ISSN 00406015, Thermal Engineering, 2014, Vol. 61, No. 5, pp. 348–356. © Pleiades Publishing, Inc., 2014. Original Russian Text © F.A. Kozlov, A.P. Sorokin, V.V. Alekseev, M.A. Konovalov, 2014, published in Teploenergetika.

NUCLEAR POWER STATIONS

The HighTemperature Sodium Coolant Technology in Nuclear Power Installations for Hydrogen Power Engineering F. A. Kozlov, A. P. Sorokin, V. V. Alekseev, and M. A. Konovalov State Scientific Center of the Russian Federation–the Leipunskii Institute for Physics and Power Engineering, pl. Bondarenko 1, Obninsk, Kaluzhskaya oblast, 249033 Russia email: [email protected] Abstract—In the case of using hightemperature sodiumcooled nuclear power installations for obtaining hydrogen and for other innovative applications (gasification and fluidization of coal, deep petroleum refining, conversion of biomass into liquid fuel, in the chemical industry, metallurgy, food industry, etc.), the sources of hydrogen that enters from the reactor plant tertiary coolant circuit into its secondary coolant circuit have intensity two or three orders of magnitude higher than that of hydrogen sources at a nuclear power plant (NPP) equipped with a BN600 reactor. Fundamentally new process solutions are proposed for such condi tions. The main prerequisite for implementing them is that the hydrogen concentration in sodium coolant is a factor of 100–1000 higher than it is in modern NPPs taken in combination with removal of hydrogen from sodium by subjecting it to vacuum through membranes made of vanadium or niobium. Numerical investiga tions carried out using a diffusion model showed that, by varying such parameters as fuel rod cladding mate rial, its thickness, and time of operation in developing the fuel rods for hightemperature nuclear power installations (HT NPIs) it is possible to exclude ingress of cesium into sodium through the sealed fuel rod cladding. However, if the fuel rod cladding loses its tightness, operation of the HT NPI with cesium in the sodium will be unavoidable. Under such conditions, measures must be taken for deeply purifying sodium from cesium in order to minimize the diffusion of cesium into the structural materials. Keywords: sodium coolant, hightemperature nuclear power installation, hydrogen power engineering, sodium purification, concentration of admixtures, degassing, loss of fuel rod cladding tightness, sodium puri fication from cesium DOI: 10.1134/S004060151405005X

The future of hydrogen power engineering will depend to a considerable degree on the efficiency of technological hydrogen production processes. As a rule, these processes use highgrade heat, the genera tion of which requires considerable energy resources. The same applies to efficient petroleum products refinery processes. Gas and petroleum products rather than renewable sources of energy are presently used for obtaining such heat. However, this method does not hold much promise [1], because it involves addi tional depletion of resources, the natural reserves of which are far from being unlimited, and the possibility of using nuclear energy for these purposes is being dis cussed. A sodiumcooled fastneutron nuclear reactor may become one of possible sources producing high grade energy [2–4]. At present, an innovative project of a 1200 MW BN1200 sodiumcooled reactor plant intended for its serial construction is being developed [4]. This project involves the use of new technical solutions aimed at achieving better technicalandeconomic indicators and enhanced safety level; part of these solutions has been validated during operation of sodiumcooled reactors in Russia, and the other part has been sub stantiated by the relevant research and experimental

development works for the BN800 reactor. The results from a performed analysis show that the use of traditional sodium technology, central to which is removal of hydrogen from sodium by means of cold traps [5], does not hold promise for development of hightemperature nuclear power installations (NPIs) [6]: the intensity of hydrogen flows from the tertiary to the secondary coolant circuit in a hightemperature NPI increases by several tens of times as compared with that in the BN600 reactor. This is attributed to the following two main factors. First, this is the kinetics relating to the physicochemi cal processes in the liquidmetal system: in accor dance with the Arrhenius equation, the constants determining this kinetics are exponential functions of temperature (as the temperature increases by a few hundreds of degrees, the values of these constants grow by several orders of magnitude). And second, the pressure of hydrogen in the process systems of the ter tiary coolant circuit is essentially higher both in case of hydrogen production and in refinement of petroleum products than it is in the tertiary coolant circuit of modern nuclear power plants (NPPs). Of all admixtures that are contained in hightem perature liquidmetal systems, only hydrogen and

348

THE HIGHTEMPERATURE SODIUM COOLANT TECHNOLOGY IN NUCLEAR POWER

349

Ambient medium

IHE

PHE

J1/AM

Reactor

J2/AM

Primary coolant circuit (sodium) J2/1

Secondary coolant circuit (sodium)

