sixth framework programme euratom management of ... · minor actinide burning in elsy [cinotti 08],...

SIXTH FRAMEWORK PROGRAMMEEURATOM

Management of Radioactive Waste

Project acronym: ELSY

Project full title: European Lead-cooled SYstem

Contract no.: FI6W-2006-036439





















Workpackage N°: WP2 Core Design (Task 2.5)

Identification N°: D14

Type of document: Deliverable

Title: MA burning capability of an ELSY core with nitride fuel

Dissemination Level: PP

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Status: Final (Revision 0)






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Name Partner Date Signaturewritten by: J. Wallenius KTH 2010-02-12

Task leader: J. Wallenius KTH 2010-02-12

WP leader: E. Malambu SCK-CEN

Co-ordinator A. Alemberti ANSALDO

MA burning capability of an ELSY core with nitride fuel

Jerzy Cetnar, Grażyna Domańska and Mikołąj OettingenAGH-UST, Krakow, Poland

Haileyesus Tsige-TamiratJRC-IE, Petten, e Netherlands

Kamil TučekKarlsruhe Institute of Technology, Karlsruhe, Germany

Janne WalleniusKungliga Tekniska Högskolan, Stockholm, Sweden

Summary

Mixed oxide fuel is the reference fuel for ELSY, primarily due to vast experience in fabrication and irradiation of MOX in sodium cooled fast reactors. Nitride fuel may offer competitive advantages, such as higher actinide density, higher thermal conductivity and better solubility in nitric acid. Within the ELSY project it was decided to take advantage of the higher actinide density of nitride fuels to evaluate a core with lower fuel column height, aiming at improving safety performance. In the present report, the neutronic performance and the minor actinide burning capability of nitride cores with open square fuel assemblies is assessed using the Monte Carlo codes MCB and MCNPX.

With a core averaged plutonium fraction of 16.5% and an MA mass fraction of 4.0%, it is shown that a very small reactivity swing and a Pu conversion ratio equal to unity may be achieved for the open square assembly design. Hence, a single burnup-cycle of more than five years becomes possible, while keeping power form factors below limits imposed by thermal hydraulic conditions. e low actinide inventory in the core (having an active height of ~75 cm) results in a core averaged burnup of 7.8% heavy atoms aer five full power years at 1500 MWth. Calculated temperature coefficients further indicate that the nitride core, thanks to its small height to diameter ratio, may accommodate more than 4.0% average of minor actinides in the fuel without introduction of moderator pins. e achievable minor actinide burning rate of ELSY with a nitride core may hence exceed 5.5 kg/TWhth.

Introduction

As is well known, advanced fuels, such as nitrides, carbides and metallic alloys, offer better neutronic performance than oxide fuels in fast neutron reactors [Waltar 80]. Single pin irradiations have shown that all these fuels hold the potential for achieveing a burnup of 20%. e use of advanced fuels as driver fuel has however been less common, limited to operation of BR-10 (nitrides and carbides), EBR-I and EBR-II (metallic alloys) and FBTR (carbides). Among the issues that must to be adressed for the advanced fuels are poor stability at high temperature of americium compounds, fuel-clad mechanical interaction and complexity of fabrication and/or reprocessing. In the context of minor actinide burning in ELSY [Cinotti 08], the very low dissociation temperature of americium carbide is a major problem for fabrication of the fuel. Metallic alloy fuels on the other hand are not compatible with lead and will dissolve into the coolant. Hence the remaining alternative fuel option for ELSY is the nitride fuel, which also is considered as the reference fuel for the lead cooled BREST reactor design in Russia [Adamov 97] and the SSTAR design in the US [Smith 08].

Nitride fuels have a higher actinide density, higher thermal conductivity and better solubility in nitric acid than corresponding MOX fuels [etford 03]. Hence, for identical conversion ratios and linear ratings, one may design a smaller core operating at lower fuel temperature, which will lead to improved safety characteristics. e latter fact is of particular significance for MA burning capability, as a nitride core may be expected to permit a higher loading of americium. Drawbacks include stability of AmN during fabrication processes, a low creep rate, and the production of C-14 due to (n,p) reactions on N-14.

e present report describes the neutronic performance of ELSY cores with nitride fuel. e open square assembly design has been used as basis, relying on the geometry described in Deliverable D6. e major tool has been the advanced Monte Carlo code system MCB, developed at AGH in Krakow [Cetnar 98, Cetnar 06]. is code includes convenient features for simulating fuel shuffling and control rod movement, which turns out to be highly relevant for accurate modelling of ELSY operation. e burnup calculations with MCB were performed with JEFF-2.2 and ENDF/B-6.8 nuclear data sets, processed to include gamma heating for MCB application. We also performed calculations with MCNPX [Hendricks 06] in conjuction with JEFF-3.1 data.

