) Coolant temperature 4 density 34.8 52.0 57.4 40.0 55.8 59.2 44.4 57.2 58.8 mk/°C Fuel temperature -13.5 -4.3 -0.5 -13.6 -3.7 1.7 -13.5 -3.7 2.5 mk/°C Total void effect mk 20.8 15.9 13.5 21.9 16.5 11.8 22.5 17.3 10.8 - 168 - TABLE D-2 BASE VALUES OF ECONOMIC PARAMETERS LEU In 600 MW(e) CANDU-PHW U308 Cost 117 $/kgU Heavy water cost 270 $/kgU Separative work cost 100 $/kgu SWU Enrichment plant tails 0.2% U235 Fuel fabrication cost natural uranium 55 $/kgu enriched uranium 70 $/kgU Plant capacity factor 80% Plant life 30 years Heavy water upkeep rate 0.15 kg/h Fuel disposal cost Immobilization 18 $/kgu Disposal 0.09 m$/(kW.h) Escalation rate 0 Interest rate 4% Discount rate 40% Decomissioning cost 30 m$ - 169 - TABLE D-3 TOTAL UNIT ENERGY COST SENSITIVITY LEU in 600 MW(e) CANDU-PHW Cause Enrichment Change in TUEC Derating 0.85 +0 .012 per % derated 1.00 +0.014 Fuel fabrication natural +0 .021 0.85 +0.013 per $/kgU increase 1.00 +0.010 Enrichment 0.85 +0 .002 per $/kgU 1.00 +0.004 Uranium Feed natural +0 .021 per $/kgU 0.85 +0.018 increase 1.00 +0.017 Burnup natural -0.052 0.85 -0.026 per 100 MW.d/t 1.00 -0.019 increase y - 170 - TABLE D-4 BASE VALUES OF ECONOMIC PARAMETERS FOR SYSTEM STUDIES, FUEL COMPONENT ONLY LEU IN 600 MW(e) CANDU-PHW U,0 cost 117 $/keU Separative work cost 100 $/kgU SWU Enrichment plant tails 0.2% U235 Fuel fabrication cost natural uranium 55 $/kgU enriched uranium 70 $/kgU Fuel disposal cost natural uranium 22.A $/kgli enriched uranium 26.6 5/kgU Fuel holdup times natural uranium mine to fabrication 1.0 year fabrication to reactor 0.5 year enriched uranium mine to enrichment 1.0 year enrichment to fabrication 1.75 years fabrication to reactors 0.5 year - 171 - TABLE D 5 PLUTONIUM CONTENT OF DISCHARGED FUEL Enrichment Burnup g Pu/kgU Isotopic Weight Fraction % MW.d/t Pu239 Pu240 Pu241 Pu242 0.711 7,500 3.93 0.675 0.264 0.048 0.014 0.93 14,600 5.42 0.541 0.344 0.072 0.042 1.20 21,200 6.30 0.477 0.371 0.082 0.070 TABLE D-6 ANNUAL PLUTONIUM PRODUCTION LEU in 600 MW(e) CANDU 80% LOAD FACTOR Enrichment Pu Production +/year % Total Fissile 0.711 0.333 0.241 0.93 0.236 0.145 1.20 0.189 0.106 - 172 - FIGURE D. h EFFICIENCY OF THE ONCE-THROUGH URANIUM CYCLE TAILS or CD ASSAY LJ _) 0.2% to 0.25% Ld LZ _l 0.3% MINE D I (FIS S < PE R V ERG z 0 0 % ENRICHMENT - 173 - FIGURE D.2= BUNDLES SHIFTED AS A FUNCTION OF ENRICHMENT Q LJ X CO C/) Ld _l Q z: Z> 03 .7 .0 % ENRICHMENT - 174 - FIGURE D.3= MAXIMUM BUNDLE POWER RELATIVE TO THAT FOR 4 BUNDLE PUSH THROUGH NATURAL URANIUM FUELLING 2.5 h LU O 2.0 - Q_ LU _1 a 00 LU THE NUMBERS ON THE LL) LINES ARE THE NUMBER or OF BUNDLES LOADED I I I I I I I I .7 1.0 1.5 2.0 % ENRICHMENT - 175 - FIGURE DA- NATURAL URANIUM 4 BUNDLE PUSH THROUGH E o o ca LJJ -J DISTANCE ALONG CHANNEL COOLANT - 176 - FIGURE D.5' 0.93% LOW ENRICHED URANIUM 2 BUNDLE PUSH THROUGH EQUILIBRIUM NO ADJUSTERS E NATURAL URANIUM o EQUILIBRIUM 2 f- < OH EQUILIBRIUM \ < LLJ DISTANCE ALONG CHANNEL COOLANT FLOW —- - 177 - FIGURE D.6= 1.2% LOW ENRICHED URANIUM 2 BUNDLE PUSH THROUGH EQUILIBRIUM E 20 - o NO ADJUSTERS NATURAL URANIUM £T I- lxl X ce Lxl DISTANCE ALONG CHANNEL COOLANT FLOW — - 178 - FIGURE D.7= 30 YEAR MINED URANIUM REQUIREMENT 2600 1800 ENRICHMENT % - 179 - FIGURE D.8= DIFFERENTIAL DISCOUNTED FUELLING COST RELATIVE TO NATURAL URANIUM -1.0 r- -0.8 O O -0.6 CD UJ -0.4 ai -0.2 DC UJ u_ LJ_ Q I 0.8 1.0 1.2 1.4 ENRICHMENT % - 180 - FIGURE D9; INSTALLED NUCLEAR CAPACITY SCENARIO FOR SYSTEM STUDIES 300 h 200 - S <£> 100 - 980 1990 2000 20 10 2020 2030 2040 2050 - J81 - FIGURED. 10: ANNUAL MINED URANIUM NATURAL URANIUM 980 1990 2000 20 10 2020 2030 2040 2050 - 182 - FIGURE D. I h CUMULATIVE MINED URANIUM 10 in O NATURAL URANIUM 0.93% LEV Q LJJ 980 1990 2000 20 10 2020 2030 2040 2050 - 103 - FIGURE D. 12= LEVELIZED TOTAL UNIT ENERGY COST o UJ Q Ixl M _l LU > LJ 980 1990 2000 20 10 2020 2030 ?040 2050 - 184 - E. THORIUM FUEL CYCLES IN CANDU E.I INTRODUCTION Canada's presently known uranium reserves (those costing ^30$/pound to extract) are estimated at about 4.5 x 10^ MgU. Some of this is committed already for export, leaving, in round numbers, some 3 x 10^ MgU for domestic consumption. If used in CANDU's with the once-through, natural uranium cycle, this would be enough to supply the present electrical demand in Canada for about 3 x 105/6 x 103 = 50 years. Canada is relatively well favoured among the nations of the world with respect to the ratio of uranium resources to electrical energy demand. If one does the same type of estimate for the world as a whole, the result is that the known uranium reserves (^30$/lb) would supply the present electrical energy demand for only some 15 years. No doubt more uranium will be discovered but at the same time electrical demand Is rising. The point is that, used this way, this is a relatively small resource on which to pin long-term hopes and major develop- ment. Some more effective way of using uranium would provide insurance that a major program is really worth while. Associated with this supply situation is the question of uranium price. As uranium becomes scarce it is expected that the price will rise - indeed the price has risen sharply over the last few years. The adoption of the natural uranium, once-through cycle was based on low cost uranium and it is possible that there would be a better choice on economic grounds if the price of uranium changes drastically. - 185 - Notice, though, that in talking about reserves it is not the supply of uranium that is in question, but the supply of fissile material. Indeed if all the U238 could be converted to fissile material the resource base would be increased and this, of course, is the argument in favour of the fast breeder reactor. It has been known since the earliest days of reactor physics that U233 is better suited for use in a thermal reactor than any other fissionable material. The n value (neutron yield per neutron absorbed) in a typical CANDU spectrum for U233, Pu239 and U235 is 2.28, 1.94 and 2.04,respectively. This implies that U233 has more spare neutrons available for converting fertile material into fissile material. Unfortunately U233 does not exist in nature and has to be produced via neutron capture in Th232. The abundance of thorium in the earth's crust is known to be approx- imately three times as great as uranium. If all the thorium could be used to produce U233 the resource base of fissile material could be increased compared with the natural uranium base. This then forms the argument for the thorium cycle in CANDU systems. E.2 SOME FEATURES OF THORIUM FUEL CYCLES Since thorium contains no naturally occurring fissile isotope it must be enriched with eithei U235 or fissile plutonium if it is to be used as a reactor fuel. In our studies the uranium isotopes generated in the thorium are recycled back into the thorium after the fission products are removed. This process is repeated until the uranium isotopes come into an equilibrium composition. To this mixture may be added extra quantities of fissile material to give us the "topping cycles" which result in burnup ranging from about 20 to 40 MW.d/t HE for topping additions of from 1 to 5 g of fissile isotopes per kg of heavy elements. Alternatively, no topping need be added and in this case we have what has been designated the self- sufficient equilibrium thorium cycle (SSET); that is to say that once the equilibrium isotopic composition has been established no further supply of mined uranium is required. Typical equilibrium isotopic compositions are given in Table E-l. - 186 - It so happens that both Pu and U235 initiated and topped cycles result in initially identical burnups and have very similar feed and discharge isotopic composition although the utilization of mined uranium by means of enriched uranium is more efficient than by means of Pu extracted from spent fuel. The reason for this is not difficult to trace. In a uranium enrichment plant It is possible to extract %5 g of U235 per kilogram of natural uranium whereas the spent fuel from the natural CANDU reactor contains only 2.7 g of fissile Pu per kg of initial uranium. Of course, the natural uranium fuel has produced useful energy, but the detailed calculaticns which take this into account still indicate that the utilization of rained uranium is more efficient if separated U235 rather than fissile Pu is used. Figure E.I contains relevant information on uranium resource utilization in relation to the type and quantity of fissile material used, fuel burnup and processing losses. It should be noted that processing losses must be kept below 1% if a burnup of ^10000 MW.d/t is to be achieved with the SSET cycle. The SSET cycle is especially sensitive to both processing losses and the neutron economy incorporated into the reactor core. Another version of Figure E.I is shown in Figure E.2 where the effect of a change of 10 mk (l%<5k/k) in the neutron economy is shown. Because of this great sensitivity of the SSET cycle it presents the greatest challenge to the reactor physicist. One of the characteristics of thorium is that its absorption cross section is about three times greater than that of U238; consequently more fissile material is required to establish an equilibrium fuel cycle than would be the case with U238. However, once established, in a thermal reactor the thorium cycle consumes less U235 than a uranium cycle. - 187 - E.3 SOME REACTOR PHYSICS PROBLEMS IN THE THORIUM CYCLE Because of the extreme sensitivity of the SSET cycle all the problems presented are taken from this regime. Why, first of all, does this extreme sensitivity occur? I" Figure E.3 the dashed lines show how the fission and absorption macroscopic cross sections for a natural uranium bundle vary with Irradiation. There is an initial transient in the fission cross section as Pu builds in, followed by a steady decrease as the Pu239 saturates and the U235 component continues to decrease. The absorption cross section shows the same initial rise as plutonium is created, followed by a gradual tapering off in the rate of rise as Pu239 saturates. At long irradiations the positive slope is due to accumulation of the non-saturating fission products. The net effect is that after an initial transient the neutron yield is decreasing and the absorption increasing so that the reactivity vEf/2a decreases. Contrast this with the case of an SSET fuel bundle, where the effect of the U233-Pa233 transient causes absorption and fission cross sections to decrease initially. Equilibrium is soon established, and Ef levels out. It is mainly the gradual accumulation of non-saturating fission products that causes Za to increase faster than £j. As a consequence the ratio v£j/£a drops very slowly so that a slight reactivity increment will affect the maximum burnup prediction profoundly. Roughly, we estimate the dependence to be 500 MW.d/t per mk of reactivity, about 5 times that of natural uranium. And so we must ask, "How well can we calculate fission product poisoning?" In LATREP and SOLACE we have available two fission product