Tertiary coolant circuit (medium with hydrogen)

J3/2

J2

HRS Fig. 1. Scheme of hydrogen mass transfer in the loops of a hightemperature NPI.

cesium were analyzed in this study. It is shown that, if the intensity of hydrogen sources becomes several tens of times higher, new technologies for removing hydro gen from sodium must be developed, with a yield essentially higher than that implemented at modern NPPs. An increased concentration of hydrogen in sodium—100 to 1000 times higher than in the sodium circulating in the secondary coolant circuit of modern NPPs—is the main condition for implementing such technology. In this study, we estimated the possible ingress of cesium into sodium through the tight fuel rod cladding and showed that the sodium circulating in a highpres sure NPI must be purified from cesium to a deeper extent than it is done at the presently existing NPPs equipped with fastneutron reactors, e.g. the BN600 reactor. REMOVAL OF HYDROGEN FROM SODIUM IN A HIGHTEMPERATURE NUCLEAR POWER INSTALLATION The requirements imposed on hydrogen removal systems (HRSs) from sodium are obvious: it is neces sary to ensure the required throughput and capacity of an HRS with respect to sodium hydride and removal of hydrogen from the coolant to the preset concentra THERMAL ENGINEERING

Vol. 61

No. 5

2014

tions. In so doing, both the HRS operational safety and its cost must be taken into consideration. The change of hydrogen concentration in the sodium circulating in the secondary coolant circuit with the preset operating parameters of a hightem perature NPI and with the known design solutions can be calculated from the balance equation [7]. For the secondary coolant circuit this equation has the follow ing form (Fig. 1, without a hydrogen removal system in the primary coolant circuit): (1) J 3 2 = J 2 + J 2 AM + J 1 AM, where J3/2 is the hydrogen flow rate from the tertiary to the secondary coolant circuit, which is calculated from the formula

(

)

J 3 2 = K 3 2 p30.5K S − C2 ;

(2)

J2 is the hydrogen flow rate from the secondary coolant circuit to the HRS in case of hydrogen removal by means of a cold trap (CT), which is determined from the expression (3) J 2 = GCT2β ( C2 − C') ;

J 2 AM = K 2 AMC 2;

(4)

(5) J 1 AM = K 1 AMC1 are the hydrogen flow rates from the primary and ter tiary coolant circuits into the ambient medium (AM); K3/2, K2/AM, K1/AM are the hydrogen permeability coef

350

KOZLOV et al.

ficients of the process heat exchanger (PHE), the sec ondary coolant circuit pipelines and equipment, and the secondary coolant circuit HRS, which were deter mined taking into account the operating parameters of the hightemperature NPI and the adopted design solutions, kg/s; KS is the Sieverts constant equal to 4.34 × 10–7 Pa–0.5; p3 is the partial pressure of hydrogen in the tertiary coolant circuit, Pa; С1 and С2 are the hydrogen concentrations in the primary and second ary coolant circuits, kg/kg; C' is the hydrogen satura tion concentration at the cold trap outlet temperature, kg/kg; GCT2 is the sodium flow rate through the cold trap in the secondary coolant circuit, kg/s; and β is the hydrogen retention factor. The value of C' was calculated using the expression

C' = 10 0.467−3023 TCT ,

(6)

Temperature, °C, of: IHE 900–800 PHE 875–800 Reactor vessel (average value) 850 Secondary circuit pipelines: Hot 875 Cold 800 The calculated permeability coefficients, kg/s, are given below: K3/2 3.2 K2/1 1.6 K2/AM 0.7 K1/AM 5e–3

In view of uncertainty in the selection of structural material and adopted design solutions, the calculation results are tentative in nature.

where TCT is the temperature at the cold trap outlet, K [8]. DETERMINING THE INITIAL PARAMETERS FOR CALCULATING HYDROGEN MASS FLOW RATE For determining the permeability coefficients of pipelines and equipment, we must know their surface areas, thickness, and temperature. For estimating the surface areas of the intermediate heat exchanger (IHE) and reactor vessel, we used the data of [6]. The surface area of the secondary coolant circuit pipelines was evaluated proceeding from the supposed layout of components, namely, taking into account that the process facilities are arranged at a dis tance of 500 m from the reactor plant and are connected with each loop of the secondary coolant circuit by for ward and return pipelines made of Grade 12Kh18N10T steel with a standard size of 325 × 12 mm. The following initial data of the hightemperature NPI were used in the calculations: Number of secondary circuit loops, pcs. Number of heat exchangers, pcs.: Intermediate Process Surface areas, m2, of: IHE (one module) PHE (per one loop) Reactor vessel Secondary circuit pipelines per one loop Wall thickness, mm, of: IHE PHE Reactor vessel Secondary circuit pipelines

