L

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NEACRP-A- g52m/

Paper prepared for the 30th Meeting of the OECD/NEA Committee on Reactor Physics, Sept. 14-18, 1987, Helsinki, Finland.

Inherent Safety Performance of a Mixed-Oxide-Fueled 1000 MWe Loop-Type LMFBR under A'JWS Conditions

K. Yamaguchi

Power Reactor and Nuclear Fuel Development Corporation 4002 Narita, 0-arai, Ibaraki 311-13, Japan

S. Ohta

Nippon Atomic Industry Group Co., Ltd. 4-l Ukishima, Kawasaki, Kanagawa 210, Japan

and H. Endo

Toshiba Corporation 8 Shin-Sugita, Isogo, Yokohama, Kanagawa 235, Japan

ABSTRACT

A system dynamics analysis was applied to a 1000 MWe loop-type liquid-metal fast breeder reactor (LMFBR) with mixed oxide fuel to examine inherently safe, passive shutdown performance of the reactor under anticipated transients without scram (ATWS) conditions. The analysis included all of the reactivity feedbacks having been employed in current analyses of hypothetical core disruptive accidents (HCDAs). In addition, the present analysis stressed inherent responses

0 of the reactor system by including structural reactivity feedbacks due to axial expansion of control rod driveline (CRD) and radial expansion of the reactor core

a driven by the expansion of the core support plate (CSP). The ATWS initiator examined was an unprotected loss-of-flow (ULOF) accident caused by a site power blackout.

Some design options were considered: An upper-core flow chimney was introduced to make the CRD expansion feedback effective. The flow coastdown rate of primary pumps was selected to have a halving time of 10 seconds. The initial position of the control rods (CRs), the pony-motor flow level, and so on, were treated as parameters.

By assuming that the CRs were initially inserted into the active core by about 0.2 m and that a pony-motor has a capacity of driving more than 20 % of the rated flow, the ULOF consequence was mitigated and the peak sodium temperature was suppressed below boiling point. The CRD expansion feedback controlled the short-term transient, and the CSP expansion feedback became dominant in the later phase. The asymptotic temperature of the primary heat transport system against the upset was successfully lowered below 650 OC by optimizing the design parameters.

1

NOMENCLATURES

A : B :

reactivity feedback coefficient proportional to power change [Dollar] reactivity feedback coefficient proportional to the change in power-to-flow ratio [Dollar] reactivity feedback coefficient proportional to the change in inlet temperature [Dollar]

c :

CFp :

fl" :

'PP :

QDHR : q' : T : TF : t :

a B :

GTin : AT= : ATF : ;=CR :

P : ~11.2 :

specific heat [kJ/kg'C] normlized flow [X] pony-motor flow [%] thermal conductivity [W/m"C] prompt neutron life time [s] pressure [MPa], and, normalized power [%] IRACS heat removal amount [MW or %] linear heat rate of fuel pins [kW/m or %] temperature ["Cl fuel temperature ["Cl time [s] volume of fuel [m3] mass flow rate [kg/m*s] reactivity coefficient [Tdk/dT, Ak/kk',or Akfkk'/'C] delayed neutron precursor fraction [-I, and linear expansion coefficient [l/r&] inlet sodium temperature change from its nominal value ["Cl sodium temperature increase across the core ["Cl fuel temperature increase over sodium temperature ["Cl initial delayed density halving

control rod insertion length [m] neutron precursor decay constant [s-l] Pdm31, and reactivity [Dollar] time of primary flow coastdown [s]

INTRODUCTION

0 The traditional approach to reactor safety is based on defense in depth and

includes three levels of safety design. The first level emphasizes prevention of accidents with reliable equipments and intrinsically stable design features.

0 The second level of design focuses on protection against anticipated operat'ional occurrences which may arise as results of equipment failures, for instance. The third level of design is focused on design margins for extrem&ly unlikely and hypothetical events. Based on this approach, the designs of liquid-metal fast breeder reactors (LMFBRs) have included reliable engineered-safety-systems composed of redundant, diverse and independent reactor shutdown functions. Large thermal and structural margins have been added to the plant and containment to accomodate anticipated transients without scram (ATWSs) and/or hypothetical core disruptive accidents (HCDAS). Obviously, the high levels of reliability required bring significant penalties of increased complexity, cost, and difficulty in operatiofi and maintenance.

A fundamentally different approach to reactor safety is evolving with the development of inherently safe LMFBR designs. The approach to inherent safety has been directed toward the third level of safety design: The shutdown and subsequent heat removal have been accomplished by natural processes even if safety equipment in the first two levels of safety design had failed. An integral part of the strategy in eliminating an HCDA from design consideration

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0

is showing that the reactor is protected even for ATWSs. In broad term, it should be verified that the reactor plant acts to restore a proper balance between heat generation in the reactor and heat removal from it.