options. The first is a set of pseudo fission products, consisting of the three saturating nuclides Xel35, RhlO5 and Sml49 and three pseudo fission products whose yields are fitted to Westcott epithermal indices and flux levels. In fact, Sml49 is really an amalgam of the saturating materials Cdll3, Sml49, Sml5l, Eul55 and Gdl57. The yields and fits for this model were published in the 1960s and tests have confirmed that they give adequate representations for the thermal fission of U235 and Pu239. - 188 - The second option is a transmutation chain containing 48 of the most important nuclides plus one accumulating pseudo fission product. The evaluated chain and direct yields and cross sections generally agree with ENDF/B-IV data. The first-order coupled differential equations are solved by advanced Runge-Kutta techniques, and there is no assumption of saturation - all transients are explicitly followed. We also have the fission product accumulation code FISSPROD-2 available, which contains 799 explicit fission products and a library based largely on ENDF/B-IV. We have performed a number jf comparisons between the three models (Figure E.4). The total fission product poisoning for thermal fission of U233 in barns/fission as a function of irradiation calculated by FISSPROD-2 is shown, together with the differences from this curve obtained using the two LATREP options. It is seen that the old pseudo fission product option is 6-8% high, whereas the detailed fission product representation is quite adequate. Our conclusion from this and other tests is that 48 nuclides plus one pseudo fission product is an adequate representation for a very detailed calculation of fission product poisoning in thermal reactors, and the older pseudo fission product model is adequate only for U235 and Pu239. For U233 fuel cycles (Figure E.5) a 15% difference In the predicted burnup can be directly attributed to the 6-8% discrepancy between the two methods of calculating fission products. The predictions of fission product poisoning are possibly 10% higher than experiment, but it would take a major change somewhere in the nuclear data to resolve the discrepancy, and it is difficult to see where this could be. If the discrepancy between experiment and prediction could be resolved, there may still be another 15-20% increased burnup in reserve, provided the discrepancy is in the absorption during irradiation. - 189 - The next item that has a significant effect on the predicted properties of thorium-uranium fuel is the fundamental heavy element data. As might be expected in fuel where U233 is the major fissile specl°s, it Is the fissLon cross section of U233 to which results are most sensitive, followed closely by the capture cross section. Our knowledge of the cross sections is not very good. To get an idea of what we're up against, let's have a look at the energy dependence of n for U233 over the whole energy range from thermal to several MeV (Figure E.6). The dashed curve, labelled material 1041, has appeared extensively in review articles over the past 3 or A years to Illustrate the utility of U233 as a nuclear fuel. This curve was obtained using ENDF/B-II as a source of data. The upper curve labelled material 1260 is obtained from ENDF/B-IV data. Between about 70 eV and 2000 eV there is a significant difference between the two data sets, although both evaluations cite the same Oak Ridge experiment as the source of data in this energy range. Although discrepancies In r| In this range will have little effect ^n the design of thermal reactors, it taught us to be a bit discerning in our choice of data, so we had a look at the 2200 m/s cross section data for U233 (Table E-2). Listed here are four sets of evaluated data, all from reputable sources. The thermal value of a is reasonably consistent among all four; the thermal fission cross section in ENDF/B-IV is not the value recommended by the Normalization and Standards Subcommittee of the Cross Section Evaluation Working Group (CSEWG-NSS) which recommends a cross section standard for ENDF; BNL-325 is odd man out In its evaluation of 0o, and the ENDF/B-IV data has an ao value which differs slightly from the other three. - 190 - Bigham's recent re-evaluation of experimental measurements of a at Chalk Rivf-r has produced a value of 527.4 + 3.1 barns which makes the CSEWG-NSS number look out of place. For interest sake, we have rerun the same SSET case, using each of these four sets of data. Both fission and absorption resonance integrals were kept constant so as to isolate the effect due to changing the 2200 m/s cross sections only. In Table E-2 is indicated the effective reactivity of the fresh fuel bundle in the calculated spectrum, as well as the predicted burnup for each set of data. Because of the sensitivity of SSET cycles to fissile conversion ratio, and the vast amounts of computation involved in determining a unique composition for each case, the last line indicates this ratio expressed as final (U233 + Pa233) per initial U233 atom. The last two columns where the two H values are comparable, but a is not, indicate that burnup predictions can be sensitive to more than just the ratios of cross sections - absolute values and competition for neutrons also play a role, although the conversion ratios differ here. This is also well illustrated by a comparison of columns two and four, where the ENDF/B-IV n value is 0.6% greater than that of the IAEA evaluation. However, the value of a here is 2% larger, resulting in a decrease in initial reactivity of about 6 ink and 2200 MW.d/t less burnup, with essentially the same fissile conversion ratio. Similarly, a comparison between columns one and three, both having similar conversion ratios and i")0 values, shows that the larger ot results in lower burnup as might be expected. So predicting the burnup of SSET cycles requires considerable confidence and consistency in the choice of thermal data. - 191 - In the epichermal energy range, the story Is analogous. Here the amount of self-shielding of resonances, which is the same as a change in the evaluated resonance integral, is of concern. In natural uranium cases it is adequate to use Westcott temperature dependent s-values, evaluated on the assumption that the materials are infinitely diluted, except for U238 and Th232« Since U233 has a large resonance integral and is present at more than twice the concentration of U235 in natural uranium, this assumption is suspect, so some calculations were done which indicated that for fuel bundles of interest, the U233 resonance is typically reduced to 78% of its infinite dilution values; for Pa233 the reduction is to 93% of infinite dilution. These are not precise results (Figure E.7), and it can be seen that predicted burnup is a sensitive function of the computed self-shielding of U233 because of the reduction in the fission cross section. Of course the sensitivity will vary with the type "f neutron spectrum, increasing with the hardness. The case shown, in common with all SSET cycles, has a reasonably thermal spectrum so there is not too strong a dependence on shielding of Pa233. Thus there are some doubts as to the adequacy of resonance treatment and more sophisticated methods are being added to our lattice codes to deal with this problem. A further factor of concern is the question of shielding between resonances of different nuclides (Figure E.8). The three materials U233, TJ234 and Pa233 all contain prominent resonances in the range 1 to 10 eV and have comparable macroscopic cross sections. In this figure is plotted the energy-dependent cross section of each of the isotopes in this energy range, weighted by a nuclide composition typical of each species when equilibrium concentrations are obtained. It can be seen that there is considerable scope for complex shielding interactions between the nuclides complicated by the fact that the relative weighting of U233 and Pa233 will vary during the initial transient to equilibrium. - 192 - Another of the physics problems unique to thorium-fuelled reactors is the problem of flux dependence. This problem is well known, and relates to the branching at Pa233 in the fuel cycle (Figure E.9). Intuitively it can be seen that as the absolute flux level Increases, there will be more captures in Pa233. This will result in an increased production of U234 and a decrease in U233, the prime fissionable nuclide. So, the reactivity of a fuel bundle decreases with increasing flux or power density, resulting in a lower attainable burnup at higher power densities. In a SSET cycle this can pose a severe restraint on reactor design (Figure E.10). Plotted here is reactivity as a function of attainable burnup at two constant levels of flux, corresponding roughly to power levels of 12.5 and 7.7 kW/cm. It can be seen that a 60% increase in power level produces a 47% reduction in attainable burnup simply due to decreased production of U233, and increased captures in Pa233. Both curves are normalized to the same excess reactivity allowance of 3% (30 mk). Since CANDU reactors may also be fuelled on-power, there is nothing to stop a designer from investigating the potential of various schemes whereby the fuel is somehow shuffled in a channel, either by removal or push-through methods. The time variation of U233 and Pa233 associated with this type of treatment make the physics analysis quite interesting. For example, certain fuelling strategies may Introduce transients in U233 concentrations which are similar to the familiar Xel35 transients. In the static case, the properties of a fuel bundle become time as well as irradiation dependent (Figure E-ll). For example, the reactivity of the SSET case calculated earlier for two flux levels was well behaved if the flux level was kept constant throughout the irradiation. The dashed lines show how the reactivity will vary, If at some burnup the flux level decreases from 5 to 3 x 10 n.cm .s or increases from 3 to 5 x 10 n.cm .s .The predicted burnup for each case is also indicated. Note that the discontinuity between the dashed and full curves is caused by the flux dependence of fission products at the two flux levels, and shows that the U233-Pa233 effect is of greater magnitude. - 193 - H.4 REACTOR PHYSICS ASPECTS Reactivity Coefficients The fuel temperature coefficient of reactivity is negative and decreases with irradiation (Table E-3). This behaviour Is contrary to the situation in the natural uranium system where the coefficient increases with irradiation. The behaviour of the natural uranium system is due primarily to the buildup of Pu239 and the increased number of fissions in the 0.3 eV resonance as the temperature increases. With the plutonium topped thorium system the Pu239 is destroyed and consequently the effect of the 0.3 eV resonance decreases with the irradiation. The U235 topped system which contains no plutonium and the dominant effect is the increased capture in the thorium resonances due to Doppler broadening. The coolant temperature coefficients of reactivity