6 6 6 577 1000 150 1050 1.5 5.0 30 12

DETERMINING THE HYDROGEN PRESSURE IN THE TERTIARY COOLANT CIRCUIT It is shown in monograph [1] that hightempera ture electrolysis, thermochemical decomposition, and hybrid cycles are the most efficient cycles for obtaining hydrogen. High levels of temperature are required for implementing these processes; as regards hydrogen pressure at different stages of the process, no detailed information is given on this matter in [1], and it is only pointed out that the pressure of all components may be around a few atmospheres. In [9], it is indicated that the hydrogen pressure in the tertiary coolant circuit at the process heat exchanger outlet may reach a few atmospheres, which is two or three orders of magni tude higher than in the tertiary coolant circuit of the BN600 reactor, the level of which is, according to our assessments, around 500 Pa. In refining petroleum products using hydraulic cracking and reforming, the hydrogen pressure is equal to several tens of atmo spheres [10]. Therefore, in estimating the effect the hydrogen pressure in the tertiary circuit has on the hydrogen flow rate in the secondary circuit, we used a parametric approach. The calculation results are given in Table 1. It can be shown that the hydrogen flow rate from the primary circuit into the ambient medium calcu lated from the expression J 1 AM = K 1 AM

K 2 1C 2 , K 1 AM + K 2 1

(7)

is a few orders of magnitude smaller than it is from the secondary circuit into the ambient medium; this is why it is not included in Table 1. THERMAL ENGINEERING

Vol. 61

No. 5

2014

THE HIGHTEMPERATURE SODIUM COOLANT TECHNOLOGY IN NUCLEAR POWER

CALCULATING THE PARAMETERS OF THE SECONDARY CIRCUIT SODIUM PURIFICATION SYSTEM In using CTs as the main component of the hydro gen removal system from sodium, it was assumed that they operate in the traditional mode; i.e., that the tem perature at the trap outlet is 120°С, and the working concentration of hydrogen in sodium was specified in a parametric way. With such mode of operation, the HRSs used in the secondary circuit should accumulate the entire hydrogen arriving from the tertiary circuit except for the leaks escaped from the primary and sec ondary circuits into the ambient medium, i.e., (8) J 2 = J 3 2 − J 2 AM − J 1 AM. It follows from (3) that the flow rate through all cold traps in the secondary circuit is given by J2 (9) GCT2 = . β ( C 2 − C ') There are two approaches to determining the CT volume. The first approach is based on the time for which sodium must dwell in the CT at the specified flow rate through it. For the presently recommended CT design, the coolant dwelling time in the CT τdw is taken equal to 900 s, hence (10) VCT = GCT2τ dw. According to the second approach, the CT volume is determined taking into account the specified CT life time τl.t with the hightemperature NPI operating at the nominal load and the cold trap capacity with respect to sodium hydride C* (it is taken that C* = 10 vol % and τl.t = 1 year) τ M (11) VCT = 100J 2 l.t NaH , C*M Hρ NaH where MNaH and MH are the molecular masses of sodium hydride and hydrogen, kg/mol, and ρNaH is the sodium hydride density, kg/m3. Using the data given in Table 1, we calculated the sodium flow rates through the hydrogen removal sys tem as functions of hydrogen pressure in the tertiary circuit and working concentration of hydrogen in the secondary circuit sodium. The calculation results are given in Table 2. With the hydrogen concentration in the secondary coolant circuit corresponding to operation of modern NPPs at the nominal parameters (0.1–0.2 ppm) [11], and with the hydrogen pressure in the tertiary coolant circuit equal to 0.1 MPa, the sodium flow rate through the CT required for maintaining the hydrogen con centration at a level of 0.2 ppm must be no less than 1400 m3/h, and the CT volume must be no less than 340 m3. Technically, such a system can be constructed. For example, in case of using cold traps similar to those used in the BN600 reactor, the system will consist of THERMAL ENGINEERING

Vol. 61

No. 5

2014

351

Table 1. Hydrogen flow rates from the tertiary circuit to the secondary circuit and from the secondary circuit to the ambient medium vs. the hydrogen pressure in the tertiary II

circuit and its concentration in the secondary circuit C H2 , ppm (for six loops) Hydrogen flow rate, kg/s Hydrogen pressure in the C HII2 = 0.2 C HII2 = 100.0 tertiary circuit pHIII2 , MPa J × 104 J 7 4 7 3/2 2/AM × 10 J3/2 × 10 J2/AM × 10 0.1 0.4 0.9

