Nine thermocouples and one platinum resistance thermometer (PT-100) were used to monitoring the reactor pool temperature. The thermocouples were positioned in a vertical aluminum probe and the first thermocouple was 143 mm above the core top grid plate. The reactor operated during a period of about eight hours at a thermal power of 265 kW before the steady state was obtained. The forced cooling system was turned on during the operation. This experiment is important to understand the behavior of the water temperature in the pool and evaluate the height of the chimney effect.

In the introduction we discussed how the decay heat can be calculated using two possible methods: the statistical method and via summation calculations, but any model for calculating the decay heat is only useful if it is able to describe properly experimental data. In this section we will discuss briefly how the decay heat can be measured and what benchmarks can be used to validate the calculations.

In general terms we can classify the decay heat measurements in two ways: a) radiation detection experiments and b) calorimetric experiments.

Radiation detection experiments consist of measuring the beta energy and the gamma energy coming from a small sample of fissile material that has been irradiated for a known length of time. In these measurements the aim is to increase the sensitivity of the setup for the particular goal of the study and reduce the sensitivity to any other type of radiation, which otherwise can lead to systematic errors. For example, in a beta energy measurement we will be prone to use a thin plastic detector, which has high efficiency for beta detection and a reduced efficiency for gamma rays. Conversely, in a gamma energy measurement one would use detectors of high efficiency for gamma rays and would try to avoid as much as possible the penetration of the betas in the gamma detection setup. Examples of these kinds of measurements can be seen in Refs. (Rudstam, 1990), (Dickens, 1981) and (Tasaka, 1988). These measurements have been also labelled in the past as "nuclear calorimetric" measurements, but the method itself is not truly calorimetric one, since it is only the detection of pure nuclear radiation with very high efficiency.

Real calorimetric experiments consist of absorbing the decay radiations and measuring the heat in the absorber. They follow well developed procedures that have been used extensively in studies of the energetics of chemical reactions. Many calorimetric measurements have been limited to intermediate cooling times, because of the difficulties of building an instrument that responds rapidly to changes in the power level, which is characteristic of short cooling times. Actually, for the application to the decay heat problem, the challenge is to build a calorimeter with a short time constant, or in other words to construct a setup that reacts quickly to the power release of the sample. Examples of such measurements can be found in the works of (Schrock, 1978) and (Yarnell, 1977).

There are several publications that summarize the efforts to improve the decay heat benchmarks covering different time periods. Some examples can be found in the reviews of Schrock (Schrock, 1979), Tobias (Tobias, 1980) and Tasaka (Tasaka, 1988). Nowadays the results of decay heat calculations are compared with the measurements of Akiyama et al. (Akiyama, 1982), Dickens et al. (Dickens, 1981), (Dickens, 1980) and Nguyen et al. (Nguyen, 1997). Decay heat benchmarchs can be found in the work of Tobias (Tobias, 1989), which is considered the standard in the field.

2.1 Correlation development

Experimental evidence shows that, the onset of significant voids, OSV is very close to the onset of flow instability, OFI (Lee & Bankoff, 1993; Gehrke & Bankoff, 1993). Therefore, the prediction of OFI in the present work becomes the problem of predicting OSV. Due to the complicated nature of the subcooled nucleate boiling phenomenon, it is often convenient to predict OSV by means of empirical correlations. In the present work, an empirical correlation to predict the onset of significant void is proposed takes into account almost all

By taking the logarithmic transformation of equation (2) and applying the least squares method, the constants ki, k2, k3 and k4 are evaluated as 1, 0.0094, 1.606 and -0.533 respectively. So the developed correlation takes the following form:

with all water physical properties calculated at the local bulk temperature. This correlation is valid for low pressures at heat flux ranges from 0.42 to 3.48 MW/m2 and L/dh ratios from 83 to 191.

2.2 Bubble detachment parameter

A parameter, rj (the bubble detachment parameter) which indicates the flow stability is defined as follows (Bergisch Gladbach, 1992):

where U is the local velocity, ATsub is the local subcooling and ф is the local heat flux. The physical meaning of r is that it controls the behavior of the steam bubbles formed at active sides of the heating surface. If r decreases below a certain value (rOFI), the steam bubble will detach from the wall, otherwise it will stay there. In order to be sure of the maximum power channels are protected against the occurrence of excursive flow instability, the parameter r must be higher than rOFI by a considerable safety margin. Based on the developed correlation, rOFI can be determined by:

The measurements of interfacial area have been carried out earlier in the field of chemical engineering using chemical reaction and/or chemical absorption at gas-liquid interface (Sharma and Danckwerts (1970)). A lot of experimental studies have been reported and reviewed (Ishii et al.,(1982), Kocamustafaogullari and Ishii (1983)). However, in this method, measured quantity is the product of interfacial area concentration and mass transfer coefficient. Light attenuation method and photographic method were also developed and measurement of interfacial area concentration was carried out. However, the measured interfacial area concentration using these methods is volumetric averaged value and measurement of local interfacial area concentration is impossible. In the detailed analysis of multidimensional two-phase flow, measurements of distribution of local interfacial area concentration are indispensable for the validation of interfacial area transport model. Therefore, the establishment of the measurement method of local interfacial area concentration was strongly required. Ishii (1975) and Delhaye (1968) derived following relation among time averaged interfacial area concentration, number of interfaces and velocity of interface. They pointed out local interfacial area concentration can be measured using two or three sensor probe based on this relation.

Here, T and N are time interval of measurement and number of interfaces passing a measuring point during time interval T. nij and vij are unit normal vector and interfacial velocity of j-th interface. For bubbly flow, assuming that shape of bubble is spherical and sensor of probe passes any part of bubble with equal probability, Eq.(78) can be simplified to

Here, vsz is the z directional (flow directional) component of velocity of interface measured by double sensor probe as shown in Fig.8. vsz is obtained by

vsz = s (80)

sz At

where As is spacing of two sensors (Fig.8) and At is the time interval where interface passes upstream sensor and downstream sensor.

Later, based on local instant formulation of interfacial area concentration, Kataoka et al. (1986) proposed three double sensor probe method (four sensor probe method) as shown in Fig.9. Using this method, time averaged interfacial area is measured without assuming spherical bubble and statistical behavior of bubbles. The passing velocities measured by each double sensor probe are denoted by vsk which are given by

vsk =-A (k=1,2,3) Atk

Then time averaged interfacial area concentration is given by

Most of recent experimental works of local interfacial area measurement are carried out by double sensor probe or three double sensor probe (four sensor probe) using electrical resistivity probe or optical probe.

For practical application, Kataoka et al.(1986) further proposed a simplified expression of Eq.(88) for double sensor probe which is given by

1

1111
1 — cot2 a0 ln(cos^ag) — tan-^a0 ln^in-^a0)

where a0 is given by

Here, |viz| and ctz are the mean value and fluctuation of the z component interfacial velocity.

Hibiki, Hognet and Ishii (1998) carried out more detailed analysis of configuration of gas — liquid interface and double sensor probe and proposed more accurate formulation of interfacial area concentration measurement using double sensor probe. It is given by

Here (Bo is given by

Double sensor probe or three double sensor probe (four sensor probe) has finite spacing between sensors. In relation to sensor spacing and size of bubble, some measurement errors are inevitable. In order to evaluate such measurement errors, a numerical simulation method using Monte Carlo approach is proposed (Kataoka et al., (1994), Wu and Ishii (1999)) for sensitivity analysis of measurement errors of double sensor probe or three double sensor probe. Using this method, Wu and Ishii (1999) carried out comprehensive analysis of accuracy of interfacial area measurement using double sensor probe including the probability of missing bubbles. They obtained formulation of interfacial area concentration measurement similar to Eqs.(91) and (92). The method using Eqs.(89) and (90) underestimated the interfacial area concentration up to 50%.

For adiabatic two-phase flow, many research groups all over the world, carried out measurements of interfacial area concentration mainly using double sensor or four sensor electrical resistivity probes. Most of experiments were carried out for vertical upward air — water two-phase flow in pipe. Some data were reported in annulus or downward flow. Flow regime covers bubbly flow to bubbly-to-slug transition. Some data are reported for annular flow. The experimental database of interfacial area concentration for non-boiling system described above is summarized in Table 1 (Kataoka (2010)).

Measurement of interfacial area concentration in boiling two-phase flow is quite important in view of practical application to nuclear reactor technology. However, in boiling two-phase flow, measurement of interfacial area is much more difficult compared with the measurement in non-boiling two-phase flow because of the durability of electrical resistivity and optical probes in high temperature liquid. Therefore, the accumulation of experimental data in boiling system was not sufficient compared with those in non-boiling system. However, recently, based on the establishment of measurement method of interfacial area as described above and improvement of electrical resistivity and optical probes, detailed measurements of interfacial area concentration become possible and experimental works have been carried out by various research groups. Most of experiments are carried out in annulus test section where inner pipe is heated. However, recently, some experimental studies are reported in rod bundle geometry. The experimental database of interfacial area concentration for boiling system described above is summarized in Table 2 (Kataoka (2010)).