are positive and show a complicated irradiation dependence. For fresh fuel the coefficient is large for the thorium cycle but is less for the equilibrium fuel. The coolant void effect is positive but smaller than that for the natural uranium system. The major effect here is the decrease in the resonance integral of the fertile material on voiding the coolant. Because tl.e effective resonance integral in thorium is less than in U238 the coolant void effect is correspondingly smaller. The overall power coefficient of reactivity is illustrated in figure E.12. Xenon and Similar Effects The xenon problem in uranium reactors is well known. In a thorium reactor the situation ie a little different since the direct yields of Xel35 and 1135 from the fission of U233 are different from the natural uranium values. Typical values of yields and resulting effects are shown in Table E-4. For thorium reactors both the equilibrium and peak xenon reactivity worths are reduced. - 194 - It has been noted that the preseru.e of Pa232 which decays to U233 presents problems unique to the thorium system. Very slow spatial instabilities, analogous to the xenon effect, may develop and on shutdown the decay of Pa233 to U233 causes a slow reactivity increase which can rise to approximately 100 ink. The shutdown system of the reactor must be able to cc^e with this situation. The spatial distribution of U233 after shutdown will not be the same as before shutdown. The Pa233 and U233 distribution prior to shutdown tends to flatten the power distribution, the U233 generated after shutdown will tend to cause power peaking at the centre of the core which may require additional zonal control capability. Kinetic Parameters The effective delayed neutron fraction increases with irradiation for thorium fuel while it decreases for uranium fuel. The reason is the buildup of Pu isotopes in the uranium system, which in the thorium system are either pot present or depleting. Prompt neutron lifetimes are generally larger in the uranium reactor than in the thorium reactor due to the higher absorption in the latter. The apparently worse control situation in the thorium system is compensated by the differing reactivity coefficients and xenon parameters so that overall the control aspects are similar or even better than for the natural uranium reactor. - 195 - Fuel Management Fuel management is more complicated in the thorium reactor due to the dependence of burnup properties on the absolute flux level. Realistic calculations must take into account the detailed flux history of each fuel element. The presence of Pa233 allows a rather different fuelling stre.egy for thorium systems to be considered. In this scheme the decay properties of Pa233 are utilized tc allow operation at lower flux levels without decreasing the instantaneous power density. This is done by periodically removing fuel from the reactor and allowing U233 to buildup from the decay of Pa 233 and re-inserting the now more reactive fuel. Initial studies indicate that this strategy can result in increasing burnup and decreasing system fuel inventories as the design channel power increases, which is in the right direction to improve the economics. TABLE E-l Isotoplc Compositions and Fissile Inventories of Several CANDU Thorium Cycles* U 3ecyc11nc and Pu Top.plng U Recyq ing and U235 Tonpln (93< enr U) Flsslle Topping, 0 1 5 0 5 g/kg burnup, MW.d/kg 10 8.7 37 2 10 19.5 37.5 Burnup Category Burnup Category Self-sufficient 1ntermedI ate High Self-Sufficient nterrodiate High Eq Feed Discharge Feed Dlsj^arge Feed Discharge Feed Discharge Feed Discharge Feed Discharge Isotopes, g/kg HE Th232 974.8 964.0 973.0 955.0 967.4 934.0 974.8 964.0 970.5 952.0 962.1 928.0 U232 0.008 0.008 0.014 0.014 0.027 0.027 0.008 0.008 0.014 0.014 0.025 0.025 U233 15.1 15.1 15.2 15.2 15.0 15.0 15.1 15.1 15.1 15.1 14.8 14.8 U234 5.7 5.7 5.8 5.8 5.7 5.7 5.8 5.8 5.8 5.8 5.7 5.7 U235 1.3 1.3 1.3 1.3 1.4 1.4 2.7 1.7 2.7 !.7 6.9 1.9 U236 3.0 3.0 3.1 3.1 3.0 3.0 4.8 4.8 4.8 4.8 7.5 7.5 U236 0.06 0.06 0.06 0.06 0.05 0.05 1 .1 1.0 I.I 1.0 3.0 2.6 Pu239 - - 0.9 - . 4.5 ------Pu240 - - 0.5 - 2.3 ------Pu24l - - 0.1 - 0.5 ------Pu242 - - 0.2 — - ** •Isotoplc compositions and fissile Inventories of some CANDU thorium cycles. Natural uranium CANDU fuel contains 2.7 g fissile Pu/kg He at discharge. Notes (I) Discharge concentrations are g/kg of Initial HE In feed fuel. (2) Pu Is discharged at the end of each Irradiation and Is not recycled. The quantities are small and the l~u240 concentration too high for It to be of valuo. - 197 - TABLE E-2 BNL-3251-1 ENDF/B-IV2^ CSEWG-NSS3) IA" (1973) (1974) (1974) (1975) 47.7+2 45.9 46.2+1.3 45.3+.9 31.1+1 .3 525.11 533.7+2.7 529.9+1.4 ,x =OC/Of .0899+_.004 .0874 ,0866+_.0024 .O855+_.OO2 o o o !] =v/(l+a ) 2.287+.007 2.2972 2.2841+.0088 2.283+.006 o o 140+6 134. L.5 764+13 761. .183 .176 — — u.5 burnup 11,900 10,800 13,900 13,050 (MW.d/t) thermal (ka~l) .2156 .2109 .2200 .2168 (Q233+Pa233)/U233 1.003 1.014 1.008 1.012 initial 1) Neutron Cross Sections, Volume 1, S.F. Mughabghab and D.I. Garber, BNL-325, Ed.3, V.I, TID-4500. 2) Material 1260 on ENDF/B-IV tapes; evaluated by N.M. Steen. 3) Thermal Data for Fissile Nuclei in ENDF/B-IV, J.R. Stehn, Trans. ANS, 18,351(1974). 4) The Third IAEA Evaluation of the 2200 m/s and 20°C Maxwellian Neutron Data for U233, U235, Pu239 and Pu241, H.D. Lemrael, in N.B.S. Special Publication 425, R.A. Schrack and CD. Bowman, eds, Oct. (1975). - 198 - TABLE E-3 Reactor Physics parameters for ThO, and nat. UCL fuelled PHWs Th PHW 1200 U-nat PHW 1200 fresh equilibrium fresh equilibrium fuel fuel fuel fuel Fvsl temperature -0.70 -0.82 -1.49 -0.56 5 coefficient %2.a t x, {-^.lO ] T =753°C T =753"C T =890°C T =890°C dT r C r r r r Coolant temperature 3.78 2.71 1.35 3.64 dT coefficient ~ at Tc [4^.10 J Tc=293°C Tc=293°C Tc=282°C Tc=282"C Coolant void effect pno cool - pcool at T .10 6.96 7.28 18.3 10.7 Effective 2.56.10"3 3.13.10~3 7.04.10 3 4.80.10 3 delayed neutron fraction 6 + (7,n) 2.69.10 3.43.10 7.30.10 5.00.10 Effective prompt neutron lifetime &* [ms] 0.43 0.52 0.70 0.63 - 199 - TABLE E-4 XENON PARAMETERS FOR TOPICAL REACTORS evaluation is - at 1.2 n/kb irradiation - for PHW 1200 reactors U233/Th NATURAL U Xel35 direct yield 1.086% 0.562% 1135 cumul. yield 5.411% 6.212% from fission Equilibrium Xel35 worth (ink) -24.6 -29.0 Peak Xel35 worth (ink) after complete shutdown -88.9 -154.0 PROCESSING LOSSES % 40i- TOPPING ENRICHMENT GRAMS FISSILE PER Kg a. z K O CD o Pu TOPPED THORIUM CYCLES o U235 TOPPED THORIUM CYCLES H _L 0 10 20 30 40 50 60 ANNUAL URANIUM REQUIREMENTS AT I 00% LOAD FACTOR MgU PER GWe " 1 1 1 i 0 O.I O.E 0.3 0.4 FIGURE E.I- ANNUAL URANIUM REQUIREMENTS RELATIVE TO NATURAL URANIUM CYCLE - 201 - i i f r 40,000 2.5% NEUTRON WASTAGE 30,000 00 20.000 3.5% NEUTRON WASTAGE 10,000 I I I 0 10 20 30 40 50 U CONSUMPTION (g/kWa ELECTRICAL) FIGURE E.2= BURNUP VERSUS U CONSUMPTION FOR (Th + U233+ Pu239) FUEL CYCLE - 202 - — 233U-Th BUNDLE NATURAL URANIUM BUNDLE I J_ I • I • 0 1000 3000 5000 7000 9000 BURNUP MWd/t FIGURE E.3: NORMALIZED! ,Z AND REACTIVITY VS BURNUP f a BARNS/FISSION — (\3 no —o— Ol o Ol o Ol 56 o o o o o o m rn — CD m 5 CO CO CO > z o > £ O -D 2 m §3 m m 70 5 > o m co oo PERCENT ERROR - £03 - "FRESH FUEL" WITH 1.090 - EQUILIBRIUM POISONS .070 - 1.050 - O < o LLJ 48 EXPLICIT FISSION PRODUCTS .030- PSEUDO FISSION PRODUCT 3 PSEUDO FISSION 3 SATURATING FISSION PRODUCTS 1.0 10 0 1000 3000 5000 7000 '9000 ' I 1000 8500 10 100 BURNUP (MWd/t) FIGURE E.5: BURNUP PREDICTION USING TWO SETS OF FISSION PRODUCTS 233 1 u V ENDF/B-IV (MATERIAL 1260) 3.00 3.00 2.75 ENDF/B- I I (MATERIAL 1041) 2.75 2.50 2.50 2.25 2.25 2.00 2.00 1.75 1.75 1.50 .00 0 IOC 1000 |04 I05 I06 I 0' ENERGY (eV) FIGURE E.6= U233 n =va /a VERSUS ENERGY r a - 20G - I 1000 U233 SHIELDING 10500 Pa233 SHIELDING + REFERENCE REFERENCE 10000 00 9500 9000 I I I 70 75 80 85 90 95 100 SHIELDED RESONANCE INTEGRAL/INFINITE DILUTION (%) FIGURE E.7= SENSITIVITY OF PREDICTED BURNUP TO SELF- SHIELDING OF U233 AND Pa233 - 207 - 10° 2 3 4 5 I01 NEUTRON LNERGY (eV) FIGURE E.8- MACROSCOPIC CROSS SECTION OF U233, u"4 AND Pa233 AS A FUNCTION OF ENERGY IN A BUNDLE WITH AN EQUILIBRIUM CONCENTRATION OF THESE ISOTOPES. - 208 - 233 234 235 u u U 6.7 HOURS 27 DAYS THERMALLY THERMALLY FISSILE FISSILE 233 Pa Pa 234 KEY 22 MINUTES -DECAY 233 Th 232 Th (n, FIGURE E.9* TRANSMUTATION CHAIN LEADING TO THE FORMATION OF U233 FROM NEUTRON CAPTURES IN Th232 ,.080 .060 < >- =- 1.040 O o < u 13 2 1 .020 . <£=5xlO n-cm" • s" (12.5 kW/cm) 0,1 50 2 1,800 .000 .1 0 4000 8000 12000 16000 20000 24000 BURNUP (MWd/t) FIGURE E. 10= VARIATION OF REACTIVITY WITH BURNUP AT TWO LEVELS OF FLUX .080 CONSTANT FLUX .060 - STEP CHANGE IN FLUX >- £ 1.040- o < 1.020 - .000 8000 I2C00 20000 24000 BURNUP (MWd/t) FIGURE E.I 1= VARIATION OF REACTIVITY WITH BURNUP USING DIFFERENT FLUX LEVELS AND HISTORIES - 211 - 0.06 LU CD 0 05 z: 0 04 VE R CH / j> o 0. 03 Q_ EQUIL U0 0.02 % \- Ld 0.01 O 0 COEF F oc -0 .0 1 u EQUIL ThO, o Q_ -0. D2 100 110 120 30 REACTOR POWER (%) FIGURE E. 12= COMPARISON OF POWER COEFFICIENTS FOR ThO2 AND NAT. U02 FUELLED PHW's - 212 - F. RKSOORCE UTILIZATION AND ECONOMICS WITH THE THORIUM SYSTEM F.I THORIUM CYCLES IN CANDU REACTORS Two types of thorium cycles in CANDU reactors are currently of most interest to Atomic Energy of Canada Limited (AECL). These are: i) cycles with an external feed of thorium and fissile plutonium to a CANDU reactor and recycle of the uranium produced in the reactor. The fissile plutonium is produced in CANDU reaccors of current design fuelled with natural uranium. Thus two reactors are required: one fuelled with thorium and one with natural uranium. ii) cycles with an external feed of thorium and highly enriched U235 to a CANDU reactor and recycle of the uranium from the reactor. In this case only one reactor type is required. AECL has carried out a general survey of the economic and performance characteristics of these cycles. Some typical results from this survey will now be briefly discussed. These particular results apply to cycles using fissile plutonium, supplied from a natural uranium fuelled CANDU-PHW, plus thorium in a CANDU-PHW with uranium recycle. All reactor units are a 1200 MW(e) size (i.e. electrical power) with a fairly standard lattice design, and use 37-element oxide fuel bundles in 4" (nominal inner diameter) pressure tubes. The design power for the maximum rated channel is 6.5 hW (thermal). The important parameters for the natural uranium fuelled reactor are a fuel burnup of 7.5 MW.d/kgU with spent fuel containing 2.7 g fissile plutonium per kg. The more important economic parameters used are listed in Table F-l. Note that in the general survey studies no allowance was made for fissile material losses in fuel fabrication and reprocessing. - 213 - Figure F.I shows equilibrium burnup per pass and conversion ratio for the thorium fuelled CANDU-PHW as a function of fissile Pu content of the feed fuel (Th + recycled U + Pu). Also shown are system total unit energy cost and system