Aer describing MCB , we present results from modelling of the open square reference MOX core, showing that MCB not only well reproduces the outcome of the design work made with ERANOS, but also points out issues that only can be discovered and adresssed with proper modelling of control rod movement. en we present the design of a nitride core with smaller height and minimal reactivity swing, and evaluate its performance in terms of irradiation cycle characterization, evolution of power profile and reactivity coefficients. e results of C-14 production and comparison with the oxide reference fuel are also presented. e minor actinide burning capability of this core is discussed in comparison to the oxide core.

Monte Carlo burnup code system – MCB

e Monte Carlo Continuous Energy Burn-up Code (MCB) is a general-purpose code for the calculation of nuclide density evolution with time (aer burn-up or decay). e code performs eigenvalue calculations of critical and sub-critical systems as well as neutron transport calculations in fixed source mode to obtain the reaction rates and energy deposition that are necessary for evaluation of the burn-up. MCB internally integrates the well-known MCNP code (currently - version 5), which is used for neutron transport calculations, and a novel Transmutation Trajectory Analysis (TTA) module, which forms on-line, analyses and solves specific transmutation chains and then calculates the nuclide density evolution [Cetnar 06]. Version MCB1C became available to the scientific community on a freeware basis through the Nuclear Energy Agency Data Bank (Package-ID: NEA-1643) in 2002. e MCB code was applied in many fuel cycle studies of nuclear reactor systems including “PDS-XADS” [Cetnar 05] of the 5th FP, which concerned a lead-bismuth cooled subcritical reactor. e code has been under constant development for twelve years; recently added features include statistical analysis of burnup, emitted particle collection, thus enabling direct helium production assessment, thermal-hydraulic coupling and power form factors assessment [Cetnar 10]. Further development is directed towards improved and more detailed description of advanced reactors and development of new analytical tools of Monte Carlo burnup analysis.

In MCB, the decay schemes of all possible nuclides and their isomeric states are formed and analysed based on decay data taken from two sources. e first one – TOI.LIB, which is prepared based on “Table of Isotopes”[Firestone 96], describes decay schemes for over 2400 nuclides and includes excited state nuclide formation.

Cross-section libraries and data sets can be loaded in support of adequate calculation of reaction rates and nuclide formation probabilities. e data treatment is separate for every zone, which concerns the cross section temperature, energy dependent distribution of fission product formation, and energy dependent formation of isomer nuclides.

Reaction rates are calculated exclusively by the continuous energy method, with usage of point-wise transport cross-section libraries or, in case of lack of such data, by using dosimetry cross section libraries. e contributions to reaction rates are being scored at every instant of neutron collision occurring in cells filled with burnable material. For this purpose, the code uses the track length estimator of neutron flux.

Fission product yield is calculated from incident energy dependent distributions of fission products prepared separately for every fissionable nuclide.

Heating is calculated automatically in a similar way as the reaction rates during neutron transport, by using heating cross section, that is KERMA factors, contained in standard cross section tables. e code also automatically calculates decay heating rates, where energy of decay is taken from an ORIGEN library.

e time evolution of nuclide densities is calculated with a complete set of linear transmutation chains which is prepared for every zone and time step, hence being automatically adjusted to time evolving transmutation conditions. e code uses an extended linear chain method, based on the Bateman approach, to solve prepared-on-line a set of linear chains that noticeably contribute to nuclide formation. e program calculates transmutation transitions from nuclide to nuclide and prints them out to an output file. Transmutation chains that are formed by the code can be also printed for the nuclides of interest.

Advanced modelling of material processing and system rearrangements is possible, which includes fuel shuffling, reloading and control rod operation (insertion or withdrawal).

Automatic calculation of various system performance parameters is available including power profiles and form factors, neutron multiplication factors, radiotoxicity, decay heat and others.

Analysis of reference oxide core

Prior to the investigation of the nitride fuel option for ELSY, the reference oxide core with open square sub-assembly design has been analysed using MCB. Pin, pellet and SA dimensions were adopted as defined in ELSY delivery report D6 [Sarotto 08]. Note that some of these design parameters were later changed for the final design of the reference core in D8 [Sarotto 09]. is concerns a decrease in smear density of the MOX fuel to allow for higher burn-up. e decreased actinide density then had to be compensated with higher plutonium fractions in the fuel in D8, resulting in a higher reactivity swing than the one calculated in D6. e calculations presented below should therefore be compared to the ones presented in D6.

e MCNP input file for stationary calculations was prepared by ENEA. A detailed analysis of power distribution applying Monte-Carlo methodology has been done, including the statistical fluctuations and the dependence on nuclear data. is has been done for BOL with assumption of 1500 MW thermal power.

e investigation of neutron library influence on the neutronic results has been carried out for two major cross section data libraries: JEFF-2.2 and ENDF/B-6.8. ese neutron libraries of MCB system code are available for temperatures from 300 K to 1800 K, in steps of 300 K, which differ from the reference temperatures; therefore, for the calculation model we have applied the nearest ones shown in Table 1 shows. Since JEFF-2.2 does not include individual nuclide cross sections for lead and silicon, JENDL-3.2 data have been used for these nuclides.