Unprotected loss-of-flow (ULOF) and other AlWS tests ware performed in the Rapsodie and EBR-II reactors(1)"(5). These tests implied the following feedback mechanisms: Thermal expansion of control rod drivelines (CRDs) caused the net insertion of control rods (CRs) into active core; Thermal expansion of core support plate (CSP) drove the radial core expansion; These phenomena provided overall negative reactivity feedbacks which could reduce the reactor power during the coolant heatup. and Coffield et a1.(7)

The ULOF evaluations by Chang et al.(2),(6) indicate that these two feedback mechanisms become

effective in different time phases: During the first phase, which extends over the initial 100 to 200 seconds of the accident, the CRD elongation contributes primarily to reducing the fission power down to an acceptable power level so that boiling and fuel damage are precluded. After this initial period, the CSP becomes hot enough to enhance radial core expansion. By this expansion, the reactor power continues to decrease toward the decay power level. The major concern left thereafter is assuring that acceptable temperatures are maintained by a decay heat removal system (DHRS) and a main heat transport system (MHTS), while issues of decay heat removal will not be further discussed here.

Extensive thermal-hydraulic tests of EBR-II validated the inherent safety approach. Supporting analyses(S)9(9) d' II-I lcated that the characteristics observed can be incorporated into large commercial liquid-metal-cooled reactors (LMRs) and be used as the basis for a totally new shutdown strategy. This conclusion was, however, obtained from the findings in a small metal-fueled reactor having negative or small positive sodium density and small Doppler coefficients. In such a reactor, the positive reactivity insertion by the sodium temperature increases is rather small. On the contrary, the sodium density coefficient is positive and the Doppler coefficient is very large in the cases of mixed-oxide- fueled commercial reactors. Therefore, the passive reactor shutdown capability available to a small metal-fueled reactor does not necessarily hold for large reactors with mixed-oxide fuel, and the enhancement of the passive shutdown

e mechanisms is required especially in the field of large LMFBR developments.

The purpose of this paper is to investigate the influence of several

0 significant design features on mitigating consequences of AlWSs. The work reported in the previous paper was extended and an additional sensitivity analysis was performed on a typical 1000 MWe loop-type LMFBR. Emphasis is laid on investigating the optimized design by which the short-term temperature overshoot is suppressed below boiling point and the long-term asymptotic temperature of the reactor system is lowered below 650 "C. The ATWS initiator examined here is the ULOF caused by the site power blackout.

REACTOR DESCRIPTION

The reactor plant cited in the present study.is not a specified but a typical loop-type IXFBR designed to operate at a thermal power output of 2600 MW and a net electrical power output of about 1000 MW. A schematic representation of the plant is given in Fig. 1. The plant has four MHTSs. Each system consists of a primary loop, a secondary loop and a water/steam loop. The rimary sodium at 360 'C is driven by the pump with a rated flow rate of 1.31~10 P kg/h. It enters ~the common cold plenum of the reactor vessel, goes up through the

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0

e 0

reactor core being heated, and reaches the common hot plenum where it mixes-with the bulk sodium and becomes 500 "C. Leaving the reactor vessel, it moves to the intermediate heat exchanger (IRX) and is cooled again to 360 'C. The IHX transfers the 650 MW energy generated in the reactor to the secondary sodium system and ultimately to the steam generator (SG) and the turbogenerator. The rated flow rate and the hot-leg/cold-leg sodium temperatures of the secondary circuit are 1.14~10~ kg/h and 4701310 "C, respectively. The secondary circuit has an intermediate reactor auxiliary cooling system (IRACS) as a DHRS.

The reactor vessel is 8 m in diameter and 15 m in height, containing a core of 6 m in diameter and 4 m in height. The reactor core is composed of 354 oxide-fueled driver subassemblies, 25 CRs, 150 radial blanket and 372 radial shield subassemblies. The active core is 1 m in height and is radially subdivided into inner and outer cores whose diameters are 2.5 and 3.3 m, respectively. The lengths of the upper and lower axial blankets are 0.35 m. The fuel subassembly, spaced triangularly with a radial pitch of 160.8 mm, contains a 271-pin, wire-spacer bundle having around 0.35 MPa pressure drop at the nominal operating condition. The diameter of fuel pin is 7.4 mm and the maximum heat flux is 43 kW/m.

ANALYTICAL MODEL

A flow network analysis code used in the previous study was applied again to the present study. The code models flow mixing in the plenum, heat transport along the flow paths and heat transfer between sodium and structures. It has a capability of treating transient thermal-hydraulics with a fine node scheme.

Figure 2 shows the flow network model applied to the reactor vessel. It is assumed that 90 % of rated flow entering the reactor vessel moves straight to high pressure plenum located just below the CSP, and that residual 10 % of the primary flow mixes with the bulk sodium at the lowest cold pool and joins thereafter to the main flow at the high pressure plenum. The high pressure plenum supplies the fuel subassembly flow, while some part of the sodium flow branches into the low pressure plenum and then separates into radial blankets and neutron shields depending on the pressure drop characteristics of these subassemblies and structures.

The hot sodium introduced into the upper plenum during the flow coastdown rises toward the top of the plenum, thereby providing the thermal expansion of the CRD. To enhance this heatup-induced elongation of the driveline, a design of an upper core flow chimney was considered in the flow network model. The entrance of the chimney is located to include coolant stream from the CR and its neighboring six fuel subassemblies, and encompasses each CRD in the upper plenum. The effect of the chimney on the flow distribution is, however, sensitive to design details. Therefore, it is assumed~from the geometrical configuration of the chimney that 40 % of sodium flow from the core passes through the chimneys.