equilibrium uranium requirements. The conversion ratio is defined as the ratio of total fissile atom production to total fissile atom destruction. As expected the burnup increases with increasing fissile Pu content in the feed fuel. As the burnup increases the conversion ratio decreases and the system equilibrium uranium requirements increase. The total unit energy cost has an optimum, but the curve is quite shallow. Figure F.2 gives exactly the same information as Figure F.I but in a slightly different form. The characteristics have been cross-plotted against system equilibrium uranium requirements to stress that aspect of the results to be dealt with in more detail. For zero system equilibrium uranium requirements the burnup is finite and the energy cost is reasonable. This we call a self- sufficient equilibrium thorium (SSET) cycle since at equilibrium no fissile Pu topping is required for the feed fuel (only recycled uranium). Hence at equilibrium no natural uranium reactors are needed in the system, which can operate with only an external feed of thorium. Reasons for Interest in Self-Sufficient Equilibrium Thorium Cycles i) They provide insurance against uranium shortages and high uranium prices. - 214 - Figure F.3 shows the estimated mined uranium requirements for the cycles of Figures F.I and F.2. The total uranium require- ments over a nominal reactor lifetime (23 full power years) are lowest for the SSET cycle. Furthermore, since the inventories occur only once for a given installed capacity, the advantage increases as further reactor generations are considered. The inventories include both out-reactor and in-reactor components. The out-reactor component is based on an assumed delay time of 1.5 years for fuel cooling, reprocessing, fabrication and hold- up. In fact for the SSET cycle, a given installed power capacity can be maintained indefinitely with no increase in cumulative uranium requirements. A point is also shown on Figure F.3 for an SSET cycle based on highly enriched U235 for the initial fissile requirements rather than fissile Pu. A tails assay of 0.2% for the enrichment plant was used in deriving this estimate. Figure F.4 shows estimated energy costs for the same cycles at various yellowcake prices. The solid parts of the curves are for conditions giving energy costs equal to or less than those from a natural uranium fuelled reactor under the same uranium price assumption. Low energy costs can be maintained even at very high uranium prices. (A price of 18.4 S/kg for thorium is assumed throughout. The price for thorium is not expected to increase substantially since it contains no fissile material. Recycling the thorium provides an upper limit on cost.) ii) The characteristics of SSET cycles are such as to make their design, operation and economic behaviour most sensitive to input data, methods of estimation and efficient design. The discipline imposed by this sensitivity tends to improve the state of knowledge of other thorium cycles and their performance. - 215 - Key Problem Areas The key problem areas in assessing SSET cycles arise from the sensitivity of their performance to economic parameters and achievable burnup. Figure F.5 shows the sensitivity of estimated energy costs to reprocessing cost and fuel fabrication penalty associated with active fabrication. AECL plans to accumulate experience in these areas and this will eventually provide a better basis for estimates. Figure F.6 shows the sensitivity of estimated energy costs to burnup. The SSET cycles have the greatest sensitivity. The survey calculations described to this point have a number of shortcomings with respect to burnup estimates associated with physics data and methods, allowance for fissile material losses in fabrication and reprocessing, and reactor design parameters. The next section discusses some of these problems in more detail. F.2 BURNUP ESTIMATES F.2.1 General The main difficulties in making accurate burnup estimates for SSET cycles arise from uncertainties in allowances for fissile material losses due to fuel fabrication and reprocessing, and inaccuracies in reactivity estimates. Fissile material losses are equivalent to a change in conversion ratio. The effective change in conversion ratio, given by: (CR)eff = -A f - 216 - where A is dependent on the fissile content of the spent fuel and is 15 MW.d/kg for the cycles of the previous section, x is the fraction of fissile material lost per cycle, and B is the burnup per cycle in MW.d/kg. Figure F.7 is a plot of conversion ratio as a function of burnup for the cycles of Figure 3. Also shown are the effective conversion ratios corresponding to both 1% and 2% losses. The burnup achievable for SSET cycles, for which the effective conversion ratio must equal unity, is seen to be extremely se.isitive to losses. In fact, for the case shown, no SSET cycle is possible witn losses of 2%. The crucial reactivity estimates are those associated with parasitic neutron absorption (in structural materials and irradiation products), neutron leakage, fast fission and ri-values for fissile materials. Reactivity changes associated with the ratio of fertile to fissile material neutron absorptions are of secondary importance from the point of view of feasibility, but would be important for design purposes. Reactivity changes of the first type are also equivalent to changes in conversion ratio at constant burnup, with the equivalence given by: -3 (CR) = n k x 10 where n is the number of neutrons produced per neutron absorption in fissile material, and k is the reactivity change in ink. - 217 - Figure F.8 shows the conversion ratio as a function of burnup and the effects of changes of +10 ink and -10 ink in the reactivity. Note the sensitivity of burnup at constant conversion ratio to reactivity In this range - approximately 0.5 MW.d/kg of heavy element (HE) per mk. Inaccuracies and uncertainties in reactivity will be discussed under the headings of "physics" and "design". F.2.2 Physics The following method has been used for estimating equilibrium characteristics of SSET cycles. The reactivity limited burnup, using an idealized fuel management scheme, for an initial fuel material (such as ThO2 plus Pu) in the assumed reactor lattice is calculated using a lattice code- The isotopic and total uranium content of the fuel at the end of irradiation is determined and used to specify a new fresh fuel composition allowing for reprocessing and fabrication losses. This process is repeated until the isotopic composition of the uranium ceases to vary from iteration to iteration. During the first few iterations extra fissile material (e.g. Pu) must be added to each step but eventually is no longer required to obtain the equilibrium burnup. The information on required additions of extra fissile material in the early stages is used to estimate "inventory" terms. The method is in effect a simulation of the procedure which would be used in an actual reactor to establish equilibrium. Although the first few steps are somewhat arbitrary a unique equilibrium result is obtained. - 218 - A set of nominal isotopic compositions for an equilibrium SSET cycle is given below. This set can be used as an initial composition to speed convergence to equilibrium. Total U/Th : 2.6% U Composition 232 : Trace 233 : 61% 234 • 23% 235 6% 236 9% 238 Trace The main uncertainties in feasibility arising from this procedure are associated with nuclear data and fuel management. Those arising from the lattice calculational methods are thought to be relatively minor. Further experimental work is required for final verification. A great deal more experimental and theoretical work would be required to obtain accurate detailed information for actual design and operation. Nuclear Data Uncertainty in the r]-val\te of U233 is estimated to contribute an uncertainty in reactivity of some 3 ink. Considerable effort has been put into inter-comparing various approximations for non-saturating fission products and experimental data. This has led to a change in our best estimates from these used in the survey studies. In the survey studies three saturating and three pseudo fission products were used, whereas in the detailed studies to be presented 48 detailed fission product chains plus one - 219 - non-saturating fission product were used in addition to the saturating fission products. The uncertainty Introduced from this source Is estimated to be about 0.3 B mk (where B is the burnup in MW.d/kg HE). Uncertainties in other nuclear data are less important although further assessment of mutual resonance shielding between U233 and Pa233 would be desirable. Fuel Management The calculation of spatial flux distributions in Th fuelled reactors Is complicated by the flux dependence of U233 holdup in Pa233. This makes the study of fuel management schemes difficult and time- consuming. Preliminary studies of three different axial fuelling schemes suggest a reactivity uncertainty of up to 5 mk associated with fuel management estimates for SSET cycles, mainly due to variations In the neutron leakage associated with different fuel management schemes. This is the area which would probably require the most effort in upgrading information to a level required for design and operation. Figure F.9 shows the effect of reducing the lattice pitch from 28.6 to 22.9 cm. For a given fissile topping, the burnup is reduced leading to an increase in uranium consumption and fuelling costs. Figure F.10 shows the effect of average fuel specific power (or channel power) on burnup and uranium consumption. The effect of increasing the average fuel specific power is to decrease burnup and increase uranium consumption for a given amount of fissile topping. - 220 - This effect is caused primarily by neutron absorption into Pa233. The Pa233 level increases with the flux level and hence with specific power at a given enrichment. Although the burnup increases and uranium consumption decreases as the fuel specific power is decreased, the fuelling cost may not decrease. This is because the Inventory of fissile material increases with decreasing specific power. The effect of specific power on fuelling costs is illustrated in Figure F.ll (l^Og cost is again taken to be 50 $/kgU). Several, more subtle design assumptions were made in the early studies, based on experience with natural uranium fuelled CANDU reactors. These included fixing heavy water isotopic composition, control margins (including '•aactivity effects of structural material in control, safety and dete •••or devices) and pressure tube material. The conditions under which SSET cycles would be attractive are so different from the conditions pertaining to our present CANDU design that some of these assumptions should be re-examined. D?0 Isotopic Composition The normal design isotopic composition of the moderator (as distinct from the coolant) in a CANDU reactor is 99.8 atom %. Consideration has been given in the past to increased purity but, while this appears feasible, the incentive has not been great