Roy and et al. 1994

R-113 Annulus, Di=19.2mm Do=38.1mm Vertical up L=2750mm

Tl=43.0 — 50.3 C, ATsub=37.1 -29.8 C

Zeitoun et al. 1994,1996 Water Annulus, Di=12.7mm Do=25.4mm Vertical up L=306mm

Tl =11.6-31.1 C

Situ et al. 2004a,2004b

Water Annulus, Di=19.0mm Do=38.1mm jL=0.5 — 2.02m/s Double sensor electrical resistivity Vertical up L=1700mm q= 5 — 200kW/m2 0.1292 — 0.1481 MPa

Tl =95.0-99.0 C

Bae et al. 2008

Water Annulus, Di=21.0mm Do=40.0mm

Vertical up L= 1870mm

Tl =95.0-99.0 C

Yun et al. 2008

Water 3×3 rod bundle pitch: 16.6mm,

G=250 — 552Kg/m2/s Vertical up L=1700mm

Tl =96.0- 104.9 C

Lee et al. 2002,2008

Water Annulus, Di=19mm Do=37.5mm G=478 — 1049.5Kg/m2/s Double sensor

electrical resistivity

Vertical up L=1670mm q= 88 — 359kW/m2 0.01147 — 0.1698 MPa

Tl=84.3- 100.4 C

Table 2. Summary of Experimental Database of Interfacial Area Concentration for Boiling System

Georgy L. Khorasanov1, Anatoly I. Blokhin1 and Anton A. Valter2

1Institute for Physics and Power Engineering Named After A. I. Leypunsky, Obninsk

2Institute for Applied Physics, Sumy 1Russian Federation 2Ukraine

1. Introduction

In critical and subcritical fast reactors functional materials fulfill various tasks, including:

— heat transfer from pins to heat exchangers,

— heat removal from ADS target under dissipation of high energy intensive proton beam in the liquid metal — a source of spallation neutrons.

As such of materials molten heavy metals — mercury, lead, eutectic of lead (45%) and bismuth (55%) and others are using or to be used in future.

Heavy metals posses acceptable for FRs and ADSs neutron and physical characteristics while due to some their properties, for example chemical passivity to water, high boiling temperature, they are better as coolant in comparison to liquid light metal which is now used in sodium cooled FRs such as BN-600 and BOR-60 in Russia.

One of important parameters of functional material considered is a value of neutron absorption in coolant because it is desirable the neutron losses in the core of FR and ADS blanket have to be minimized.

Ways of minimization of neutron absorption in FR are well-known: it is offered to use wrapper less fuel assemblies, low neutron absorbing nitrogen isotope 15N in nitride fuel contents, structural materials with low cross section of neutron capture, etc.

The authors of this paper are pointing out on one more possibility of reducing the neutron losses in the core cooled with lead: it is connected with enrichment of lead isotope, lead-208, from its value in the natural lead isotope mix, equal to 52.3%, up to the value of 99.0% [1-9]. Lead-208 as a twice magic nucleus possesses a very low cross section of neutron radiation capture. This unique feature leads to economy of neutrons in the core and other profitable factors which are listed in the Part I of this paper.

The limiting factor of usage highly enriched 208Pb as the coolant is its high price in the world market. In the ISTC #2573 project [10], executed in the RF, the opportunity of creation of the plant for separation of lead isotopes using selective photoreactions was considered. The complex of calculations and theoretical works were carried out, the outline sketch of the separation installation was developed, and economic and technical estimations of industrial production of highly enriched 208Pb were made. Developers of the ISTC #2573 project expect that at the scale of manufacture equal to 150-300 kg of 208Pb per year its price will be of US $200/kg [11]. But these theoretical predictions have not been confirmed experimentally yet.

Presently lead isotopes are separated in gaseous centrifuges in using tetra methyl of lead Pb(CH3)4 as a working substance. According to estimations given in Ref. 12 the price of lead — 208 with enrichment of 99.0% will be about 1000-2000 US $/kg, which is relatively high for nuclear power plants. For comparison, another heavy metal coolant, Pb-Bi costs approximately 50 US $/kg.