Table 1: Reference temperatures and temperatures used in MCNP model.

MaterialTemperatureTemperatureTemperature

MaterialReference JEFF-2.2 ENDF/B-6.8

Fuel INN - MOX with 13.4% Pu 1223 K 1200 K 1200 KFuel INT - MOX with 15.0% Pu 1223 K 1200 K 1200 KFuel OUT - MOX with 18.5% Pu 1223 K 1200 K 1200 KCladding Steel - T91 753 K 600 K 900 KStructural Steel - T91 713 K 600 K 600 KAssembly Plates - T91 713 K 600 K 600 KLead Chambers - T91 713 K 600 K 600 KControl Cloak - T91 + 15% B4C with 90% 10B 713 K 600 K 600 KControl Assemblies - T91 + 70% B4C with 90% 10B 713 K 600 K 600 KLead Coolant - Pb 713 K 600 K 600 K

Figure 1 presents the flux radial distribution calculated on fuel pin level along the main coordinate axis (OX). e data denote axially averaged values, which are smaller than the peak values at the mid-plane. e absolute values of flux are normalized to a core power of 1500 MWth. e vertical lines show fuel zone boundaries: inner (INN), internal (INT) and outer (OUT) with plutonium enrichments of 13.4%, 15% and 18.5% respectively. e results obtained with two different libraries - JEFF-2.2 and ENDF/B-6.8 - slightly differ from each other in the inner zone. is difference does not occur in the average power density radial distribution, shown in Figure 2, where the visible difference is created mainly by statistical fluctuations.

Figure 1: Average flux radial distribution in the reference ELSY core with oxide fuel.

e power density has been calculated in an homogenized structure of fuel and coolant. e lower flux in the calculations using JEFF2.2 can be explained by higher heating energy cross sections. e radial power distribution is very well levelled out by the introduction of zones with different enrichments. Within the intermediate and outer zone, where the enrichment is higher, the power density peaks in locations close to the inner boundary, but only by few percent above the average value. is is result of the higher fluxes there.

Figure 2: Average power density radial distribution in the reference ELSY core with oxide fuel.

Other differences occur due to presence of control rod assemblies. ese become visible in the two-dimensional power distribution shown in Figure 3, as well as in the sub-assembly power density distributions shown in Figures 4 and 5. e green squares in the figures represent the control rod assembles, which are fully withdrawn in the power calculation model. Here, the control rods locations suppress the flux and create peaks in the locations between them.

Figure 3: Two-dimensional visualization of average power density distribution in the ELSY core with oxide fuel.

Figure 4. Average power density [W/cm3] distribution in the ELSY core with oxide fuel.

Figure 5: Peak power density [W/cm3] distribution in the ELSY core with oxide fuel.

A summary of the power deposition is shown in Table 2. e total power released in the active core is about 1470 MW; the remaining 30 MW, or 2% of 1500 MW total power, is deposited in the surrounding coolant or in exterior structures. e above results confirm that the zone enrichments were chosen adequately to level out the power distribution at beginning of life (BOL). e average power density changes from zone to zone at maximum by 2.6%. e differences in power density on the zone level due to cross section libraries also are smaller than 1%.

Table 2: Summary of power deposition in oxide fuel SA-s of the reference ELSY core.

Fuel Zone Number of SAJEFF-2.2JEFF-2.2JEFF-2.2 Endfb6.8Endfb6.8Endfb6.8

Fuel Zone Number of SA[MW] [W/cm-3] [MW/SA] [MW] [W/cm-3] [MW/SA]

INN 132 727 107.0 5.51 733 107.8 5.55INT 72 393 106.2 5.46 393 105.9 5.45

OUT 68 350 100.0 5.15 345 98.6 5.08TOT 272 1470 105.0 5.41 1470 105.0 5.41

e axial power distribution is shown in Figure 6. ese results confirm that the power peak occurs in the second fuel zone. One can also notice the influence of lead on the power distribution. e values at the core edges are a bit higher due to neutron reflection in the surrounding lead. e influence of withdrawn CR-s is visible in higher points of the outer zone, where power is suppressed. e neutron absorption there overshadows the effect of neutron reflection from lead. Some differences in local power deposition might be thought of being the due to applied cross section libraries. is, however, may be also explainable, at least partially, by statistical fluctuations.