enough to make the change. For an SSET cycle the incentive is much greater. Therefore it seems reasonable to assume moderator purities as high as 99.95%. This change corresponds to a reactivity gain of approximately 5 mk. - 221 - New Materials Units P3 and P4 at Pickering use Zr-2^wt%Nb pressure tubes while Pi and P2 use Zircaloy-2. The reactivity difference is about 5 mk in favour of P3, P4. The previous survey studies assumed pressure tubes designed to the P3, P4 criteria. However, development work is continuing on stronger pressure tube materials and it seems reasonable to expect up to a further 5 mk gain with respect to the "standard" design based on P3, P4 criteria. Control Margins In the survey studies an allowance of 6 mk was made to provide for control margin and structural material associated with control, safety and detector devices. This is nominally the same allowance as for a natural uranium fuelled reactor. One might argue that a slightly lower value would be suitable on the basis of a more heavily absorbing core, and fuel properties which are almost constant over the irradiation lifetime. On the other hand the presence of Pa233 leads to some new requirements which could increase the required allowance. Xenon OverRide It has been conventional to allow about 1/2 hour decision and action time in CANDU-PHWs for xenon override after a shutdown. This amounts to some 18 mk in reactivity in the form of absorber rods which can be removed after a shutdown. The same rods are used to flatten the power distribution under normal operation and their reactivity effect is partly due to neutron absorption in the rods ( The characteristics of a typical SSET cycle are such that only 40% as large a reactivity allowance is required to provide the same override. It is doubtful that this small amount ('W mk) will be enough to provide the desired power flattening. However, there is considerable scope for new methods of achieving radial power flattening in SSET cycles with minimum neutron wastage. For example, the fissile content of the feed fuel for a central reactor core region could be made different from that for an outer region. Thus the full saving from reduced xenon override requirements should be realized- A further saving could be achieved if it were feasible to use thorium as the removable absorber for xenon over- ride. F.3 RESULTS A detailed study of SSET cycles has been done and some typical results will now be discussed. One of the parameters used in this discussion will be "reactivity allowance". The "reactivity allowance" includes the reactivity requirements for neutron leakage, control margins, and xenon override. The results have been derived in this way to permit quick assessment of various options and uncertainties. Information will be given for two designs: a "standard design" and a "potential design". The "standard design" has pressure tubes designed to th.* P3, P4 criteria and a moderator D2O purity of 99.8%. The "potential design" has pressure tubes with a 5 mk reactivity saving relative to tubes designed to the P3, P4 criteria and a moderator D2O purity of 99.95%. The reactor is a straightforward design with 6 m long, 4" (nominal) diameter channels on a 28.6 cm square lattice pitch, and a maximum channel power of 6 MW(t). The oxide fuel is in the form of 37 element bundles. - 223 - The range of reactivity allowance believed to be applicable to 1200 MW(e) units operating on an SSET cycle is 25-35 ink, based on the following approximate breakdown. TABLE F-2 Reactivity Allowance (mk) Pessimistic Reference Optimistic Leakage + Xe override 28.5 25 21.5 Control Margin 6.5 5.0 3.5 Total 35 30 25 Figure F.12 shows achievable burnup as a function of reprocessing and fabrication losses for both designs and various reactivity allowances. The burnup is very sensitive to losses, giving a big incentive to keep the losses below 1% if possible. Above about 2%, SSET cycles are not feasible. Figure F.13 shows achievable burnup as a function of reactivity allowance. Also shown is the point representing a natural uranium fuelled reactor. There is a big incentive to minimize the reactivity allowance. For values much above 40 mk, SSET cycles are no longer feasible at this maximum channel power. Figure F.14 g'ves an economic comparison among natural uranium cycles, the SSET cycle and other thorium cycles. For the reference economic parameters, the yellowcake price must reach 200 $/kgU for the SSET cycle to compete with the natural uranium one, and evan higher yellowcake prices are required to give energy costs as low as some of the other thorium cycles. (Note that the other thorium cycles start to compete with natural urani'im cycles at a uranium price of about 100 $/kgU.) - 224 - For the SSET cycle shown in Figure F.14 the estimated cumulative mined uranium requirements for indefinite operation of a given capacity are about 2 MgU/MW(e). Figure F.15 shows the economic behaviour for similar systems to those in Figure F.14 except that U235 topping Is used. For the SSET cycle with U235 topping, the cumulative mined uranium requirements for operation of a given capacity are about 1 Mg/MW(e), assuming 0.2% tails assay in the enrichment plant. F.4 CONCLUSIONS (a) SSET cycles in 1200 MW(e) CANDU reactor units, achieving burnups of 10 MW.d/kg HE per cycle, are feasible provided fabrication and reprocessing losses of fissile material can be kept below 1% and careful attention is paid to neutron economy. (b) Achievable cumulative uranii."«i requirements for indefinite operation of such cycles are: i) 2 Mg/MW(e) if fissile plutonlum produced in a natural uranium CANDU is used for initial fissile material requirements. ii) 1 Mg/MW(e) of highly enriched U235 is used for initial fissile material requirements. (c) Such cycles are not expected to compete economically with natural uranium cycles until yellowcake prices reach 200 $/kgU (1978 dollars). (d) Such cycles are not expected to compete economically with other thorium cycles in CANDU reactors until prices considerably higher than 200 $/kgU (1978 dollars) are reached. - 225 - (e) Other thorium cycles are expected to compete with natural uranium cycles in CANDU reactors at about $100/kgU (1978 dollars). (f) The major uncertainties in assessing such cycles arise from estimating losses and costs Incurred in fuel fabrication and reprocessing. Resource Utilization - Growing Nuclear Power Systems The results of systems in which equilibrium has been established are of interest but can be misleading if the nuclear power capacity grows rapidly. To gain some understanding of the effect of growth rate, system studies have been made. Resource consumption in growing nuclear power systems containing natural uranium and thorium cycle reactors was determined taking into account fuel cycle process delays, fuel inventory requirements and extra fissile material requirements to reach equilibrium. For the reference cases all extra fissile material required for the thorium cycle was assumed to be produced (as plutonium) by natural uranium PHW reactors. However, the effect of having enrichment available was also considered. Two growth rate curves for total installed nuclear capacity will be discussed. These are not meant to be predictions in any sense, but are used to illustrate effects of interest. The first growth curve goes from an installed capacity of 130,000 MW(e) in 2000 to 900,000 MW(e) in 2060 and does not grow after that. It is shown as curve I in Figure F.16. The second growth curve is similar except that it goes from 130,000 MW(e) in 2000 to 900,000 MW(e) in 2030 and does not grow after that. It is shown as curve II In Figure F.16 (nuclear capacity factors are assumed to be 65% after 1995). - 226 - Natural uranium PHW reactors would supply all of the nuclear power demand until the assumed commercial introduction of the thorium cycle (or fast breeders) in 1995. Conversion of existing reactors was not permitted, so that introduction of thorium-fuelled PHW reactors was limited to expansion and replacement. The following systems were considered: 1. Natural uranium PHW reactors only 2. Natural uranium + Th (standard) PHW reactors 3. Natural uranium + Th (standard) PHW reactors with Uranium-235 enrichment available 4. Natural uranium + Th (self-sufficient) PHW reactors 5. Natural uranium + Th (self-sufficient) PHW reactors with Uranium-235 enrichment available 6. Natural uranium + liquid metal fast breeder reactors (LMFBR) The reactor characteristics and process delay times are shown in Tables F-3 and F-4. Material losses are assumed to be about 1%. Figure F.17 compares the total installed capacity of natural uranium reactors for systems 1 to. 6 for growth curve I. If no thorium reactors were built (case #1), all reactors would be natural uranium reactors as shown by the top curve (#1). However, if thorium reactors are built from 1995 onwards, the system contains only as many natural uranium reactors as are required to provide sufficient fissile material for the system. If uranium enrichment facilities are available (curves #3 and #5), natural uranium reactors are not required after 1995 and they are replaced by thorium reactors at their end of life. Note that when uranium enrichment is not available, the required natural uranium capacity is greater in the growing phase than at equilibrium, i.e. there is an overshoot in fissile Pu production capability. This is because of the extra fissile material required before equilibrium U233 levels are established. - 227 - Figure F.17 shows the same general features for growth curve II. The overshoot in natural uranium (plutonium producing) capacity is more pronounced since the growth rate is more rapid. Figures F.19 and F.20 show the uranium consumption for growth curves I and II, respectively. In the long term, the greatest savings in uranium are achieved with breeder reactors (whether they be thorium- fuelled PHWs or LMFBRs) but during periods of high growth this advantage is much less evident. However, in all cases the use of thorium cycles (or the LMFBR) results in considerable savings of uranium compared with the natural uranium cycle. One interesting facet of these curves is that more uranium is required in the initial period of introduction of thorium-fuelled reactors if uranium enrichment is available than if fissile material is obtained from natural uranium PHWs. This is because essentially all new reactors after 1995 are thorium fuelled if U235 enrichment is available and these require a large investment in fissile material before equilibrium U233 levels are reached. This is essentially an investment for the future since it leads eventually to a lower cumulative uranium requirement than for the natural uranium + thorium-fuelled systems. The curves also indicate that the uranium consumption in systems containing LMFBRs is of the same order as in self-sufficient, thorium-fuelled PHWs. Many other quantities of interest such as thorium requirements, fuel fabrication and reprocessing capacity were obtained in the system