4.3 8.7 13

1.5 1.5 1.5

1.2 5.5 9.9

0.7 0.7 0.7

Table 2. Sodium flow rate through the CTs in one second ary circuit loop, kg/s, and their volume, m3, vs. the hydro gen pressure in the tertiary circuit and working concentra tion of hydrogen in the secondary circuit sodium

Hydrogen pressure in the tertiary circuit pHIII2 , MPa

Sodium flow rate through the CTs in one secondary circuit loop and their volume

C HII2 = 0.2

C HII2 = 100.0

0.1

360/340*

0/0

0.4

730/670

0.8/600

0.9

1100/1000

1.5/1120

* The numbers in the numerator indicate the flow rate through the HRS CTs, and those in the denominator indicate the cold trap volume. The CT volume was calculated using (10) and (11), and the larger one of these two values was taken.

approximately a hundred of such traps. In our opin ion, such a solution is unacceptable not only from eco nomic, but even from aesthetic considerations! Clearly, in order to reduce the required volume of CTs and flow rates through them, it is necessary to reduce hydrogen flow rates to a great extent (by several tens of times) by using lowpermeable materials. Assessments show that in order to obtain working con centrations at a level of around 0.2 ppm, the perme ability must be reduced by a factor of 100. The data of [12] testify that for temperatures around 800°C such problem has not been solved as yet. Hydrogen flow rates from the tertiary to secondary circuit can be limited or reduced by using special design solutions, but this will entail essential compli cation of the process equipment. The necessary quantity of CTs can be reduced by using efficient regeneration methods that would allow a CT to be used several times [13–15]. In this case,

352

KOZLOV et al.

Table 3. Permeability coefficients for membranes made of different materials Coefficients A and B in equa Temperature tion (13) Source range,°C

Material

A

B

Grade 12Kh18N10T steel

1.90

3787

400–900

[16]

Nickel

3.40

2627

350–550

[17]

Niobium

2.75

1137

100–300

[18]

Vanadium

0.326 3103

100–300

[18]

approximately 2500 kg of hydrogen will be released from the cold trap per year, which will contain around 1.2 × 1014 Bq of tritium. In our opinion, it is advisable not to discharge this hydrogen into the environment, but to utilize using more advanced technologies of the future hydrogen power engineering [1]. In so doing, it should be borne in mind that since the obtained hydrogen contains tritium, its longterm storage for no less than 80 years will be required, after which its use in the industry will become possible because the con tent of tritium in it will have become by that time lower than the maximum admissible concentration. Since a considerable amount of tritium may diffuse together with hydrogen into the environment through the secondary circuit pipelines and equipment, it is also advisable to collect this hydrogen and tritium in dedicated systems. This can be implemented by enclosing the pipelines in special casings. Taking into account the results given in Table 2, we analyzed the possibility of constructing a system for removing hydrogen from sodium by evacuating it through special membranes. The hydrogen flow rate, m3/s through a membrane was estimated using the fol lowing equation: J = K p p20.5 S δ ,

(12)

where Kp is the membrane permeability coefficient for hydrogen, m 2 Pa s ; p2 is the hydrogen pressure in the secondary coolant circuit, Pa; S is the membrane sur face area, m2; and δ is the membrane thickness, m. The coefficient of hydrogen permeability through different membrane materials is determined from the formula

log K p = − A − B T .

(13)

The data on the coefficients of hydrogen perme ability through different materials are given in Table 3.

The membrane surface area S was calculated from the formula J RT K S δ (14) , S= 2 M H C2 K p where R is the gas constant equal to 8.31 J/(K mol), T is temperature, K, and MH is the molar mass of atomic hydrogen, kg/mol. The calculation results are given in Table 4, includ ing the necessary surface area and the number of tubes for the hydrogen removal system designed as a device made of 1mlong 0.3mmthick tubes with a diame ter of 10 mm operating at a temperature of 800°С. The data presented in Table 4 testify that it is possi ble to construct systems for removing hydrogen from sodium by evacuating it through membranes made of materials having high permeability for hydrogen. Steel, nickel, niobium, and vanadium were considered as such materials. For estimating whether it is possible to operate a hightemperature NPI at the nominal parameters with high concentrations of hydrogen in the sodium taking the BN600 reactor as an example, we analyzed the effect of hydrogen on the reactor nuclearphysical characteristics, on the NPI safety, on the corrosion of structural materials, on the possibility of implement ing the technological processes, and on the thermal 1