Figure 6: Axial power profile in peak subassemblies of each zone.

e calculated power distributions have been used for a concise characterisation of the core in terms of power form factors. ese form factors have been evaluated at the SA level, using average density and peak density as shown in Figure 4 and 5. ree types of form factors have been calculated, which are defined as follows:

• radial: the ratio of maximum to average SA average density• axial: the ratio of SA peak to average density calculated for the SA with the maximum average power density• total: the product of radial and axial form factors

e obtained form factors are displayed in Table 3. For the given power distribution it turns out that the intermediate zone is responsible for the form factors of the entire core. Generally, the current results remain in good agreement with the ones of deliverable D6 [Sarotto 08], and remain below the assumed limits of ffrad_core =1.20, ot_core =1.45, with the marginal exception of the JEFF-2.2 results. e observed deterioration of the form factors in the intermediate and outer region as compared to the inner region are caused primarily by the interference of control rods.

Table 3: Power Form Factors at BOL of ELSY core with oxide fuel.

Core region INN INT OUT Entire CoreForm factor JEFF-2.2 resultsJEFF-2.2 resultsJEFF-2.2 resultsJEFF-2.2 results

Radial 1.074 1.139 1.145 1.151Axial 1.083 1.256 1.160 1.256Total 1.171 1.431 1.328 1.447

ENDF/B-6.8 resultsENDF/B-6.8 resultsENDF/B-6.8 resultsENDF/B-6.8 resultsRadial 1.052 1.144 1.181 1.153Axial 1.168 1.192 1.200 1.192Total 1.228 1.363 1.412 1.374

e power peaks are located in the SA-s that are most distant from the control elements, that is along the major axes of the core (OX and OY). is may be improved if a few SA-s had lowered enrichments. It should be noted that the

above results were obtained with the assumption of withdrawn control rods. is condition influences the axial form factors adversely, apparently. erefore, it would be worth studying the influence of CR insertion level on the form factors.

Neutron Multiplication Factor

e criticality mode of MCB was used for calculation of k-effective with a simulation of 4,000,000 neutrons distributed over 400 active cycles and 40 inactive ones (10,000 per cycle), resulting in an estimated statistical error of 45 pcm. As shown in Table 4, ENDF/B-6.8 data give a k-effective 1.85% higher than JEF-2.2. Here one should recall that JENDL-3.2 data were used for lead nuclides in the ”JEFF-2.2” calculation.

Table 4: Effective neutron multiplication factor at BOL

BOL JEFF-2.2 ENDF/B-6.8keff 0.9808 0.9988

Burnup performance

A positive reactivity swing of about 900 pcm was calculated for the reference core using ENDF/B-6.8 data, as displayed in Figure 7. is is in fair agreement with burnup calculations made by ENEA using ERANOS and JEFF-2.2 data [Sarotto 08]. In order to avoid the need to compensate for the reactivity swing with control rod movement, which might affect power form factors adversely, one could consider to increase the plutonium content slightly.

Figure 7: Evolution of criticality in the oxide core calculated using MCB and ENDF/B-6.8 libraries.

e nuclide density evolution was calculated for a core configuration with 110 cm active height and 25 cm control rod insertion and is shown in Table 5. A positive plutonium balance again indicate that an increase of initial Pu fraction in the fuel may make sense. e total burnup equals 5.4% heavy atoms, a rather modest number, supporting the decision taken to redefine the reference oxide core to increase its burnup [Sarotto 09].

Table 5: Nuclide mass evolution in an oxide core with active core height 110 cm and a power of 1500 MWth.

NuclideMass [kg]Mass [kg] Mass change aer 5 yearsMass change aer 5 years

Nuclide BOL Aer 5 years [kg] [%]

Pu238 180.6 131.1 -49.5 -27.4

Pu239 4403.0 4732.0 329 7.47

Pu240 2090 2208 118 5.65

Pu241 472.6 368.4 -104.2 -22.05

Pu242 595.6 556.2 -39.4 -6.6

Pu 7741.8 7995.7 253.9 3.28

U234 1.25 6.35 5.1 408

U235 167.9 96.46 -71.44 42.5

U236 4.16 19.56 15.4 370

U238 41390 38350 -3040 7.3

U 41563 38472 -3091 -7.43

Np237 0 8.73 8.73

Np239 0 4.88 4.88

Np 0 13.61 13.61

Am241 0 77.42 77.42

Am242m 0 3.08 3.08

Am243 0 52.4 52.4

Am 0 132.9 132.9

Cm242 0 2.44 2.44

Cm243 0 0.12 0.12

Cm244 0 10.63 10.63

Cm245 0 0.46 0.46

Cm 0 13.65 13.65

Total 49305 46628 2677 -5.4

Analysis of nitride core

e application of nitride fuel core provides a potential for a significant improvement of neutronic characterization in terms of safety and neutronic performance of the core in terms of residence time, reactivity swing and power form factors. To achieve this we need to exploit specific features of the nitride fuel that may allow us to reduce the reactivity swing and prolong irradiation time, while respecting other design constraints.