The thermal-hydraulics of the reactor core is represented by the single-pin model usually used in one-dimensional safety analysis codes. Figure 3 illustrates the axial and radial node divisions. The core is described by seven channels to represent the regions of inner core, outer core, CR, fuel neighboring with CR, radial blanket, radial shield, and hottest channel having maximum power-to-flow ratio in steady state. All of these subassemblies are tightly supported by the CSP, thereby moving radially together with the CSP which expands

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radially during thermal upsets. The radial bowing of those subassemblies and the inter-subassembly flow- and heat-redistribution, both of which may be very important inherent safety features, are not taken into account. As discussed in the previous st"dy(lO), we can conclude that the present model would lead to conservative results. Furthermore, the calculation is simplified conservatively by neglecting the axial displacement of the active core, which might be caused by axial expansions of core support structures, hexcans and lower region of the pin itself.

The feedback from the expansion of the reactor vessel is positive. Following increase in outlet coolant temperature, the vessel will eventually expand and, as a result, till effectively lower the reactor core away from the CRs that are suspended from the upper deck structure. Ihe relative movement introduces additional reactivity that will drive the reactor toward higher power. However, the expansion of the reactor vessel is greatly delayed since thermal insulators are equipped around the reactor vessel. The reactor vessel expansion is, therefore, tentatively ignored in the present study.

The neutronic behavior is described by point kinetics with six groups of delayed neutron precursors given in Table 1. The reactivity coefficients are evaluated as shown in Table 2. It was pointed out that the transient fuel behavior has a possibility of introducing either of negative and positive feedbacks depending on whether or not the irradiated fuel pellets stick to the cladding on different points on their circumference along fuel column height(l). The non-fuel-stick behavior is assumed conservatively in the present calc"lations.

RESULTS OF PARAMETRIC STUDY

Calculational conditions

The dominant design parameters and the analytical models are fixed as described above. The other parameters which shall be identified as desirable from the standpoint of enhancing the inherent safety are selected by the

0 following criteria:

0 (1) The short-term temperature overshoot should be suppressed below sodium boiling point;

(2) The long-term asymptotic temperature of the reator system should be lowered below 650 OC;

(3) Designs specified by the selected parameters should have feasibility of development and flexibility for modification.

The reference case condition was determined as shown in Table 3. Of primary importance are the selections of the insertion length of CRs, AzCR, and the specifications of the primary pump, i.e., the halving time of flow coastdown, 'I/2, and pony-motor flow level, FPM. The combination of these parameters was altered from that used in the previous study because of the difficulty in pump development, viz., AzCR = 0 + 0.2 m, ~1/2 = 60 + 10 s, and FPM = 10 -f 30 % (FPM will be revised later to 20 % in the optional design case).

Table 3 also lists calculational conditions for other cases. Among 25 cases

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calculated, 2.4 cases excluding the REF case belong to the sensitivity study where effects of various parameters on short-term peak temperatures were chiefly examined by comparing the results with the standard one from REF case. The calculations of these cases were terminated at 500 seconds aftertransients, although the REF case "as calculated up to 3 hours (10,800 seconds) to confirm the steady-state conditions.

The REF case is a standard one obtained from optimizing thermal-hydraulic conditions while nuclear-kinetic parameters are unchanged from premised values. Understanding the plant response characteristics by the REF and 24 cases of short-term survey calculations, we will discuss further advantageous optional designs in the latter part of this paper. Our discussion will be focused on examining the possibility of design modification especially in the field of the core and fuel design, namely, the examination will be concentrated on linear heat rate of fuel pins, thermal conductivity of fuel materials, and reactivity coefficients of Doppler and sodium density.

0 Typical ULOF consequence

Figures 4 and 5 show the result of the REF calculation, "here Fig. 5 draws the entire results of 0 - 3 hours (10,800 seconds) while Fig. 4 focuses on only initial transient of 0 - 500 seconds. The ULOF consequence caused by the site power blackout begins with the primary and secondary pumps run down, terminating the balance of MHTS heat rejection and inducing the automatic pony-motor operations associated with the IRACS heat removal.

The flow coastdown leads to a steep increase in the sodium temperature at the hottest channel as well as at the other fuel subassemblies. Figure 4(a) shows that the initial temperature increase is so fast that the peak temperature of about 840 'C (an increase of 285 'C over the nominal exit temperature of hottest channel) is reached at 30 s after flow coastdown. The power-to-flow ratio (power/flow) takes peak value of 2.2 at 22 s and falls gradually thereafter. The flow coastdown terminates at 25 s, and then the pony-motor is switched on. Over the initial heatup phase of 30 s, the net reactivity subsides to -20 6 level and the reactor power decreases to around 50 % of rated power

0 level.

0 me can see from Fig. 4(b) that the temperatures at the upper plenum also

increase with the same trend as those of fuel subassemblies, Tout, although there is some transport delay of hot sodium from core to upper plenum. The rising hot sodium within the chimney (T35) causes promptly the heatup of CRD (T45), introducing the negative reactivity feedback after about 5 s into the transient. Figures 4(c) and 4(d) show the reactivity balance and the power and flow histories. At first, the Doppler reactivity is negative, because the sodium temperature increase provides an increase in fuel temperatures associated with the short-term heat accumulation within the fuel pins. The peak fuel temperature reached at 7 s was about 2300 'C. However, the subsequent power reduction caused by the CRD elongation and the other negative reactivity feedbacks leads to a decrease in fuel temperatures, thereby making the Doppler feedback positive after 18 s. The fuel expansion feedback shows the same tendency while the magnitude is relatively small. The CRD elongation feedback is a major negative reactivity component during this earlier phase. The steel (clad and hexcan) expansion reactivity is positive but is negligibly small. The CSP expansion feedback plays no role in this phase due to a transport delay of heated sodium flow toward the CSP and a large heat capacity of the CSP.