Meanwhile, in nature besides lead of usual isotopic content: 1.48% Pb-204, 23.6% Pb-206, 22.6% Pb-207, 52.32% Pb-208, it can be found lead with higher enrichment of lead-208. Such type of lead can be found in ores and placers containing thorium. Lead-208 is a final product of decay the radioactive nucleus Th-232 and that is why such type of lead is called as radiogenic lead. The period of half decay of Th-232 nucleus is 1.4-1010 year. In ancient ores (~3-109 year) the total content of thorium of 3-5 wt% is usual. In this case concentration of radiogenic lead reaches approximately 0.3 wt%. The enrichment of lead-208 in radiogenic lead is about 85-93%, depending on uranium content in ores and minerals. Uranium-238 produces in isotope mix the input of lead-206 which is product of uranium-238 radioactive decay.

As known, thorium containing ores and minerals can be found in India, Brazil, Australia, Ukraine, Russia and other countries. In Part 2 of this paper the possibility of reprocessing thorium containing ores and minerals for production of thorium-232 and lead-208 for nuclear engineering needs is discussed.

The power of the IPR-R1 TRIGA was raised in steps of about 25 kW until to reach 265 kW. The forced cooling system of the reactor pool was turned off during the tests. The increase of the power was allowed only when all the desired quantities had been measured and the given limits were not exceeded. After the reactor power level was reached, the reactor was maintained at that power for about 15 min, so the entire steady-state conditions were not reached in the core and coolant. The fuel temperature data was obtained by using the instrumented fuel element. The fuel temperature measurements were taken at location B1 of the core (hottest position). The outlet temperature in the channel was measured with thermocouple inserted near the B1 position. One platinum resistance thermometer measured the water temperature in the upper part of the reactor tank. Two thermocouples measured the ambient temperatures around the reactor pool. The IPR-R1 reactor has a rotary specimen rack outside the reactor core for sample irradiation. It is composed by forty irradiation channels in a cylindrical geometry. One type K thermocouple was put during the experiment in Position 40 of the rotary specimen rack (Fig. 5).

We will now concentrate on the data needs for summation calculations. Equation 3, which is the first step in the calculations requires a knowledge of decay constants, neutron capture cross-sections, decay branching ratios and independent fission yields. If this information is available, we can solve the coupled system of differential equations numerically or analytically and we will obtain the inventory of nuclei. The next step is to apply Eq. 2, which requires the inventory of nuclei previously determined from 3, a knowledge of the decay constants and the mean energies released per decay. In this chapter we will concentrate on this last part of the problem: how the mean beta and gamma energies are determined from experimental data and what is the best technique to determine them.

Most nuclear applications involving beta decay rely on data available from databases, see for example the Evaluated Nuclear Structure Data file (ENSDF) (ENSDF, n. d.). The compiled data are typically the result of the evaluation of different measurements, using different techniques, but until now they have been mainly based on the use of Ge detectors (the technique that uses Ge detectors is conventionally called the high resolution technique, since Ge detectors have a very good energy resolution (AE/E ~ 0.15%)). In such experiments the main goal is to determine the levels populated in the decay (feeding probability to the different nuclear levels), as well as the quantum numbers that characterize the levels, since these data provide the basic nuclear structure information. As part of the analysis the level scheme populated in the decay, based preferably on 77 coincidence relations, should be constructed and its consistency should be tested using the intensity balance of gamma rays that populate and de-excite the different levels. Depending on the case, these experiments can suffer from systematic uncertainties.

The ability to predict accurately the void fraction in subcooled boiling is of considerable interest to nuclear reactor technology. Both the steady-state performance and the dynamic response of the reactor depend on the void fraction. Studies of the dynamic behavior of a two-phase flow have revealed that, the stability of the system depends to a great extent upon the power density and the void behavior in the subcooled boiling region. It is assumed that the void fraction in partially developed region between onset of nucleate boiling (ONB) and the OSV equal to 0 and in the fully devolved boiling region from the OSV up to saturation, the void fraction is estmated by the slip-ratio model as:

« = l/[l + {(1 “ X)/X}SPg/Pf ] (7)

Where the slip, S is given by Ahmad, 1970 empirical relationship as:

The true vapor quality is calculated in terms of the thermodynamic equilibrium quality using empirical relationship from the earlier work of (Zuber et al., 1966; Kroeger & Zuber,1968) as:

3.4.1 Pressure drop in single-phase liquid

The pressure drop terms for single-phase liquid regime are given by:

In order to confirm the validity of transport equation of interfacial area, comparisons with experimental data were carried out mainly for bubbly flow and churn flow. The transport equation for bubbly flow is given by Eq.(68). This equation includes turbulent diffusion term
of interfacial area, turbulent diffusion term due to non-isotropic turbulence, sink term due to bubble coalescence and source term due to bubble break up. Each term is separately validated by experimental data.