studied. Figure F.21 shows the inventory of thorium required for the faster growth rate (curve II) with no recycle, 20 year storage to allow Th228 decay, and 1 year storage. - 228 - The cumulative cash flow (in 1975$) for the faster growth rate (curve II) is shown in Figure F.22. This figure illustrates that the main capital requirements are for reactors. Capital requirements for heavy water, fuel fabrication and fuel reprocessing are comparatively small. We conclude from these studies that thorium cycles in CANDU reactors offer the possibility of conserving uranium even in rapidly growing nuclear power systems. The optimum fuel cycle strategy depends on the growth rate and uranium cost. However, current CANDIJ-PHW reactors produce power with good reliability and appear tc be sufficiently flexible to accommodate thorium fuel cycles covering the whole range of conditions studied. - 229 - TABLE F-l ECONOMIC PARAMETERS Yellowcake price 75 $/kgU Reprocessing cost 80 $/kg HE Fabrication costs ThO2 50 $/kg HE uo2 46 $/kg HE Fabrication penalties U233 65 $/kg HE Plutonium 37 $/kg HE Interest rate 10% Capital charge rate 10.61% Fuel out reactor delay time 1.5 years y - 230 - TABLE F»3 REACTOR CHARACTERISTICS Reactor Power Plutonium Fissile Plutonium Burnup Thermal Type Density First Charge Content topping MW.d/kg HE Efficiency kW(t)/ g Fissile/ in first g Fissile/ kg HE kg HE charge kg HE kg Fissile/ MW(e) 1 Natural 29 0 0.8 (U235) 0 7.5 30.5 U PHW Th(std) 29 25.6 2.9 3.5 27 30.5 PHW Th(near 29 23.35 2.6 0 13 30.5 breeder) PHW LMFBR* 44.2 5i.7 2.7 -8.6 37.4 44 * Figures shown are average for core + blanket. Simple doubling time is 17.4 years at 80% capacity factor. - 231 - TABLE F-4 PROCESS DELAY TIMES Description Years Minehead to in-reactor 1.15 (no enrichment) Minehead to in-reactor 1.70 (with enrichment) Reactor out to reprocessing 1.0 out Fabrication in to in-reactor 0.5 Reactor out to in-reactor 1.5 for recycle fuel - 232 THORIUM CYCLE CHARACTERISTICS AS A FUNCTION OF Pu TOPPING o « 70 » 60 CO S 50 P CO CONVERSION 40 8 8 ——— RATIO I CD £ 30 ce g 10 - 2 - CO < CO o FISSILE Pu CONTENT OF EQ FEED FUEL (g/kgHE) (i.e. "Pu TOPPING") FIGURE F. I THORIUM CYCLE CHARACTERISTICS AS A FUNCTION OF EQUILIBRIUM U REQUIREMENTS 1 «f \A 12 60 Ul 12 2 o> 10 1.0 50 - s 10 A T 1 0 fST E CO c3s 8 - z .8 40 _ 8 o co O- 6 - DC .6 - 0. 30 - o 6 UJ > z - o .4 or 20 a. 4 o m UJ oLkJ _i 1- 2 .2 - 10 - 2 ISS I u. 0 0 0 0 0 2 4 6 8 10 12 14 SYSTEM EQUILIBRIUM U REQUIREMENTS (g/ekWa) FIGURE F.2 - 233 - TOTAL MINED U REQUIREMENTS AS A FUNCTION OF EQUILIBRIUM U REQUIREMENTS Ul z> a LU I - z 6 INVENTORY (IN & OUT-REACTOR) INVENTORY PLUS 23 FULL POWER YEARS O SYSTEM USING U235 INSTEAD OF Pu 0 10 20 30 40 50 60 70 EQUILIBRIUM U REQUIREMENTS (g/ekWa) FIGURE F.3 SENSITIVITY OF ENERGY COSTS FOR THORIUM CYCLES TO YELLOWCAKE PRICE 1000 7C0 U PRICE ($ /kgU) ,500 300 12 COMPETITIVE WITH NATURAL U REACTOR NOT COMPETITIVE WITH NATURAL U REACTOR 0 10 20 30 40 50 60 70 EQUILIBRIUM U REQUIREMENTS (g/ekWa) FIGURE F.4 - 234 - SENSITIVITY OF ENERGY COSTS FOR THORIUM CYCLES TO ECONOMIC PARAMETERS 16- 15 FABRICATION PENALTY = +35*/kgHE E 14 O 111 13 REFERENCE 12 CO A REPROCESSING COST = -50 /kgHE 0 10 20 30 40 50 60 70 EQUILIBRIUM U REQUIREMENTS (g/ekWa) FIGURE F.5 SENSITIVITY OF ENERGY COSTS FOR THORIUM CYCLES TO BURNUP 16 A BURNUP = -4 MWd/kgHE _. 15 E a i3 t- REFERENCE UJ > <- to I I 0 10 20 30 40 50 60 70 EQUILIBRIUM U REQUIREMENTS (g/ekWa) FIGURE F.6 - 235 - EFFECT OF CHANGES IN REACTIVITY VARIOUS LOSSES .00 J>v ^ 0% LOSSES "^^e^^ '%L0SSE S .96 - ^ ^^sTO- 2% LOSSES .92 - g .88 a: LU .84 o o 80 • i i i 10 20 30 40 50 BURNUP PER PASS (MWd/kgHE) FIGURE F.7 EFFECT OF CHANGES IN REACTIVITY CONVERSION R^TIO 00 REFERENCE + 1 0 mk o .96 , REFERENCE .92 Z o REFERENCE - 1 0 mk CO .88 .84 - o o .80 i i i 10 20 30 40 50 BURNUP PER PASS (MWd/kgHE) FIGURE F.8 - 236 - CANDU - PHW MAXIMUM CHANNEL POWER = 6.5 MW (t) AVERAGE FUEL SPECIFIC POWER = 9 kW/kg (HE) 70 LATTICE PITCH 28.6 cm 22.9 cm 60 50 X3 40 a. ID 30 Z OH z> GO 20 kg (HE) 0 0 I 0 20 40 60 80 100 URANIUM REQUIREMENTS (g/kW(e)a) FIGURE F.9= THE EFFECT OF LATTICE PITCH ON BURNUP AND URANIUM CONSUMPTION - 237 - CANDU - PHW LATTICE PITCH = 28.6 cm CHANNEL POWER MW (t) 3.5 5.0 6.5 8.5 AVERAGE FUEL SPECIFIC 16 22 29 38 70 POWER (kW/kg (HE)) 60 50 P 40 Q_ ID -z. 30 01 ID GO 20 10 0 I I I 0 20 40 60 80 100 URANIUM REQUIREMENTS (g/kW (e)a) HGURE F. 10: THE EFFECT OF AVERAGE FUEL SPECIFIC POWER ON BURNUP AND URANIUM CONSUMPTION - 238 - ill! 1 /o1 CANDU - PHW _ LATTICE PITCH = 28.6 cm / TOPPING 2.0 U 0 COST = 50$/kg (U) / g fiss 3 8 / kg (He) 1.8 / — / 09 O 1.6 — O CD 1.4 — ID - he ^ 1.9 LJ 1.2 LU 1.0 "^^^1 14.2 _ cr 9.3 1 0.8 — 0.6 1 1 1 1 1 I 0 15 20 25 30 35 40 AVERAGE FUEL SPECIFIC POWER (kW/kg (He)) FIGURE F.lh THE EFFECT OF FUEL SPECIFIC POWER ON FUELLING COSTS - 239 - SSET CYCLES BURNUP VS LOSSES MAX. CHANNEL POWER- 0.6 MW It) LATTICE PITCH: 2.86 mm STANDARD DESIGN POTENTIAL DESIGN ~~ "- —• ^ 30 mk REACTIVITY ALLOWANCE -o 2 z CD 0 5 10 REPROCESSING AND FABRICATION LOSSES (%) FIGURE F. I 2 SSET CYCLES BURNUP VS REACTIVITY ALLOWANCE MAX. CHANNEL POWER' 0.6 MW It) LATTICE PITCH: 286 mm STANDARD DESIGN POTENTIAL DESIGN 25 ( -O- NATURAL U FUELLED CASE) a 20 0% LOSSES 15 I* LOSSES ' 0% LOSSES "S . 3 10 - I % LOSSES 5 - 15 20 25 30 35 40 REACTIVITY ALLOWANCE (mM FIGURE F. I 3 DIFFERENCE IN UNIT ENERGY COST (m$ AWh) DIFFERENCE IN UNIT ENERGY COST (m$ /kWh) (THORIUM - NATURAL U) (THORIUM - NATURAL U) Fl T O en m cr to O Fl IBR I O , H 0 O O en ONO M CD Q FI G c POT I s 31 C CT S3S : c ro to C 0 ,0 3; -7 m 0 c e 0 en Fl / H 01 cn -JS O m O Fl cr > O . DESIG N •n m LJIREMEN T o F . 1 4 SEME N 01 REAC T I V 71 •n _\ THORIU M 0 01 0 1 0 01 0 (SW U _l K c ^* l\) cn O c>^ en 0 o 01 H 0 -< -c o |8/ l -1 0 \^ Ol CJ1 0 > O 13 m cn IN G -0 en O 0 c c > o TOTAL INSTALLED NUCLEAR CAPACITY, MW (e) O OJ o 01 ~— c70 CO m CO 71 0 CD ro 0 CO 0 CO 0 c m 0 roo ro CO o m > ro m X) O o z a o r m ro 0 7a o pi ro o o 00 o -0 o H ro o o o ro o o INSTALLED NUCLEAR CAPACITY (AS NATURAL URANIUM REACTORS) GW (e) 5 ° 35 > H r < o r m m > CD C7 33 x c: 33 ' < c m 33 m > o H O 33 CO INSTALLED NUCLEAR CAPACITY (NATURAL U REACTORS) GW (e) en 00 o o o o o o _ o o o o o o CD CD (T 73 m CO o 7o3 >~o CO o CO -1 1 m en CO -f r~ m o r~ > om 73 (G R CO o z -l I c O c 1— 73 1 c m 73 m > o H O 73 CO (TO 6.0 Tg IN 2100) 2.5r 1. ALL NATURAL URANIUM 2. NATURAL URANIUM & THORIUM (STANDARD) 2.0 3. NATURAL URANIUM & THORIUM (STANDARD) & ENRiCHED URANIUM LU 4. NATURAL URANIUM & THORIUM (SELF-SUFFICIENT) 5. NATURAL URANIUM & THORIUM (SELF-SUFFICIENT), & ENRICHED URANIUM 2 .5 6. NATURAL URANIUM & LMFBR < .0 0.5 O 0 960 1980 2020 2030 2040 2060 2080 2 100 YEAR FIGURE F. 19: CUMULATIVE URANIUM CONSUMPTION FOR SYSTEMS I TO 6 (GROWTH CURVE I) 960 1980 2000 2020 2040 2060 2080 2 100 YEAR FIGURE F.20= CUMULATIVE URANIUM CONSUMPTION FOR SYSTEMS I TO 6 (GROWTH CURVE II) 1. NO RECYCLE 2. 20 YEAR ACTIVE STORAGE 3. I YEAR ACTIVE STORAGE .2 Q LU Z 0.8 o X I- 0.4 0 970 1990 20 10 2030 2050 2070 2090 YEAR FIGURE F.2I-- CUMULATIVE THORIUM INVENTORY REQUIRED AS A FUNCTION OF RECYCLE TIME (GROWTH CURVE II) REPROCESSING f. FUEL FABRICATION D20 in r- cn CM X CO < (J LU CJ 960 980 2000 2020 2040 2060 2080 YEAR Figure F.22: Cumulative Cash Flow Required in 1975 $ (Growth Curve II - System 2). Growth Curve II - 248 - G. THE AECL ADVANCED FUEL CYCLE PROGRAM G.I INTRODUCTION There are four areas under Investigation in the Advanced Fuel Cycle Program. 1) Fuel development 2) Fuel reprocessing 3) Reactor physics 4) Assessment At this time the program is funded at the laboratory level. However, we are planning to expand the scale to enable pilot reprocessing and fuel fabrication plants to be designed and built. G.I FUEL DEVELOPMENT A small scale glovebox facility exists at Chalk River for the production of ex active fuel elements. The facility has been used to produce 540 (U-0.5 wt% Pu)02 Bruce type fuel elements and is presently being used to produce a quantity of (Pu, Th)O2 for use In the ZED-II reactor at Chalk River and the WR-1 research reactor at Whiteshell. The line will then be modifed to enable some (Th, U233)O2 fuel to be fabricated. Although U233 is only a active there will in recycle fuel be trace quantities of U232. The decay chain for this isotope includes species with very penetrating gamma emission. It is necessary, therefore, to design fabrication routes which can be followed - 249 - remotely and behind significant shielding. Conventional pellet fabrication methods are inherently dusty. This dust adversely affects the relatively complex process equipment which requires regular maintenance. In a remote environment involving active materials, reduction in maintenance is obviously important and so alternate pellet fabrication routes are being investigated. Three possible routes have been identified. a) Pellet impregnation A great attraction of this route is that powder and pressing operation are performed using inactive ThC^ outside the shielded facility. Whole stacks of ThO2 pellets can be impregnated with uranyl nitrate, producing radially impregnated pellets with the fissile material concentrated at the outside of the pellets. This, in turn, concentrates the generation of fission gases in the cooler region of the fuel where gas retention is greatest. b) Spherepac The spherepac route which eliminates all powder production steps has been applied to the production of (Th, U235)O2 fuel. A thorium-uranium sol was prepared from nitrate solutions, formed into accurately sized, fully dense microspheres, then sintered and vibration packed inside fuel cladding tubes. A smear density of 90.6% of theoretical density has been achieved. This is the equivalent of dished pellet fuel in fuel elements with small plena. - 250 - c) Extrusion In this third alternate fuel production process, (Th, U235)O2 fuels have been extruded from sol-gel derived clay. The extruded slugs had homogeneous raicrostructures and high density (9.7 g/cm3) after sintering. Irradiation experience with fuels produced using the above three production processes is limited and must be expanded to be able to properly assess these methods. The ability to fabricate very radioactive fuels safely and reliably is an essential component of the program. All process equipment must be operated and maintained behind thick biological shielding and tee technology to do this must be developed and demonstrated before the concept of large scale recycle fuel fabrication can be judged. The remote technology development program has two parallel components. One Is a generic program to develop and evaluate the technology of handling, controlling and viewing operations in inaccessible areas. Specific examples of this work include programmable robots, controlled ventilation systems and advanced television systems. The other program component is the demonstration of selected pieces of process equipment specially developed for reliable remote operation. Specific examples of this work include automatic pellet inspectors, automatic forming