hydraulic characteristics. It was shown that the con centration of hydrogen in sodium up to 150 ppm does not have an essential adverse effect on the nuclear physical characteristics. In analyzing the effect of hydrogen on the safety of a hightemperature NPI it was supposed that the equi librium pressure of hydrogen above sodium should not exceed 0.1 MPa. If the hydrogen pressure is higher than this level, then in case of an emergency situation caused by loss of tightness of the gas pocket, the pres sure of gas above the sodium will drop to atmospheric pressure, and the sodium will become “foamed” due to release of gaseous hydrogen. The concentration of hydrogen in sodium at which the equilibrium pressure of hydrogen above sodium at the nominal parameters will not exceed 0.1 MPa cal culated using the data of Banus [19] is equal to 108– 114 ppm. Hence, for safety reasons, the hydrogen concentration must be lower than the abovemen tioned values. The effect of hydrogen on the corrosion of struc tural materials and on the change of their properties in a wide range of hydrogen concentrations has not been studied. It should be expected that with oxygen con centration in sodium equal to a few ppm or less than 1 ppm (such concentration of oxygen is required in using refractory materials), a growth of hydrogen con 1 These

assessments were carried out by the employees of the Institute of Nuclear Reactors and Technologies A.P. Ivanov and V.Yu. Stogov.

THERMAL ENGINEERING

Vol. 61

No. 5

2014

THE HIGHTEMPERATURE SODIUM COOLANT TECHNOLOGY IN NUCLEAR POWER

353

Table 4. Necessary membrane surface area and number of tubes in the system for removing hydrogen from sodium by vac uumizing for different materials and with different hydrogen concentrations in the secondary circuit and J2 (at the hydrogen pressure in the tertiary coolant circuit equal to 0.1 MPa per a secondary circuit loop)

C HII2 = 0.2 ppm

C HII2 = 100.0 ppm

Parameter S, m2

number of tubes in the HRS

S, m2

3 × 105

107

74

2400

28

900

number of tubes in the HRS

Membrane material: nickel steel

1.1 × 105

3.6 × 106

niobium

2.7 × 103

8.6 × 104

0.7

22

7 × 102

2.2 × 104

0.2

6

vanadium

7.2 × 10–5

J2, kg/s

centration in sodium will not have a negative effect on the corrosion characteristics of coolant. Hence, oper ation of a hightemperature NPI with increased con centration of hydrogen in sodium is possible with respect to this indicator (the effect of hydrogen on the corrosion of structural materials). It should be pointed out that the use of structural materials at high pres sures of hydrogen in the technological processes that are supposed to be implemented in hightemperature NPIs has presently been mastered in the industry. This also testifies that it is possible to operate hightemper ature NPI systems with sodium coolant at increased concentrations of hydrogen in sodium. From technological considerations, it is dangerous if a crystalline phase (suspensions of sodium hydride) appears in sodium in cooling down the installation (a transition to the operating conditions in which the sodium temperature is decreased to 250°C). To avoid this process, it is necessary that the hydrogen concen tration in sodium did not exceed its solubility at the same temperature. The hydrogen solubility in sodium as a function of temperature is described by formula (6). It can be seen that the solubility of hydrogen in sodium does not exceed 5 and 15 ppm at temperatures equal to 250 and 300°C, respectively. Hence, for this condition to be satisfied, the admissible concentrations of hydro gen in sodium must be limited to the abovementioned values. However, if a need arises to operate at higher concentrations, this limitation can be removed pro vided that a correct combination is maintained between the temperature drop rate and extent of hydrogen removal from sodium in cooling down the installation. THERMAL ENGINEERING

Vol. 61

No. 5

2014

9 × 10–6

DIFFUSION OF CESIUM FROM UNDER THE FUEL ROD CLADDING INTO THE COOLANT OF A HIGHTEMPERATURE NUCLEAR POWER INSTALLATION Specialists of the Leipunskii Institute for Physics and Power Engineering (IPPE) showed [20] that stainless steel that had been in contact with sodium for a long period of time (120 000 h at a temperature of up to 430°C) contains cesium, which diffused into it to a depth of up to 100 µm. As was pointed out above, the physicochemical processes become more intense as the temperature increases. In this connection, and taking into account the fact that cesium had diffused into steel to a depth of 100 µm, the question arose about possible diffusion of cesium into sodium through the tight fuel rod cladding during operation of a hightemperature NPI at the nominal parameters and about possible “impregnation” of its materials with cesium in the course of their being for a long time in contact with sodium that contains cesium. Since the fuel rod cladding thickness to the curva ture radius ratio has a small value, the fuel rod cladding can be regarded as a flat wall. In view of this, the fol lowing relations are valid:

n(x, τ) = n1 + (n2 − n1) x x0 ⎛ ⎞ n2 cos(πk) − n1 sin ⎜ πk x ⎟ e +2 π k =1 k x ⎝ 0⎠