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The power reduction by the negative reactivity feedbacks and the flow reduction by the primary pump run down occur at the same time to yield a peak in Tout when the power/flow takes maximum value. The hottest sodium reaches the CRD at 41 s, thereby giving the largest negative CRD expansion feedback into the core. A moderate transient follows thereafter, because the increasing trend of To"t overshoot at the early phase diminishes, and T,"t becomes oscillatory at a high temperature level, while the oscillation disappears as time goes on. After 100 s into transient, the negative CSP expansion feedback begins to grow up due to the gradual heatup of sodium and structures in the lower plenum. The power/flow falls below unity after 273 s and the temperature increases across individual subassemblies, AT,, become smaller than their initial increases. As a result, the temperature distribution in the whole plant tends to be uniform. The status at 500 s is summarized in Table 4, where it is shown that the difference in sodium temperatures at the fuel and blanket subassemblies is around 50 "C at most. The fuel temperatures become close to sodium temperatures since the reactor power falls down to 5.5 % including the decay heat of 1.4 %.

After several hundred seconds, the CSP expansion feedback becomes a major

0 component to reduce the reactor power. Figure 5(b) shows that the sodium temperature at the CSP (T20) takes peak value at about 600 s. However, the CSP temperature (T108) does increase very slowly due to the large heat capacity of the CSP. The slow heatup of the CSP and the associated growth of negative feedback by CSP expansion continue over 1300 s. The reactor power falls to decay power level at this stage.

once the heat source becomes small enough to settle the entire temperature field, the system response becomes being dominated by the heat balance of the whole plant. When the IRACS is in operation just like in the present case, the over-cooling by the IRACS lowers the cold-leg sodium temperature to lead to a partial reduction in the negative CSP feedback. Then, the reactor power increases gradually to match with the amount of the IRACS heat removal. This phenomenon introduced long-term perturbations in power and temperature histories. Figure 5 reveals that a complete quenching condition is established after 3,000 seconds into transient at a temperature level of around 600 - 630 'C.

8 Parameter sensitivity

0 The results of survey calculations relevant to short-term Tout overshoots of hottest channel are summarized in Table 5, where sensitivities of temperature changes against one per cent increases in individual parameters are compared among all cases.

It was pointed out in the previous study (10) that the primary flow coastdown rate gives a significant effect on peak sodium temperature. The extended flow coastdown contributes to suppress not only the Tout overshoot of hottest channel but also the radial temperature difference during the earlier transient. This heat removal function must be substituted by the pony-motor operation for the fast flow coastdown, while this function is expected to be performed by the primary pump in the case of the prolonged flow coastdown. The calc"lational results of REF and PO2 cases surely,showed that the above requirement was partly covered by the high-flow pony-motor operation, namely, the pony-motor operation could apparently terminate the rising trends of sodium temperatures in the core

.to yield peaks in 5 to 10 s after pony-motor operation. However, sodium boiling occurred when the pony-motor flow level was lowered to 10 % level (PO1 case). The peak temperature increases with the rate of 7.4 OC per 1 % reduction of

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~pony-motor flow at around 20 % level.

The initial CR insertion into the active core gives large effects on sodium temperature behavior. This means that the released energy over the entire transient becomes small as the CR is inserted partially beforehand. The results of CPl and CP2 cases suggested that the peak temperature could be reduced by 2.1 "C par 0.01 m insertion of CRs at AzCR = 0 to 0.2 m region. The similar effect will be obtained by improving the design of the chimney and CRD structures to have a 10 %-larger capability of expanding the CRD than the nominal value.

The examination on the CSP design was conducted in the previous paper: The time constant of the CSP expansion was so large that the CSP feedback was insensitive to the sodium peak temperature especially in the fast flow coastdown case. Additionally, it was confirmed in the present study (SCl, SC2 and HTl case) that the operational conditions of the secondary loop and IRACS, though their conditions were related to the core inlet temperature, also played little role in mitigating the earlier transient.

The uncertainties in the reactivity coefficients are of primary concern in the safety analyses. The feedback coefficients were, therefore, altered by factors 1.2 and 0.8 in the RDPl, RDP2 (Doppler), RCCl, RCC2 (sodium), RFLl, RFL2 (fuel), RLDl, RLD2 (steel), RCRl, RCR2 (CRD), RCSl and RCS2 (CSP) cases, and the ULOF consequences calculated were compared with that of the REF case. The excess temperatures beyond the reference value at the peak condition are converted into values per 1 % increases in individual coefficients, and are shown in Table 5 together with those evaluated at 500 s into ULOF transients. The sodium density coefficient is most sensitive to the peak temperature among all positive feedbacks, and is the CRD worth among all negative feedbacks. Therefore, much efforts should be concentrated on minimizing these errors in the core de$igfi.