of stacks of pellets conforming to low tolerance stack length and pellet handling methods. G.3 FUEL REPROCESSING Canada reprocessed both U and Th fuels from research reactors in the early 1950's using both solvent-extraction and ion-exchange - 251 - processes. Further R&D continued until the early 1960's, then was suspended as it was considered unnecessary to the economics of the CANDU concept. A modest R&D program was restarted in the late 1960's as the natural uranium fuel cycle became established and advanced fuel cycles for the CANDU reactor were assessed to determine the various options for ensuring a long-terra supply of nuclear fuel for the CANDU system. The present reprocessing program is of a generic nature and is focussed on development and testing of alternative reprocessing flow sheets, including those with proliferation-resistant advantages. The program is aimed at acquiring a basic understanding of the fuel reprocessing technology required to implement a Th-U233 fuel cycle initiated with plutonium recovered from irradiated CANDU natural uranium fuel. Process flow sheet development is geared to the following situations. a) Reprocess irradiated uranium fuel b) Reprocess irradiated thorium fuel c) Recover fissile material from unirradiated scrap material The three basic approaches are a) Separation of all components, fissile, fertile and fission products, from each other b) Separation of fissile material from the rest of the irradiated c) Basic decontamination where the fissile and fertile products stay together, co-processing. - 252 - The separations of interest are listed in Table G-l. A 30% TBP process is specified for all four fuel types but the amine process could be used for the two uranium fuels. The amine process was shown to be feasible on the laboratory scale and is being tested at the pilot plant scale in a joint Canadian- Italian reprocessing experiment. A flowsheet for processing (Th-Pu-U) fuel is under development. Work has proceeded through the following phases - non-active development - computer simulation - active flowsheet verification Work on the last phase is currently in progress in a small hot cell facility using irradiated (U235, ThO2) fuel. Initial results confirm the viability of the flowsheet although tests with plutonium are still to be done. Co-processing of fuel materials without the complete separation of fissile material in pure form may offer some advantages over conventional processing. It is viewed as a technical barrier to proliferation and may result in lower processing costs and improved product recoveries. Work on co- processing has begun and distribution ratios covering a broad ranga of conditions expected in such flowsheets have been measured. Correlation of the data and definition of reference flowsheet conditions will be carried out using computer codes, testing of the flowsheet is also planned. - 253 - G.4 REACTOR PHYSICS Introduction Reactor Physics is concerned with the nuclear properties of reactor cores, under both stationary and transient or accident conditions. Of foremost Importance among these nuclear characteristics is the practicability to shape, sustain and control a suitable field of neutrons within the core. It is this neutron field which, through the fission reaction in the fissile nuclei of U233, U235 or plutonium, determines the distribution and magnitude of heat generation. But the presence of the neutron population has other Important effects too. The materials present in the reactor core are irradiated. As a result of nuclear reactions, caused by this neutron irradiation, fertile nuclides such as Th232 or U238 transmute into the fissile species U233 and Pu239 and a number of materials become radioactive. Also, the fragment nuclei generated in the fission process go through a variety of further transmuta- tions and decays. All these irradiation effects shape the composition of the fuel discharged from the nuclear reactor and determine for instance whether or not recovery of the fissile species from the discharged fuel is worthwhile. The fundamental role of the neutron field in the reactor core and the reactor physicist's expertise with its behaviour require his Input in many areas: reactor design, fuel development, fuel manage- ment, reactor control, reactor safety, shielding, crltlcality control, fuel cycle assessment and many others. - 254 - The tools of the reactor physicist are: a) Basic information on the probability of nuclear reactions, on the kinetics of the fission process and on the decay of radioactive species. This information is compiled in data libraries. b) Mathematical equations describing the Interaction of the neutron field with the core materials, because of the large number of neutrons present in an operating reactor core (typically 500 million per cnW) the physical laws represented by these equations are of a statistical nature. Because of the complexity of the required calculations, these are usually performed by large computer programs, commonly referred to as computer codes. The required nuclear data library is conveniently attached to the code on a computer accessible storage medium. Unfortunately, although the basic physical laws governing the interaction of neutrons with matter are well known, the actual solution of the pertinent equations Is so complex and time consuming that it cannot, even on the fastest available computers, be performed without some approximations. Moreover, the information in the required data library Is so detailed and voluminous that even today, more than 40 years after the discovery of nuclear fission, the data collections are periodically updated as new measurements become available, and compilation errors are removed. As a consequence of this, all approximations made in the computer codes must be reviewed when new, unfamiliar fuel cycles are to be analyzed. Subsequently, the codes in conjunction with their data libraries must be tested against relevant experimental Information to verify both code approximation and data. This stage In the analysis is usually referred to as code and data validation. - 255 - Once the tools are validated, they can be used with confidence to derive the required core and fuel cycle characteristics. Complications in the Reactor Physics Analysis of Advanced Fuel Cycles Our present reactor physics codes have been developed for the once- through natural uranium fuel cycle- This fuel cycle Is unique compared to the advanced fuel cycles now being considered. The fissile component of natural uranium is set by nature at a very uniform 0.711 weight percent U235. This enrichment, somewhat low for a very efficient use, required a number of special measures to make natural uranium reactors viable. Such measures are, for instance, the use of heavy water as the neutron moderator and the applications of neutron conserving structural materials, furthermore a highly heterogeneous fuel channel design and engineering features to allow on-power fuelling. Computational tools to deal with the numerous advanced fuel cycles of interest must be developed from this now familiar background. In doing so a number of complications are encountered which are due mainly to the following characteristics: - A number of additional nuclides enter into the analysis. - The higher fissile concentration causes the energy distribution of the neutron field to shift towards higher energies. - In thorium cycles the fuel contains protactinlum-233 and uranium-233 in concentrations which depend on the previous Irradiation history. - 256 - The characteristics of several nuclides can no longer be evaluated in isolation (resonance Interaction). Plutonium, a possible fissile topping material, depletes quickly which in turn results in changes In the energy distribution of the neutron field. The higher fissile content of the fresh fuel in connection with the highly heterogeneous core design of present CANDU reactors results In undesirable power peaking effects. - In a recycle program the exact composition of irradiated fuel Is of Interest, while for natural uranium systems only the average exit burnup was important. Particularly the concentra- tion of certain fission products and nuclides such as U232 and Th228 which are of minor importance for neutronics calculations become crucial for reprocessing and fabrication. In an effort to acquire reactor physics tools which can in a reliable and efficient manner generate the required information, a program of experiments and code development has been set up. Definition of Program Objectives The overall objective is to develop and validate efficient reacLor physics methods and to generate reliable data libraries for use in the development of advanced nuclear fuel cycles. Current sub- program objectives are as follows: 1. To carry out a program of measurements on unirradlated recycle fuels (i.e. uranium or thorium fuel containing U233, U235 or plutonium) to yield information to validate reactor physics data and methods. - 257 - 2. To carry out a program of measurements on irradiated recycle fuels to yield information on isotope and reactivity changes to validate reactor physics data and methods of calculation. 3. To develop, upgrade and validate reactor physics data and methods of calculation. 4. To develop and upgrade data and methods of assessment and calculation for radiological safety studies. These four sub-objectives are being pursued in four respective programs by the reactor physics working party (Table G.2). The Experimental Programs on Un-irradiated and Irradiated Fuels While the reactor physics tools require validation for cores with fuel of all irradiation stages, including fresh fuel, it is in practice much simpler and more precise to perform, as a first step, measurements on fresh fuel only. Such measurements can be performed in ZED-2, the small 'zero power' reactor at Chalk River Nuclear Laboratories. Even if there is not enough oi the 'test' fuel available to fill the entire ZED-2 core, there are techniques to determine the characteristics of the test fuel even if it is present only as a central region, surrounded by a reference-fuel of kiown character- istics. Measurements of this kind are referred to as substitution experiments since the test fuel substitutes part of the reference fuel. Although a substitution experiment is much more difficult to analyse than a uniform lattice, this technique requires considerably less of the very expensive fuel. - 258 - It is planned, with some generalization, to acquire about 37 fuel bundles of each of the following three fuel types: Uranium-plutonium, (Pu,U)O2 Plutonium-thorium, (Pu,Th)C>2 and Uranium-233 - thorium, (U233,Th)O2 Each fuel type will initially be analyzed in ZED-2 substitution experiments to determine its basic unirradiated characteristics and will then be inserted into a small power reactor - most likely WR-1 at the Whiteshell Nuclear Research Establishment - to study the modifications to the reactor physics characteristics which occur under irradiation. Following this irradiation