e design parameters are constrained as follows:

• e open square lattice sub-assembly design is adopted as a reference for the nitride core, with pin, pellet and SA dimensions defined in deliverable D6 [Sarotto 08] of ELSY. • e TRU vectors used are those specified in deliverable D5 [Sobolev 07] of ELSY.

• e core power is fixed to 1500 MWth• e active core height should be reduced from the 90 cm used for the oxide core, in order to reduce the coolant temperature coefficient.• e radial power form factor is to be kept below 1.3, respecting limiting values for outer clad temperatures at 550°C (without GESA) or 600°C with GESA treatment.• e N-15 enrichment of nitrogen is a priori set to 99% to achieve a C-14 production rate comparable to that of the oxide reference core.• Fuel porosity at BOL is fixed to 15%, in order to allow for 1% swelling per percent burnup in peak pin. is corresponds to experimental data for UN swelling at T = 1173 K [Zaboudko 06]• MA fraction in the the MA burning core variant is to be in the the range of 2.5 – 5.0 % (MA: Np, Am, Cm)• ermal expansion data for (U,Pu)N is adopted from the review of Zabudko [Zaboudko 06]• Other physical properties of fuel are taken from etford and Mignanelli [etford 03].• A total cooling time of 7 years is subdivided into 5 years before reprocessing and 2 years aer fabrication of the fuel before the next core loading.• Recommended negative reactivity margin: 5%.

Our task is possible due to bigger flexibility in composition of the denser nitride fuel, which allows for reduction of the fissile fraction in the fuel, and/or introduction of an inert matrix fraction to the fuel, adding another degree of freedom to achieve better fuel zoning. e higher thermal conductivity of the nitride fuel further improves margins to fuel failure for a given power linear rating.

In the search for the optimal configuration, we aimed at extending the irradiation period without any rearrangement of the core. Necessarily, during this irradiation period the limiting constraints of core parameters must be satisfied. A substantial extension of of the irradiation periods up to and beyond 5 years of effective full power would allow the reactor to run in one batch fuel cycle without intermediate refuelling and reshuffling. Such a feature significantly simplifies reactor design, since systems dedicated to decay heat removal during reshuffling can be avoided. is incentive should be recognized as an important one, particularly for reactors dedicated to MA burning , where decay heat removal requirements during reload or reshuffle are significant. is is even more important in the suspended core concept of ELSY, where a batch of subassemblies needs to be taken out of the lead coolant during the reshuffle.

e reactivity swing is the first limiting factor that may prompt fuel reloading or at least shuffling. However, keeping it within operational limits will not permit extended reactor operation if the power profile does not provide proper thermal-hydraulic conditions. erefore, the radial power form factor is the second limiting factor for the design. e power profile issue becomes more important in case of breeder reactors like ELSY, since the core fuel composition and its spatial distribution undergoes large changes along with irradiation.

Search for an optimal configuration of the nitride core

In order to flatten the power profile at BOL and then keep it under constraints during fuel cycle, it was found that the core could to be divided into four radial regions with different volume of ZrN inert matrix, which gradually decreases outwards. e optimised core configuration is shown in Figure 8. Green squares represent control rod sub-assemblies. e relative fraction of plutonium to minor actinides is also a variable and increases outwards, though only two enrichment levels are used. However, due to inert matrix fractions the absolute Pu and MA mass densities in the fuel gradually increase from region to region outwards. e composition of the four fuel zones is presented in Table 6. is configuration produces the desired power distribution and does not deteriorate with burnup, as will be demonstrated in next sections. Some minor adjustment may be needed latter in order to provide the optimal level of control rod insertion giving zero reactivity at BOL.