Kataoka et al. (2011b, 2011c) carried out the validation of turbulent diffusion term of interfacial area, turbulent diffusion term due to non-isotropic turbulence using experimental data of radial distributions in air-water two-phase flow in round pipe under developed region. Under steady state and developed region without phase change, coalescence and break up of bubbles are negligible. Under such assumptions, transport equation of interfacial area concentration based on turbulent transport model, Eq.(68) can be simplified and given by following equation.

Here, R is pipe radius and y is distance from pipe wall. Kataoka’s model for turbulent diffusion of interfacial area concentration, (Eqs.(61) , ( 65) and (67)) was used.

Kataoka et al. (2011c) further developed the model of turbulent diffusion term due to non­isotropic turbulence for churn flow. In the churn flow, additional turbulence void transport terms appear due to the wake of large babble as schematically shown in Fig.10.

Fig. 10. Wake in Churn Flow Regime

For interfacial area transport due to wake of churn bubble, interfacial area is transported toward the center of pipe. The flux of interfacial area concentration in radial direction J^, due to churn bubble is related to the terminal velocity of churn bubble. The flux of interfacial area concentration toward the center of pipe is large at near wall and small at the center of pipe. Then, it is simply assumed to be proportional to the distance from pipe center. Finally, the flux of interfacial area concentration in radial direction, Jai due to churn bubble is assumed to be given by

Jai = KCaiR^ {0.3^/gD}a1 (94)

Then, transport equation of interfacial area concentration based on turbulent transport model in churn flow is given by

In order to predict radial distribution of interfacial area concentration using Eq.(93) or Eq.(95), radial distributions of void fraction, averaged liquid velocity and turbulent liquid velocity are needed. These distributions were already predicted based on the turbulence model of two-phase flow for bubbly flow and churn flow (Kataoka et al. (2011d)).

Using transport equation of interfacial area concentration for bubbly flow (Eq.(93) and churn flow (Eq.(95)), the radial distributions of interfacial area concentration are predicted and compared with experimental data. Serizawa et al. (1975, 1992) measured distributions of void fraction, interfacial area concentration, averaged liquid velocity and turbulent liquid velocity for vertical upward air-water two-phase flow in bubbly and churn flow regimes in round tube of 60mm diameter. Void fraction and interfacial area were measured by electrical resistivity probe and averaged liquid velocity and turbulent liquid velocity were measured by anemometer using conical type film probe with quartz coating. Their experimental conditions are

Liquid flux, JL: 0.44 — 1.03 m/s

Gas flux, JG: 0 — 0.403 m/s

For empirical coefficient, Kcai is assumed to be 0.01 based on experimental data. The condition of flow regime transition from bubbly to churn flow is given in terms of area averaged void fraction, a based on experimental results which is given by

a = 0.2 (96)

Figures 11 and 12 show some examples of the comparison between experimental data and prediction of radial distributions of interfacial area concentration in bubbly flow and churn flow. In bubbly flow regime, distributions of interfacial area concentration show wall peak of which magnitude is larger for larger liquid flux whereas distributions interfacial area concentration in churn flow show core peak. The prediction based on the present model well reproduces the experimental data.

300

о о о о Л

Atmospheric Pressure D=0.06 m jL=0.442 m/s jG=0.134 m/s

0 0.01 0.02 Distanc from wall y(m)

Atmospheric Pressure D=0.06 m jL=0.737 m/s jG=0.135 m/s

<U> C3

J 001 1 002 Distanc from wall y(m)

300

«г 200|- of s-h

Distanc from wall y(m)

Fig. 12. Distributions of Interfacial Area Concentration for Churn Flow

Hibiki and Ishii (2000a) carried out the validation of their own correlations of sink term due to bubble coalescence (Eq.(72)) and source term due to bubble break up (Eq.(70)) using experimental data. They carried out experiments in vertical upward air water two-phase flow in pipe under atmospheric pressure. In order to validate their interfacial transport model, evolutions of radial distributions of interfacial area concentration in the flow direction were systematically measured. Experimental conditions are as follows.