0

Effects of reducing the fuel temperatures on mitigating the earlier transient were examined by changing the linear heat rate of fuel pins and thermal conductivity of fuel materials in FPl (0.5xq'), FP2 (0.75xq') and FCl (5xk) cases (the subassembly heat output was conserved by increasing the number of

0 pins contained in the subassembly in FPl and FP2 cases). The results are summarized in Table 6. The positive Doppler feedback is lowered drastically as compared with that of REF case, since the Doppler feedback is given by

0 aDln(TF/TF,o)~and In(TF/TF o) becomes small for a Small initial fuel temperature condition. In addition, the initially stored energy in the fuel pins, which is given by c~~VTF,~, becomes small. These two factors introduce a larger and negative net reactivity during the earlier heatup phase, and therefore leads to a situation of smaller and shorter temperature overshoot than the REF case.

DESIRED TRENDS IN SAFETY DESIGN

Quasi-static reactivity balance

Based on the following quasi-static reactivity balance equation, Planchon et a1.(8) and Wade(g) analized the temperature reponse characteristics:

6p = A&P + BG(P/F) + C&Tin (1)

where a change in reactivity, 6p, is correlated with changes in power, 6P, power-to-flow ratio, 6(P/F), and inlet temperature, GTin, by the above lumped

8

equation using the coefficients A, B and C. The power coefficient, A, includes reactivity feedbacks proportional to power change, and is expressed by:

A = -(aD + oF)ATF = -(Doppler + Fuel Axial Expansion)

x (Ave. Fuel - Ave. Sodium Temperature) (2)

The power/flow coefficient, B, includes feedbacks which are proportional to the power-to-flow ratio or the flowing sodium temperature rise across the core. It is given by:

B = -(aD + CXF + oRa + oS + oCRD)AT,/Z = -(Doppler + Fuel Axial Expansion + Sodium density

+ Steel density + CRD Expansion) x (Ave. Sodium - Inlet Sodium Temperature) (3)

The inlet temperature coefficient, C, includes feedbacks proportional to inlet temperature change. It is formulated by:

0 c = -(‘XD + OF + a)Ja + OCSp) = -(Doppler + Fuel Axial Expansion + Sodium Density

+ CSP Expansion) (4)

Using these equations, we can examine the desirable relationships among the reactivity coefficients A, B and C for the ULOF event in question.

In addition to the above quasi-static discussion, we can point out the following for the short-term temperature overshoot: Since the flow coastdown time constant, rl/z, is smaller than the delayed neutron time constant, l/X, the power reduction induced by the negative reactivity of power/flow increase, B6(P/F), is held back by delayed neutrons, and the power cannot decrease quickly. The power/flow mismatch is thus exacerbated, causing the Tout overshoot. The sequential change from the primary pump flow coastdown to the pony-motor operation alters the tendency of the Tout overshoot and suppresses the unlimited temperature increase. The Tout overshoot disappears after

0 several delayed neutron lifetimes as the quenching condition, 6p = 0, establishes itself. Although the Tout overshoot relative to its asymptotic value cannot be

0 quantified without dynamic analysis, Planchon et al. and Wade pointed out that the relative overshoot is reduced when:

T~/~X(A + B)*/B >> 1 . (5)

This shows that one should extend the pump coastdown and/or increase the power coefficient relative to the power/flow coefficient.

For the end state of the ULOF transient, Eq. (1) leads to

0 = A(P - 1) + B(P/F - 1) + CGTin

Therefore, the quenching~temperature, &Tin, is given by:

6Tin = (A + B)/C - [AP + B(P/F)]/C (6)

Substituting the REF calculational conditions, i.e., P = 0.046 and F = 0.3, into Eq* (6), we can obtain:

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GTin = (O-9548 + 0.847B)/C

0

0

Tentatively assuming the typical coefficient values~of A = -(0.25 - 2.0) $ B = -(0.25 - 0.75) $, and C = -0.005 $/"C(g), th e quenching temperature becomes &Tin = 90 - 510 'C, i.e., Tin = 450 - 870 "C. Apparently, the desired trends in reactivity coefficients A; B and C are:

Power coefficient A small and negative Power/flow coefficient B small and negative Inlet temperature coefficient C large and negative

Partial design modification

The above discussion suggests that A + B should be large for limiting short-term Tout overshoot, but should be small for suppressing the asymptotic temperature. This conflicts can be solved in favor of the long-term objective, because the effort of suppressing the short-term Tout overshoot has been taken by the design for long T and high Fp~*. For a long-term coolant behavior, the desired trends in reactivity coefficients are available from Eqs. (2), (3) and (4). One can also use the result of the parametric survey that improvements of fuel characteristics make consequences of the ULOF event desirable. To investigate the advantage of choosing more desirable parameters than the standard design examined in the REF calculation, one additional case (LNG case) "as selected and the entire plant dynamics calculation "as conducted.