!Figure 8: Nitride fuel core map for the open square SA design of ELSY

Table 6: Composition of fuel zones with inert matrix in near optimal configuration

ZoneVolume fraction [%]Volume fraction [%] Atomic fraction in HM [%]Atomic fraction in HM [%]

ZoneZrN (U,Pu,MA)N Pu MA

I 30 70 15.0 2.5II 20 80 15.0 2.5III 10 90 17.5 5.0IV 0 100 17.5 5.0

Neutronic characterization at BOL

For the given core configuration, neutronics at BOL was investigated. First, the required level of CR insertion in order to achieve zero reactivity zero was obtained. e results are presented in Table 7. All 12 rods were considered to have the same insertion. e control rod insertion level was measured form the core mid plane to the bottom of the CR. e total CR movement of 110 cm spans a higher range than the active core height of about 77 cm. e insertion depth is measured from the top edge of the active core, therefore the CR full withdrawal level has negative insertion depth.

It may be noted that the entire reactivity worth of 5211 pcm is not symmetrically distributed with respect to the mid plane level, since 50% worth reactivity insertion is obtained for a CR level about 12 cm above the mid plane. Criticality is obtained for 33,5 cm insertion depth. e optimal level of CR insertion depth at BOL cannot be known before assessing the core performance in the irradiation cycle, where values of criticality swings requiring compensation are assessed. As a deeper insertion may help in extending irradiation time by larger reactivity swing compensation and power profile balancing, a shallow insertion provides a larger negative reactivity margin required for safety reasons. is issue will be discussed in more detail later. Further analysis has been carried out for two insertions levels of CR: 25 and 33,5 cm.

Table 7: Reactivity response to CR insertion

CR insertionlevel [cm]

Insertion depth [cm] k-eff ρ [pcm] Δρ [pcm]

55 -16,5 1,0379 3648

11,5 27 1,0086 857 2791

8,5 30 1,0034 335 3313

5 33,5 1,0001 9 3639

3 35,5 0,9983 -167 3815

-55 93,5 0,9846 -1563 5211

Fuel cycle analysis

e core performance was studied for two levels of CR insertion: 25 and 33.5 cm. e single irradiation cycle was assumed to last for a period of 2000 full power days, or about 5.5 years. We first consider the case with CR-s insertion level being fixed over the entire cycle.

Case of 25 cm CR insertion

For the first case, the results of criticality evolution are shown in Figure 9. e results were obtained using MCB with a statistical fluctuation of k-eff about 50 pcm. During the first half of the cycle, the trend line is almost flat with the calculated criticality fluctuating around the trend line with the range allowed by normal distribution. In the second half of the, aer 850 days, reactivity losses of about 1,1 pcm/day are introduced. e total reactivity compensation of about 1150 pcm for a period of 5 years is smaller than the CR reactivity compensation margin of 2790 pcm. e second important parameter to be kept under constraint is the radial power form factor. Its time evolution is shown in Figure 10. A maximal value of about 1.23 is obtained at BOL, then it decreases to reach a minimum of about 1.13 in the middle of period. In the second half of the cycle is grows moderately and does not exceed the BOL value.

Figure 9: Evolution of keff in the nitride core with CR insertion fixed at 25 cm

Figure 10: Evolution of radial power form factor in the nitride core with CR insertion fixed at 25 cm

ese calculations were obtained with a fixed level of CR insertion, which can pose the question how the power profile would change aer CR level adjustment. e required compensation could be achieved by shiing the rods upwards by a few centimetres, which would increase the power density in outer regions. is will slightly lower the radial power form factor as at EOC the power density there is below the average. Below, we confirm that this is the case by performing a burnup simulation explicitly including control rod movement.

Case of 33.5 cm CR insertion

For the second case, the results of criticality evolution are shown in Figure 11, whereas the evolution of radial form factor is shown in Figure 12. is case also satisfies the required constrains concerning the power profile. e criticality swing can also be compensated. ere are, however, some differences in the performance in comparison to the case of shallower CR insertion. In the first half of the cycle, reactivity increases, but very modestly – less than 200 pcm. en it decreases with a similar rate as in the earlier case. e radial form factor performance is changed more significantly in terms of reaching the maximum and minimum values. Generally, the range of change is roughly the same and the power form factor remains well below the constraint.

Figure 11: Evolution of keff in the nitride core with CR insertion fixed at 33.5 cm

Figure 12. Evolution of radial power form factor in the nitride core with CR insertion fixed at 33.5 cm

Modelling of control rod operation

Since the fuel conversion depends on neutron flux distribution, it is important in design studies to include the modelling of the actual insertion level of control rods during the irradiation cycle, due to its influence on the flux. In figures 13, 14, 15 and 16 the effect of applying explicit modelling of control rod operation is displayed for both the above studied cases. As one can see this effect is positive on the core performance, since the radial form factor is stabilized at a lower value. In both cases, it is well below the limits over entire irradiation period i.e. over 5 years. It can be noted that the insertion level of 35.5 cm is better than 25 cm, since maximum radial form factor there is kept below 1.17. Analysis of both cases shows that the required performance of the core can be achieved with different values of initial insertion levels of CR, hence the insertion level is not a critical factor, and may be adjusted to meet other constraints, such as shut down margin.