Condition I

Pipe diameter D: 25.4mm, Measuring positions z from inlet: (z/D=12, 65,125),

Liquid flux jl=0.292 — 3.49 m/s, Gas flux jg=0.05098 -0.0931 m/s Condition II

Pipe diameter D: 50.8mm, Measuring positions z from inlet: (z/D=6,30.3, 53.5)

Liquid flux jl=0.491 — 5.0 m/s, Gas flux jg=0.0556 -3.9 m/s

Figures 13 and 14 show the result of comparison between experimental data and prediction using transport equation of interfacial area with sink term due to bubble coalescence (Eq.(70)) and source term due to bubble break up (Eq.(72)). Predictions agree with experimental data within 10% accuracy.

Fig. 14. Comparison between Experimental data and prediction for the variation of interfacial area concentration along flow direction for 50.8mm diameter pipe (Hibiki, T. and Ishii, M. 2000a One-Group Interfacial Area Transport of Bubbly Flows in Vertical Round Tubes, International Journal of Heat and Mass Transfer, 43, 2711-2726.Fig.9)

7. Conclusion

In this chapter, intensive review on recent developments and present status of interfacial area concentration and its transport model was carried out. Definition of interfacial area and rigorous formulation of local instant interfacial area concentration was introduced. Using this formulation, transport equations of interfacial area concentration were derived in details. Transport equations of interfacial area concentration consist of conservation equation of interfacial area concentration and conservation equation of interfacial velocity. For practical application, simplified transport equation of interfacial area concentration was derived with appropriate constitutive correlations. For bubbly flow, constitutive correlations of turbulent diffusion, turbulent diffusion due to non-isotropic turbulence, sink term due to bubble coalescence and source term due to bubble break up were developed. Measurement methods on interfacial area concentration were reviewed and experiments of interfacial area concentration for non-boiling system and boiling system were reviewed. Validation of transport equations of interfacial area concentration was carried out for bubbly and churn flow with satisfactory agreement with experimental data. At present, transport equations of interfacial area concentration can be applied to analysis of two-phase flow with considerable accuracy. However, the developments of constitutive correlations are limited to bubbly and churn flow regimes. Much more researches are needed for more systematic developments of transport equations of interfacial area concentration.

1.1 Reducing of neutron absorption in cores of FRs and ADSs

In Fig. 1 microscopic cross sections of radiation neutron capture, s(n, g), by the lead isotopes 204Pb, 206Pb, 207Pb, 208Pb and the natural mix of lead isotopes natPb in the ABBN-93 (Abagian — Bazaziants-Bondarenko-Nikolaev of 1993 year) system of 28 neutron energy groups [13] are given. The cross sections are cited on the basis of files of the evaluated nuclear data for the ENDF/B-VII.0 version.

As can be seen, the microscopic cross sections of radiation neutron capture by the lead isotope 208Pb for all of the 28 neutron energy groups of the ABBN-93 system are smaller than the cross sections of radiation neutron capture by the mix of lead isotopes natPb, and this difference is especially large, by 3-4 orders of magnitude, for intermediate and low energy neutrons, En <50 keV.

In Fig. 2 microscopic cross sections of radiation neutron capture, s(n, g), by the lead isotope 208Pb and the eutectic lead(45%) — bismuth (55%) in the ABBN-93 system of 28 neutron energy groups are given. The cross sections are cited on the basis of files of the evaluated nuclear data for the ENDF/B-VII.0 version.

As can be seen, the microscopic cross sections of radiation neutron capture by the lead isotope 208Pb for all of the 28 neutron energy groups of the ABBN-93 system are smaller than the cross sections of radiation neutron capture by mix of lead natPb (45%) and bismuth, Bi (55%), and this difference is especially large, by 3-5 orders of magnitude, for intermediate and low energy neutrons, En <50 keV.

Share of neutrons with energies less than 50 keV, En<50 keV, usually is about 20-25% of all neutrons in FR or ADS cores and it increases in lateral and topical blankets of the core.

In Table 1 one-group cross sections of neutron radiation capture by two types of coolants — Pb-208 or the eutectic of Pb-Bi — in the lead-bismuth fast reactor project named as RBEC-M and designed in the Russian Kurchatov Institute [14] are given.

From Table1 follows that the coolant from lead-208 is characterized with minimum one — group cross section, about <s>=0.93-0.94 millibarns. In standard lead-bismuth coolant the value of the same one-group cross section is by ~4 times bigger, about <s>=3.62-3.71 millibarns. In lateral and topical blankets one-group cross sections for Pb-208 by ~6-7 times are less than for Pb-Bi. The small values of one-group sections in RBEC-M cooled with lead — 208 and corresponding excess of neutrons can be used for minimization of fuel load of the core, increasing fuel breeding and transmutation of long-lived fission products in lateral and topical blankets.