∑

⎡ Dn1 ⎢ n ∂ = 1+ J Cs(τ) = −D x0 ⎢ ∂x x = x0 ⎣

(15)

2

∞

∞

⎛ ⎞ −⎜ π k ⎟ D τ ⎝ x0 ⎠

∑ cos(πk)e k =1

;

⎛ ⎞ −⎜ πk ⎟ Dτ ⎤ ⎝ x0 ⎠ ⎥ 2

⎥ ⎦

,(16)

354

KOZLOV et al.

D × 1015, m2/s

n(x, τ)/n1 1.0

10–1

[1000; 0.185]

1 2

10–2 3

[1000; 9.5 × 10–3]

0.5

10–3

4 2 3

1 10–4 700

800

900 Temperature, °C

1000

0

The distribution of cesium concentration over the fuel rod cladding thickness n(x, τ) was obtained, and the cesium flow rate into the coolant JСs(τ) through the tight fuel rod cladding at a temperature of 1000°C was estimated. It can be seen from Figs. 3, 4 that for the assumed fuel rod service time (equal to 2.5 years), cesium will diffuse into the steel cladding from the fuel–fuel rod cladding boundary to the fuel rod–coolant boundary (the fuel rod cladding thickness is 0.3 mm) and to a depth of 0.1 mm from the fuel–fuel rod cladding boundary in case of using a molybdenum cladding. The flow rate of cesium at the fuel rod–coolant boundary reaches the value of 2 × 10–16 kg/(m2 s) for the cladding made of steel, and for the cladding made of molybdenum its flow rate will be negligibly small. This means that release of cesium may take place through the 0.3mmthick tight fuel rod cladding made of steel dur ing the fuel rod service life, whereas the ingress of cesium into sodium through the 0.3mmthick tight fuel rod cladding made of molybdenum can be neglected. It fol lows from (15) and (16) that the ratio of stationary flow rates of cesium through the steel and molybdenum claddings is equal to the ratio of diffusion coefficients through them: (1.88 × 10–16/9.2 × 10–18) = 20, and the time taken for the process to achieve steadystate parameters is equal to the inverse quantity; i.e., this time for the molybdenum cladding is a factor of 20 longer than it is for the steel cladding.

x, mm

0.2

Fig. 3. Distribution of cesium concentration over the 0.3mmthick fuel rod cladding in molybdenum (1, 2) and steel (3, 4) for different intervals of time: 1, 3—1 year and 2, 4—2.5 years.

Fig. 2. Diffusion coefficient vs. temperature. (1) Cesium in steel [21], (2) cesium in steel [22], and (3) cesium in molybdenum [23].

where x varies from 0 to x0, m; τ is time, s; n1 is the concentration of cesium on the cladding surface on the fuel side, kg/m3; n2 is the concentration of cesium at the cladding–coolant boundary, kg/m3; and D is the coefficient of cesium diffusion in the fuel rod clad ding, m2/s, which is an exponential function of tem perature (Fig. 2).

0.1

The ingress of cesium into sodium through the tight fuel rod cladding can be reduced by varying the parameters of the complex D τ x 02 . For example, if we increase the steel cladding thickness by a factor of 2, the ingress of cesium into sodium through it can be neglected. If a loss of fuel rod cladding tightness occurs, oper ation of the hightemperature NPI with certain con tent of cesium in the sodium will be unavoidable. In view of this, diffusion of cesium into the structural materials should be minimized in order to improve the radiation environment. To this end, the sodium must JCs(τ), ×1015 kg/(m2 s) 1 10–1

2 10

–2

10–3 0

20

40

60

80

τ, years

Fig. 4. Time dependence of cesium flow rate into sodium for a 0.3mmthick cladding made of steel (1) and molyb denum (2). THERMAL ENGINEERING