The calculational condition of the LNG case shown in Table 7 "as selected as follo"s: ATf appears only in A among Eqs. (Z), (3) and (4), and can be used to design for a low value of A. The Doppler coefficient and the linear heat rate of the fuel pin "era, therefore, reduced by 20 % from their nominal values, namely, the power coefficient, A, "as lowered to about 65 % level in LNG case as compared with that of REF case. The positive sodium density coefficient "as also reduced by 20 % to obtain better trends in coefficients B and C. Physically interpreted, this modification is almost equivalent to selecting a design option of smaller height-to-diameter-ratio reactor core having large neutron leakage and hardened neutron spectrum. The pony-motor flow "as reduced simultaniously from 30 to 20 % level from the point of MHTS design availability.

0 Modified ULOF consequence

Figure 6 shows the result of the LNG calculation. The peak temperature calculated "as 877 "C (37 'C higher than the RRF case) and it "as reached at 44 seconds of transient. The decreased heat removal capability of coolant due to the reduced pony-motor flow level "as nothing to worry about. The reduction in stored energy in fuel pins was so superior to all counter effects that the period of temperature overshoot beyond 650 "C "as extremely shortened in the hot-leg components, i.e., REF + LNG = 730 + 500 s at the upper core structure (UCS), and even disappeared in the cold-leg components.

The power reduction continued until about 1,000 s. The minimum power level "as halved compared with that of RBF case. Therefore, over-cooling by the IRACS occurred more strongly than the REF case, leading to larger-amplitude temperature perturbations during 1,000 to 3,000 s period. The mismatch between heating and cooling conditions disappeared to yield a complete quenching condition at about one hour after flow coastdo"n. The attained temperature level "as 550 - 600 'C. This level "as lower than the REF case by 36 OC, typically.

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It is evident that a large part of the reason of achieving an advantageous quenching condition comes from suppressing the power coefficient, A, which will be easily attained for metal fueled reactors. However, the present study validated that the design of an inherently safe, mixed-oxide-fueled commercial LMFBR will be available at least within the framework of loop-type reactor designs.

CONCLUSION

A system dynamics analysis was applied to a 1000 MWe loop-type LMFBR with mixed-oxide fuel to examine inherently safe, passive shutdown performance of the reactor under AlWS conditions. The analysis included all the reactivity feedbacks employbd in the current analysis of HCDA. In addition, the present analysis stressed inherent responses of the reactor system by including structural reactivity feedbacks due to axial expansion of CRD and radial expansion of CSP. The analytical results were summarised and the following conclusions were drawn:

(1) A chimney type UCS was considered in the analysis to introduce effective CRD expansion feedback. With this design, the ULOF accident caused by the site power blackout was mitigated and sodium temperatures were suppressed below boiling point by using the primary pump having a 10 s halving time of flow coastdown aided by the pony-motor having a capacity of driving more than 20 % of rated flow and by assuming that the CR was initially inserted into the active core by about 0.2 m. It was found out that the CRD expansion feedback and the flow coastdown characteristics controlled the earlier thermal upset, and that the CSP expansion feedback and the cooling condition of DHRS dominated the long-term transient.

(2) Desired trends in safety design especially in the field of nuclear-kinetics were discussed. A smaller aspect-ratio reactor core gives better trends in reactivity coefficients. The asymptotic temperature can be suppressed below 650 "C with sufficient safety margin when thermal-hydraulic and nuclear-kinetic parameters are well optimised in the plant design.

ACKNOWLEDGMENTS

The authors would like to thank Mr. A. Oiwa, Dr. T. Aoki, and Dr. K. Shiba for their suggestions and advice. This study was conducted under the sponsorship of Plant Engineering Section of PNC. The authors are indebted to the project manager, Y. Taniyama.

(1) C. Essig et al., "Dynamic Behaviour of Rapsodie in Exceptional Transient Experiments," International Topical Meeting on Fast Reactor Safety, cONF-850410, Knoxville, TN, Apr. 1985.

(2) L. K. Chang et al., "An Experimental and Analytical Investigation of Unprotected Loss-of-Flow Events in EBR-II," Trans. Am. Nucl. Sot., Vol 50, pp. 350-351, San Francisco, CA, Nov. 1985.

(3) G. H. Golden et al., "Evolution of Thermal-Hydraulics Testing in EBR-II," Nucl. Eng. Des., Vol 101, pp. 3-12, Apr. 1987.

11

(4)

(5)

(6)

(7)

(8)

(9)

0 (10)

D. Mohr et al., "Loss-of-Primary-Flow-without-Scram Tests: Pretest Predictions and Preliminary Results," Nucl. Eng. Des., vol 101, pp. 45-56 Apr. 1987.

,

E. E. Fedlman et al., "EBR-II Unprotected Loss-of-Heat-Sink Predictions and Preliminary Test Results," Nucl. Eng. Des., Vol 101, pp. 57-66, Apr. 1987. i; K. Chang and D. Mohr, "The Effect of Primarv Pumu Coastdown Ch laracteristics on Unprotected Loss-of-Flow Transients in EBR-II w Second swcialists' Meeting on Decay Heat Removal and Natural Convectio: in

3 Island, NY, Apr. 1985. L.MFBRs, J.0nj R. D. Coffield et al., "Thermal-Hydraulic and Nuclear Kinetic Considerations Involved with LMR Inherent Safety," Second International Topical Meeting on Nuclear Power Plant Thermal Hydraulics and Operations, Tokyo, Japan, Apr. 1986. H. P. Planchon et al., "Implications of the EBR-II Inherent Safety Demonstration Test," Nucl. Eng. Des.,, Vol 101, pp. 75-90, Apr. 1987. D. C. Wade, "LMR Core Design for Inherent Safety," 29th Meeting of the OECD/NEACRP, Chalk River, Canada, Sept. 1986. K. Yamaguchi and H. Endo, "System Dynamics Analysis for Mitigating ATWS Consequences of a 1000 MWe Loop-Type LMFBR," 29th Meeting of the OECD/NBACRP, Chalk River, Canada, Sept. 1986.