Figure 13: Evolution of keff in the nitride core with control rods inserted 25 cm at BOL. Red line – fixed CR; blue line, CR in operation.

Figure 14: Evolution of the radial power form factor in the nitride core with control rods inserted 25 cm at BOL. Red line – fixed CR; blue line, CR in operation.

Figure 15: Evolution of keff in the nitride core with control rods inserted 35.5 cm at BOL. Red line – fixed CR; blue line, CR in operation.

Figure 16: Evolution of the radial power form factor in the nitride core with control rods inserted 35.5 cm at BOL. Red line – fixed CR; blue line, CR in operation.

Nuclide density evolution in the nitride core

e evolution of nuclide densities for an irradiation cycle of 5 years is displayed in Table 8. is calculation was made with MCB for an active core height of 72 cm and a control rod insertion of 25 cm at BOL. As aimed for, the plutonium inventory remains nearly constant. anks to the smaller core height and the presence of the inert matrix, the actinide inventory of the nitride core is just 75% of the reference oxide core, which leads to a correspondingly higher burnup for the same total power and cycle length. e nitride core hence reaches an average burnup of 7.8% heavy atoms, with a peak pin burnup of the order of 10%. e minor actinide consumption is 22% at EOL, corresponding to a burning rate of 5.0 kg/TWhth. Albeit not nearly as efficient as a dedicated MA burner like EFIT, the nitride version of ELSY is considerably more capable of burning minor actinides than its oxide counter-part. is is mainly due to better safety coefficients, which permit higher fractions of minor actinides in the fuel. Taking into

account that nearly all Cm-242 and a significant fraction of Cm-244 decays into plutonium during cooling and reprocessing, the actual consumption rate of minor actinides would be closer to 5.5 kg/TWhth.

Burnup calculations with MCNPX and JEFF-3.1 data exhibited a larger reduction of Am-241 and a smaller production of meta-stable Am-242. e latter can be attributed to the fact that MCB properly takes into account the energy dependence of the branching ratio in isomer production following capture in Am-241. In a fast spectrum, the branching ratio into the meta-stable state is about 20%, while for a thermal spectrum it is of the order of 10% [Mann 77]. Hence one may expect MCNPX to significantly underestimate the production of Am-242m, and consequently also slightly over-estimate reactivity losses. JEFF-3.1 data on the other hand should be more accurate for the capture cross section of Am-241 than JEFF-2.2 and ENDF/B-VI.

Table 8: Nuclide mass evolution in the nitride core. Active core height 72 cm, Control rod insertion 25 cm at BOL. Power 1500MWth.

NuclideMass [kg]Mass [kg] Mass change aer 5 yearsMass change aer 5 years

NuclideBOL Aer 5 years [kg] [%]

Pu238 140.91 277.99 137.08 97.3%Pu239 3435.9 3425 -10.9 -0.31%Pu240 1631 1679.5 48.5 2.97%Pu241 368.81 255.42 -113.39 -30.7%Pu242 464.74 458.89 -5.85 -1.26%

Pu 6041.36 6096.8 55.44 0.92%U234 0.876 80.86 79.98 9131%U235 120.1 59.855 -60.20 -50.1%U236 2.946 15.43 12.48 424%U238 29601 26945 -2656 -8.97%

U 29725 27101 -2624 -8.83%Np237 56.35 48.54 -7.81 -13.9%Np239 0 4.187 4.187

Np 56.35 52.73 -3.623 -6.43%Am241 1115 693.2 -422. -37.8%Am242 0.00 0.12 0.12

Am242m 3.76 55.73 51.97 1383%Am243 239.1 199.2 -39.87 -16.7%

Am 1358 948 -410 -30.2%Cm242 0 26.94 26.94Cm243 0.986 2.544 1.558 158%Cm244 44.91 96.64 51.72 115%Cm245 17.12 18.94 1.824 10.7%Cm246 1.33 3.27 1.94 145%

Cm 64.35 148.35 83.99 131%

Total Actinides 37 245 34 347 2 898 -7.8%

Total MA 1 479 1 149 330 -22.2%

e relatively good minor actinide burning capability obviously has to be based on a solid safety performance. Here, the smaller H/D ratio of the nitride core is expected to constitute an additional advantage.