Vol. 61

No. 5

2014

THE HIGHTEMPERATURE SODIUM COOLANT TECHNOLOGY IN NUCLEAR POWER

be subjected to deep purification from cesium to the minimal concentrations by means of the methods used in NPIs, such as BOR60, BN600, and EBRII. It should be pointed out that in calculating the sol ubility of hydrogen in sodium and permeability of hydrogen through structural materials as functions of temperature, we had to extrapolate the data presented in the literature from the temperature region 300– 400°С to the interval of temperatures typical for high temperature NPIs. But since such extrapolation has not hitherto been substantiated experimentally, numerically, or theoretically, it should be emphasized once again that the presented data are tentative in nature. CONCLUSIONS (1) It has been shown from the results of the per formed investigations that in case of using hightem perature NPIs (and other innovating technologies) for obtaining hydrogen, the sources of hydrogen entering from the tertiary to secondary circuit have intensity two to three orders of magnitude higher than the sources of hydrogen at an NPP equipped with a BN 600 reactor. Under such conditions, with using the tra ditional technology for operation of a sodiumcooled NPI, and with the need to maintain the working hydrogen concentration equal to 0.1–0.2 ppm, bulky metal and energyintensive systems for removing hydrogen from sodium will have to be developed and operated in the case of using both cold traps and other proposed innovative technical solutions, including those based on evacuating hydrogen from sodium through materials with high permeability with respect to hydrogen. (2) The proposed innovative sodium treatment technology in case of hydrogen concentrations at a level of several tens of ppm taken in combination with removing hydrogen from sodium by evacuating it through membranes made of vanadium or niobium makes it possible to develop hightemperature com pact and highefficient systems for removing hydrogen from sodium. (3) A preliminary study aimed at analyzing the effect of high hydrogen concentrations in sodium on the nuclearphysical characteristics of the reactor, on its safety, on the corrosion of structural materials, on the implementation of technological processes, and on maintenance of the required thermalhydraulic parameters has shown that it is in principle possible to use such technology in hightemperature NPIs. (4) The possible ingress of cesium into sodium through the tight fuel rod cladding has been estimated using the diffusion model. It has been shown from this analysis that ingress of cesium into sodium through the tight cladding can be excluded by varying such param eters as cladding material, its thickness, and time of operation. THERMAL ENGINEERING

Vol. 61

No. 5

2014

355

(5) For securing better radiation environment, measures must be taken for purifying sodium from cesium to the minimal concentrations by applying the methods used, for example, in the BOR60, BN600, and EBRII reactors. REFERENCES 1. E. E. Shpil’rain, S. P. Malyshenko, and G. G. Kuleshov, An Introduction to Hydrogen Power Engineering, Ed. by V. A. Legasov (Energoatomizdat, Moscow, 1984) [in Russian]. 2. V. I. Rachkov, “Nuclear power engineering as a factor of utmost importance for sustainable development of Russia in the 21st century,” Energosber. Vodopodg., No. 6, 2–4 (2006). 3. E. O. Adamov, A. V. Dzhalavyan, A. V. Lopatkin, N. A. Molokanov, E. V. Murav’ev, V. V. Orlov, S. G. Kalyakin, V. I. Rachkov, V. M. Troyanov, E. N. Avrorin, V. B. Ivanov, and R. M. Aleksakhin, “The conceptual statements in the strategy for develop ment of the Russian nuclear power industry in the out look up to 2100,” At. Energ. 112 (6), 319–330 (2012). 4. V. I. Rachkov, V. M. Poplavskii, A. M. Tsibulya, Yu. E. Bagdasarov, B. A. Vasil’ev, Yu. L. Kamanin, S. L. Osipov, N. G. Kuzavkov, V. N. Ershov and M. R. Ashirmetov, “The concept of a prospective power unit equipped with a BN1200 fastneutron sodium cooled reactor,” At. Energ. 108 (4), 201–205 (2010). 5. V. V. Alekseev, Yu. P. Kovalev, S. G. Kalyakin, F. A. Kozlov, V. Ya. Kumaev, A. S. Kondrat’ev, V. V. Matyukhin, E. P. Pirogov, G. P. Sergeev, A. P. Sorokin, and I. Yu. Torbenkova, “Sodium coolant purification systems for a nuclear power station equipped with a BN1200 reactor,” Therm. Eng. 60 (5), 311 (2013), DOI: 10.1134/S0040363613050019. 6. V. M. Poplavskii, A. N. Zabud’ko, E. E. Petrov, M. K. Ovcharenko, V. V. Popov, V. B. Bogush, N. I. Loginov, V. A. Tarasov, V. B. Polevoi, V. A. Khorom skii, and A. S. Mikheev, “Physical characteristics and problems associated with development of a sodium cooled fastneutron reactor as a source of highgrade thermal energy for producing hydrogen and for use in other hightemperature technologies,” At. Energ. 106 (3), 129–134 (2009). 7. F. A. Kozlov, L. G. Volchkov, E. K. Kuznetsov, and V. V. Matyukhin, LiquidMetal Coolants of Nuclear Power Installations: Purification from Impurities and Their Monitoring, Ed. by F. A. Kozlov (Energoatomiz dat, Moscow, 1983) [in Russian]. 8. A. C. Wittingam, “An equilibrium and kinetic study of the liquid sodiumhydrogen reaction and its relevance to sodiumwater leak detection in LMFBR systems,” J. Nucl. Mater. 60, 119–131 (1976). 9. A. V. Morozov and A. P. Sorokin, “Methods for obtain ing hydrogen and prospects of using a hightemperature sodiumcooled fastneutron reactor for its produc tion,” in Proceedings of the International Seminar “HighTemperature Projects,” Kalpakkam, India, November 14–15, 2011.