Table 1 Neutron kinetics parameters

Parameter Value

Effective delayed neutron precursor fractions (Beff) 3.67~10~~

Six-group components: 61 7.76~10-~

a 62 8.13~16~ Xl 1.29x10-2

x2 3.11x10-2 03 6.56~10-~ 84 1.33x10-3

x3 1.34x10-1

0 x4 3.31x10-1

85 6.21~10-~ 1.26 S6 1.74x10-4

x5 ?'6 3.21

Prompt neutron lifetime (lp) 0.45 US

Table 2 Temperature coefficients of the reactivities

Reactivity coefficient Value

Doppler (Tdk/dT) -1.07x10-2 Sodium (Ak/kk') +1.87x10-2 Fuel (Akfkk') +3.60x10-l Steel (Aklkk') -6.63~10-~ CRD (Ak/kk') -1.14x10-1 CSP (Ak/kk'/"C) -1.15x10-5

12

.

Table 3 Calculational conditions of short-term ULOF consequences

Category

Reference

Pony-motor flow level

0 IRACS heat removal

CR position

CRD expansion

Linear heat rate

Thermal conductivity

0 Reactivity Coefficient

0

Case ID

REF

PO1 PO2 SC1 SC2

HTl

CPl CP2

CEl CE2

FPl FP2

FCl

RDPl RDP2 RCCl RCC2 RFLl RFL2 RLDl RLD2 RCRl RCR2 RCSl RCS2

Calculational conditions

4-100~ primary- and secondary-pump trip Primary: Secondary:

~1/2 = 10 S, FPM = 30 % 5.5 s, 7.5 %

CR partial insertion: AZCR = 0.2 m IRACS heat removal: QDHR = 30 MWfloop (= 4.6 X) Linear heat rate: 4' = 43 kW/m (= 100 %) Thermal conductivity: k = nominal CRD expansion: S = nominal Reactivity coefficients given in Table 2

Primary FPM = 10 % 20 %

Secondary FPM = 5 % (Natural,circulation flow) 10 %

QDHR = 15 MW/loop (= 2.3 %)

AzCR = 0 m (Full extraction) 0.1 m (Partial insertion)

2xBcRD (Doubled CRD expansion coefficient) 1?5X&RD (Intermediate case)

0.5xq' (Halved heat flux) 0.75xq' (Intermediate case)

5xk (Intermediate case between metal and mixed-oxide fuels)

1.2xDoppler O.BxDoppler 1.2xSodium 0.8xSodium 1.2xFuel 0.8xFuel l.ZxSteel O.BxSteel 1.2xCRD O.BxCRD 1.2XCSP O.BxCSP

13

Item Value

power: . Fission . Decay heat . Total

Table 4 Status of variables at 500 s into ULOF transient

Item Value

Reactivity balance: 4.1 % . Doppler 65.9 6 1.4 . Sodium 25.5 5.5 . Fuel 5.0

. Steel 1.1

. CRD -61.1 . CSP -48.3 . Total -11.9

Sodium/fuel temperature distribution: . Hottest channel 7141785 OC . Inner core 7oai781 . Outer core 6991744 . Radial blanket 6621678

Table 5 Summary of the parametric survey calculations

0 category

Reference REF 31.0 840.0

Pony-motor flow level . Primary loop

. Secondary loop

PO1 PO2 SC1 SC2

50.0 boiling ----- 44.5 913.8 -7.38 31.0 840.0 31.0 840.0 0

IRACS heat removal

CR position

HTl 31.0 840.0 0

0 CRD expansion

0 Linear heat rate

Thermal conductivity

Reactivity Coefficient Doppler

Sodiuin

Fuel

CPl CP2 CEl CE2

34.0 882.6 -2.13 33.0 870.8 ("C/cm) 28.0 789.2 29.5 816.8 -0.25

FPl FP2 FCl

36.0 795.8 32.5 820.3 M.88 29.5 795.5 -0.09

Steel

CRLI

CSP

RDPl RDP2 RCCl RCC2 RFLl RFL2 RLDl RLD2 RCRl RCR2 RCSl RCS2

31.5 843.4 30.0 836.6 +0.17 31.0 845.4 31.0 834.8 +0.27 31.0 840.3 31.0 839.7 +0.02 31.0 840.3 31.0 839.8 +0.01 30.0 825.2 32.0 856.8 -0.79 31.0 840.0 31.0 840.0 0

Case ID Peak conditions

Time Temp. Sens . (s) ("C) ("C/%)

14

status at 500 s Temp. Sens.