phase each fuel batch will undergo a detailed destructive chemical and isotopic analysis. Data and Methods The reactor physics tools necessary to simulate and predict the nuclear behaviour of advanced CANDU reactor cores can bt classified as follows: - Nuclear data libraries Lattice codes - Reactor burnup and fuel management codes - Reactor kinetics and dynamics codes - Nuclide analysis codes The nuclear data libraries contain the basic interaction data between neutrons and the nuclei of the various materials in a reactor core. - 259 - Lattice codes are used to derive from this basic information and from a description of a single fuel channel with its share of surrounding moderator the detailed nuclear characteristics of this channel. These characteristics can include the changes that occur in the fuel when the fuel channel - a reactor lattice cell - is irradiated. Reactor burnup and fuel management codes integrate the individual fuel channel characteristics into a description of the full reactor core, including control devices and the actual design elements that are necessary for an operating core. This class of codes is also capable of simulating various refuelling and fuel-management schemes. In short, these programs predict the stationary long-term behaviour of a complete reactor core. In contrast, reactor kinetics and dynamics codes analyze the short- term response of the neutron population to a perturbation of the core. They are used to determine the margins of stability, to simulate the influence of control manoeuvres, and are part of accident analyses. Nuclide analysis codes are employed to predict the exact nuclide composition of fuel material, both in the reactor and outside, in any of the fuel cycle stages when many nuclides are transformed by radioactive decay. Although data libraries and codes exist in all areas for application to natural uranium fuelled reactors, these tools must be extended to cover the wide field of advanced fuels. Accordingly, Table G-2 shows sub-objective for the data libraries as well as for each class of codes requiring implementation of models for those phenomena that are not present in a natural uranium cycle. All new data and code modification require validation by comparison with results from the experimental program. - 260 - Unfortunately, in most cases, such comparison cannot be made directly. This is why work is also being pursued in parallel on the proper analysis of these reactor physics experiments. Radiological Safety While assistance in design and simulation of advanced reactor cores is the primary objective of the reactor physics working party, it is essential that all stages of a nuclear fuel cycle be safe from radiological hazards. This is why a separate sub-objective concerning the description and analysis of radiological accidents is included in the nuclear fuel cycle program. Some of the relevant areas have, of course, been explored in the past in connection with the natural uranium fuelled reactor. It is clear, however, that the introduction of advanced fuel cycles with their higher fissile loadings and the inclusion of reprocessing and fabrication stages requires considerably more work in all five areas listed. Accomplishments and Present Work Experimental Program: Detailed prep?rations are being made for the insertion of a small region of AECL made (Pu,U)02 fuel into ZED-2 (channel simulation), as well as for substitution experiments with about 37 Italian made (Pu,U)02 bundles. Also, assessment and planning activities are underway for the reactor physics measurements in WR-1. - 261 - An analysis of the suitability of present WR-1 measurement systems has recently been completed. Data and Methods: An effort is being made to base all data libraries presently in use on the comprehensive data collection (ENDF/B) maintained by the National Neutron Cross section Center of the U.S.A. This involves the writing or adaptation of processing codes and the establishment of interfaces between the various library formats. While, because of the much longer timescale, experimental data from the experiments on advanced fuel types are not yet available, i^odels for new phenomena are introduced into present codes and are sub- sequently tested. In some cases a sensitivity analysis of integral results to the new phenomena has been performed. Examples for such activities are the investigation into the effects of fuel inhoraogen- eities (fissila particles in the fuel matrix) and a detailed analysis of nuclear data in the resolved resonance region. Radiological Safety: Work in the area of radiological safety has so far concentrated, where available, on the collection of experimental critlcality data for advanced fuels. For the most part this data has now been compared with results of computer calculations. G.5 ASSESSMENT The work on assessment has four major areas a) To determine the characteristics, both technical and economic, of CANDU reactors using different fuel cycles. - 262 - b) To evaluate the Impact on resources, economics and facility requirements of advanced reactor designs when introduced Into a national program. c) To compare CANDU systems with other reactor systems. d) To provide, review and update an economic data bank for the various facilities associated with these nuclear systems. In the studies concerned with resources, facility requirements, economics and comparison with other systems, the major problem Is in the economic data base for reprocessing and fuel fabrication and there is still some uncertainty in the reactor capital charge due to lack of precise knowledge on the attainable channel power rating. A major work area is that of reactor characteristics which will be concerned with fuel management during the approach to equilibrium and at equilibrium, determination of power distributions from in- core detectors, shutdown system analysis, control and effects of using purer materials. Specific examples which illustrate some of these aspects are: a) The bundle radial power distribution For thorla fuel bundles there is both a larger radial power gradient and greater variation in power distribution with time, Figure G.I. Modified designs including both a variation in geometry and enrichment grading will be investigated. The fuel performance and thermal hydraulic behaviour of these new designs will have to be established. - 263 - b) Determination of the reactor power distribution Precise knowledge of the power distribution is needed to ensure that the licensing limits are adhered to. This distribution is Inferred from measurements of intra cell fluxes. The greater variation in the bundle power to cell edge flux ratio with burnup makes the conversion from flux to power less certain. Figure G-2. G.6 SUMMARY The advanced fuel cycle R&D program is designed to enable the complete thorium fuel cycle to be demonstrate' by the end of this century. - 264 - TABLE G-l SEPARATIONS OF INTEREST Complete Fissile Co-Processing Separation Separation Fuel Products Waste Products Waste Products Waste Natural U U,Pu FP ?u U.FP U.Pu FP U-Pu U,Pu FP Pu U.FP U.Pu FP Th-U-Pu Th-U-Pu FP Pu,U Th.FP Th,U,Pu FP Th-U Th-U FP U Th,FP Th.U FP - 265 - TABLE G-2 SUB-PROGRAMS OF THE REACTOR PHYSICS WORKING PARTY EXPERIMENTAL PROGRAM - unirradiated fuels. CANADA/ITALY (Pu,U)O2 fuel. AECL (Pu,Th)O2 fuel. AECL (U233,Th)O2 fuel- Upgrading of test facilities. EXPERIMENTAL PROGRAM - irradiated fuels. AECL (Pu,Th)O2 fuel. AECL (U233,Th)O2 fuel DATA AND METHODS Nuclear data evaluation and processing. Lattice code development and validation. Reactor burnup and fuel management code development and validation. Reactor kinetics and dynamics code development. Development of analysis techniques for the proposed reactor physics experiments. Analysis of reactor physics experiments. Validation of nuclide analysis codes. RADIOLOGICAL SAFETY Criticality analysis. Shielding analysis. Nuclide analysis. Accident initiation analysis. Consequence analysis. - 266 - FIGURE G.I-. BUNDLE RADIAL POWER DISTRIBUTIONS .5 r- .45 Pu/U-233/Th 8 SSET Pu/Th E o NATU .4 37 ELEMENT LLJ FUEL BUNDLE o .35 NATU SSET Pu/U-233/Th .3 _ 240 MWh/kg (U) 480 I 100 200 300 400 500 600 700 800 TIME (d) - 267 - FIGURE 6.2- BUNDLE POWER TO BOUNDARY FLUX RATIO NAT U .0 SSET E Pu/U - 233/Th o a: LxJ G.5 — > 286.5 mm < o Q. Z (th) Q_ 240 MW • h/kg (U) 480 I I h . i I i I 100 200 300 400 500 600 700 800 TIME (d) - 268 - BIBLIOGRAPHY CANDU Nuclear Power System AECL report number TDSI 105, 1981 January A.A. Pasanen Fundamentals of CANDU Nuclear Design AECL report number TDAI-244, 1980 August M.H.M. Roshd The Physics of CAÜDU Reactor Design AECL report number AECL-5803, 1977 April A. Okazaki & Determination of Lattice Parameters Using a Few Rods D.S. Craig AKCL report number AECL-2593, 1967 March A.P. Baudouin & Analysis of Transient Measurements in the ZED-2 Reactor F.N- McDonnell AECL report number AECL-5967, 1977 December A.M. Lopez, Reactor Physics Measurements at Bruce NGS-A J.R. Enselnoz (OH) Ontario Hydro Report CNS-54, 1977 October & V.K. Mohindra J. Griffiths The Effectiveness of LATREP Calculations: A Survey and Detailed Comparison with Experiment AECL report number AECL-3739, 1971 September M.F. Duret Plutonium Production in NPD. A Comparison Between Experiment and Calculation AECL report number AECL-3995, 1971 August E. Critoph Reactor theory and Power Reactors AECL report number AKCL-6150 A.L. Wight CANDU-PHW Fuel Management Ontario Hydro Report 80023, 1980 May E. Critoph Prospects for Self-Sufficient Equilibrium Thorium Cycles in CANDU Reactors AECL report number AECL-5501, 1976 March S.R. Hatcher, Thorium Cycle in Heavy V.'ater Moderated Pressure Tube et al. (CANDU) Reactors AECL report number AECL-5398, 1976 May M.S. Mllgrara & Some Physics Problems in the Design of Thorium Fuelled W.H. Walker CANDU Reactors AECL report number AECL-5561, 1976 August M.S. Milgram Potential of Axial Fuel Management Strategies In Thorium Fuelled CANDU's AECL report number AECL-6182, 1978 June - 269 - ACKNOWLEDGEMENTS The material presented in these lectures was taken, mostly wc-d for word, from published AECL and Ontario Hydro documents. To the many original authors my thanks. Their efforts made the task of preparing these lectures far less onerous than would otherwise have been the case. ISSN 0067 - 0367 ISSN 0067 - 0367 To identify individual documents in the series Pour identifier les rapports individuals faisant we have assigned an AECL- number to each. partie de cette *e>ie nous avons assign^ un nume'ro AECL- a chacun. Please refer to the AECL- number when re- Veuillez faire mention du nume'ro AECL- si questing additional copies of this document vous demanded d'autres exemplaires de ce rapport from Scientific Document Distribution Office Service de Dictribution des Documents Officials Atomic Energy of Canada Limited LEnergie Atomique du Canada Liriite'e Chalk River, Ontario, Canada Chalk Rivwr, Ontario, Canada KOJ 1J0 KOJ 1J0 Price $11.00 per copy Prix $11.00 par exemplaire © ATOMIC ENERGY OF CANADA LIMITED, 1982 562-83