356

KOZLOV et al.

10. N. L. Plate and E. V. Slivinskii, Chemical Principles of Monomer Technology: A Handbook (IAPC Nauka/Inter periodika, Moscow, 2002) [in Russian]. 11. Yu. V. Chechetkin, V. D. Kizin, and V. I. Polyakov, Radiation Safety of an NPP Equipped with a Sodium Cooled FastNeutron Reactor (Energoatomizdat, Mos cow, 1983) [in Russian]. 12. D. A. Karpov, I. F. Kislov, S. N. Mazaev, V. Ya. Moiseev, A. P. Shabanov, G. V. Dubinin, V. A. Dubrovskii, and I. A. Khazov, “Studying the possibility of using protec tive coatings for reducing hydrogen penetration through the structural materials,” URL: http://www.rvs. itsoft.ru:8000/article/sart.html?id=239&conf_id=4. 13. F. A. Kozlov and L. G. Volchkov, “A method for regen erating the cold traps of impurities in alkali metals,” USSR Inventor’s Certificate No. 429468, Byul. Otkr., Izobr., No. 19 (1974). 14. F. A. Kozlov and L. G. Volchkov, “A new method for regenerating the traps for purifying alkali metal cool ants,” At. Energ. 39 (4), 310 (1975). 15. F. A. Kozlov, L. G. Volchkov, B. I. Tonov, Yu. P. Nali mov, and V. A. Likharev, “A method for regenerating the cold traps of impurities in alkali metal coolants,” USSR Inventor’s Certificate No. 473221, Byul. Otkr., Izobr., No. 21 (1975). 16. G. Eshbakh, U. Gross, K. Llayapcoeprep, and S. Shul lien, “Hydrogen diffusion and penetration through steel,” in Sorption Processes in Vacuum, Ed. by K. N. Myznikov (Atomizdat, Moscow, 1966) [in Russian].

17. J.L. Courouau and D. Labatut, “Hydrogen and tri tium distribution in FBRs. Modelling and crossover calculations in a reference case for the PHENIX and BN600 reactors,” Note Technique, No. 97036 dated 25.06.97 (Cadarache, 1997), pp. 17–20. 18. L. F. Belovodskii, V. K. Gaevoi, and V. I. Grish manovskii, Tritium (Energoatomizdat, Moscow, 1985) [in Russian]. 19. E. Fromm and E. Gebkhart, Gases and Carbon in Met als (Metallurgiya, Moscow, 1980) [in Russian]. 20. A. I. Lastov, E. A. Pavlinchuk, E. U. Konovalov, I. G. Sheinker, L. F. Lastova, I. A. Efimov, and A. N. Mezentsev, “Distribution of radionuclides over the primary circuit pipeline wall thickness in a BR10 reactor,” At. Energ. 60 (4), 262–264 (1986). 21. F. A. Kozlov and M. A. Konovalov, “Assessment of cesium release through the steel claddings in a high temperature sodium coolant,” in Proceedings of the Sci entificTechnical Conference “Thermal Physics2012,” GNTs RFFEI, Obninsk, 2012, pp. 115–116. 22. H. J. Matzke and G. Linker, “Study of the diffusion of cesium in stainless steel using ion beams,” J. Nucl. Mater. 64 (1–2), 130–138 (1977). 23. V. I. Isaichev, L. N. Larikov, and L. F. Chernaya, “About the diffusion mechanism of group IV elements in Mo and W,” Ukr. Fiz. Zh. 25 (4), 591–598 (1980). Translated by V. Filatov

THERMAL ENGINEERING

Vol. 61

No. 5

2014