("C) ("C/%)

713.6

----- -----

748.2 -3.46 712.6

'713.9 -0.26

716.8 -0.05 ("c/Mw)

794.9 -4.06 762.0 ("C/cm) 660.1 687.7 -0.27

607.7 662.1 +2.12 629.1 -0.17

730.6 693.6 +0.93 721.4 706.2 +0.38 715.0 712.1 +0.07 713.9 713.2 +0.02 696.4 735.2 -0.97 702.4 726.0 -0.59

0

Table 6 Effects of the decrease in fuel temperatures

Case ID FPl FP2 FCl REF

Calculational conditions 0.5xq' 0.75xq' 5xk

Initial temperature ("C) Hottest fuel center Hottest fuel surface

1,340 717

1,829 1,216 839 960

Time of power/flow > 1 (s) 185 210 185

Peak temperature conditions Sodium peak temperature ("C) 796 Fuel center peak temperature ("C) 1,356

820 796 840 1,843 1,241 2,283

Conditions at 500 seconds Hottest sodium temperature ("C) 608 Hottest fuel center temperature ("C) 617

surface temperature ("C) Doppler reactivity (6) 25.3 Sodium reactivity (6) 15.9 CRD reactivity (8) -26.7 CSP reactivity (C) -38.1 Total reactivity (C) -19.4

662 694

46.5 20.9

-43.5 -43.6 -14.6

629 714 639 785 633 722

35.5 65.9 18.0 25.5

-33.2 -61.1 -40.5 -48.3 -17.4 -11.9

Table 7 Calculational conditions of long-term ULOF

Item Calculational conditions REF LNG - -

Pump flow coastdown characteristic

0 primary: ~1/2 = 10 S, FPM = 30 % secondary: 5.5 s, 7.5 %

0 CR position AzCR = 0.2 m

IRACS heat removal QDHR = 30 MW/loop (= 4.6 %)

Power density q ' = 43 kW/m (= 100 %)

Thermal cond. k = nominal

CRD expansion S = nominal

Reactivity coefficients Doppler aD = nominal Sodium cr~~ = nominal Fuel aF = nominal Steel aS = nominal CRD oCRD = nominal CSP aCSp = nominal

15

9’9 k

2,276 960

237

consequences

'l/2 = 10 s, Fp,., = 20 % 5.5 s, 7.5 %

AzCR = 0.2 m

QDHR = 30 Mw/100p (= 4.6 %)

4 ' = 34.4 kW/m (= 80 %)

k = nominal

S = nominal

a,, = 80 % aNa = 80 % a* = nominal as = nominal OCRD = nominal uCSP = nominal

l e

n Air

__---- --- Sodium

e

Pump Gl

ml- 1~ of 4 loops 3lfJ”C, i.l4XlO’kg/h

I--------

Pig. 1 Schematic flow diagram of a 1000 MWe loop-type LMFBR

. ’

a l

/TIM MT105

CRD chimney

Fuel CR Fuel SA SA SA

Detail of the UCS and chimney

Fig. 2 Flow network model of the reactor vessel

III I l!!l II E J=MTOP

CR M (Inter-assembly

heat transfer) - ’

I II

&dial noding

Inter-assembly flow redistribution

Fig. 3 Single-pin, multi-channel model

17

00

8o01 C~D(TW. CRD(T‘W Outlet(T12,) t- z*

400 fROL Sodium ---structure

I 1 I I 0

I 100 200 300 400 5m-l

Time (s) b) Sodium and structure temperatures

at the reactor vessel

010 Time (s)

a) Sodium temperatures at subassembly exits

-0.6 FD -0.23; ’ I I I I I I I

100 200 300 400 500 Time (s)

- c) Reactivity balance

Fig. 4 Predicted short-term ULOF consequence of the REF case

0

N” I=

E 0.01 I I I f 4 I I I 100 200

2 300 400 500

Time (s) d) Power, flow and power/flow

0.6 -

.

I

CRD(T35). CRD(T45)

Inlel(T155)

CSP(TlO6) -Sodium --- S1ruc1ure

r I I I 0 I I I I I I I 0 2000 4000 6000 8000 10000

Time (s) b) Sodium and structure temperature

at the reactor vessel

“““I 800 -

Sodium /

Fuel /

/

Radial blanket

500.

400 # I I I I I I I I I

0 2000 4000 6000 8000 10000 -~ I I I 1 I I I I 1 I I 0 2000 4000 6000 8000 10000

Time (s) Time (s)

a) Sodium temperatures at subassembly exits

c) Reactivity balance

Fig. 5 Predicted long-term ULOF conseqwnce of the REF case

‘I’ 1 i _------------------_----------- k I I ,FlOW FIOW 4

0.01; ' I I I I I I I I I 2000 4000 6000 8000 10000

Time (s)

d) Power, flow and power/flow 1.2 I I I I I I I I

/ Doppler ,-

l 9. .

. .

L CRD(T35), CRD(T45)

-Sodium ---SWucture

I I I I I I 0 2000 4000 6000 8000 10000

Time (s) b) Sodium and structure temperature

at the reactor vessel

800 - I,

Hottest channel

0 2000 4000 6000 8000 10000 Time (s)

a) Sodium temperatures at subassembly exits

Time (s) d) Power, flow and power/flow

0.8 Doppler

-0.8 I I , I I I I 0 1

0 2000 4000 6000 8000 10000 Time (s)

c) Reactivity balance

Fig. 6 Predicted long-term ULOF consequence of a modified case by desired design parameters