Assessment of safety parameters

Fast neutron reactors may in principle operate with positive coolant temperature coefficients, as long as prompt Doppler feedback and semi-promt axial expansion of the fuel can compensate reactivity insertions. Americium however is strongly detrimental for the magnitude of the Doppler feedback, and even a small fraction of Am in the fuel can reduce the Doppler coefficient to insufficient values. e oxide ELSY core, for instance, suffers a reduction of 24% in the Doppler constant when introducing 2% MA in the fuel [Tucek 09].

e effect of Doppler broadening of capture resonances in fertile nuclides has been evaluated for the core with 25 cm inserted CR-s. e fuel temperature reactivity coefficient also known as Doppler coefficient is defined by

!

e coefficient αD multiplied by fuel temperature T is known as the Doppler constant KD, since αD x T is known to be independent of temperature, at least in sodium cooled reactors with oxide fuel [Waltar 80]. is feature can be used for calculating the Doppler constant from core reactivity calculations with different fuel temperatures. Here, we apply fitting by linear regression to the following expression:

ρ = KD ln T + C

Results of calculated reactivity at eight points of Doppler broadened data libraries for the fuel temperature are presented in Table 8. e standard deviations of presented values equal 5 pcm.

Table 9: Keff and ρ for different temperatures.

T [K] ln(T) keff ρ

500 6.215 1.01185 0.01171600 6.397 1.01097 0.01085700 6.551 1.01015 0.01005800 6.685 1.00960 0.00951900 6.802 1.00909 0.00901

1000 6.908 1.00859 0.008521200 7.090 1.00782 0.007761800 7.496 1.00620 0.00616

Figure 13 shows reactivity as function of ln(T), confirming the logarithmic dependence of reactivity with temperature for the nitride core. e fit results in a calculated Doppler constant of -433±8 pcm.

e Doppler constant may then be used for calculating the Doppler coefficient as a function of fuel temperature, as tabulated in Table 10. If we compare these data to the Doppler feedback of the oxide core [Tucek 09], we may note that for identical MA fractions, the harder spectrum of the nitride core results in a smaller Doppler constant. Since the nitride fuel has a much higher thermal conductivity, fuel temperatures will be considerably lower. For the ELSY oxide core, an average fuel temperature of 1480 K has been assumed [Tucek 09]. e corresponding average temperature of the nitride fuel may be expected to be about 1000 K. Hence, a Doppler coefficient of -0.4 pcm/K can be inferred for the here investigated core, containing 4.0% weight minor actinides on average. us the operational Doppler feedback of the nitride core is of similar magnitude as that of the oxide core.

Figure 13: Fit of calculated data for reactivity as function of fuel temperature.

Table 10: Doppler coefficient as a function of fuel temperature

T [K] αD [pcm/K]500 -0.866600 -0.722700 -0.619800 -0.541900 -0.481

1000 -0.4331200 -0.3611800 -0.241

A summary of safety related parameters is presented in Table 11. One may note in particular that the coolant temperature coefficient is estimated at +0.23±0.05 pcm/K, which is significantly lower than for the oxide core having an active height of 90 cm (≈ +0.47 pcm/K). Comparing with the Doppler coefficient the present core should not only be able to survive a reactivity insertion accident, it may also hold the potential for raising the minor actinide content even further. e delayed neutron fraction (beta) was estimated with MCNPX at 377 pcm at BOL. Beta-effective then would reside in the region 300-320 pcm.

C-14 generation

C-14 production rate was assessed for both fuel options. In the nitride fuel core with 99% enrichment of N-15 the C-14 production rate is about 25 [g/year], while in the oxide fuel core it is about 15 [g/year]. While C-14 today is a significant contributor to releases of radioactivity from oxide fuel reprocessing plants, it is not the leading term. [Beaty 95]. Hence we may conclude that 99% N-15 enrichment is sufficient. An updated cost analysis would have to be carried out to show that the corresponding cost penalty would not be prohibitive.

Table 11: Safety parameters of the nitride core

Source of changeReactivity change [pcm]Reactivity change [pcm]

Source of changeBOL EOL

Doppler constant -433 (±8) -414 (±8)Core 2% axial expansion -967 (±15) -939 (±15)Core 2% radial expansion -292 (±15) -257 (±15)Coolant temperature 200 K change 47 (±10) 46 (±10)Void worths:Void worths:Void worths:Density change of coolant in active part of core

–10%–20%–30%

305 (±15)537 (±15)730 (±15)

191 (±20)277 (±20)394 (±20)

Voiding of active part of core (100%) 3790 (±20) 1450 (±20)Reactor vessel voiding to the level of:- top of the fuel assembly- bottom of the fuel assembly

-883 (±15)-7700 (±20)

-560(±20)-4270(±20)

Control rod worths (12 CR) 5700 (±20) 4500 (±20)

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http://www.sciencedirect.com/science/journal/00223115

http://www.sciencedirect.com/science/journal/00223115

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