Reactivity feedback

One is the secondary- coolant reduction test by partial secondary loss of coolant flow in order to simulate the load change of the nuclear heat utilization system. This test will demonstrate that the both of negative reactivity feedback effect and the reactor power control system brings the reactor power safely to a stable level without a reactor scram, and that the temperature transient of the reactor core is slow in a decrease of the secondary coolant flow rate. The test will be perfonned at a rated operation and parallel-loaded operation mode. The maximum reactor power during the test will limit within 30 MW (100%). In this test, the rotation rate of the secondary helium circulator will be changed to simulate a temperature transient of the heat utilisation system in addition to cutting off the reactor-inlet temperature control system. This test will be performed under anticipated transients without reactor scram (ATWS). 

LMRs with oxide-fueled core Models modified and newly developed mto the code so far mclude models for reactivity feedback effects and pool thermal-hydraulics In order to venfy the logic of the models developed, and to assess the effectiveness of the inherent safety features based upon the negative reactivity feedbacks m achieving the safety design objectives of passive safety, a preliminary analysis of UTOP and ULOF/LOHS performance has been attempted... 

It is essential for the safety of the reactor to exclude the possibility of super prompt critical state at all times. This requires that the inserted reactivity at potential events should be below 1 under conservative conditions, neglecting reactivity feedback coefficients. 

Table 6.3 illustrates the magnitudes of various reactivity feedback coefficients and their variation with fuel burnup. 

The turbulent flow control can be implemented through different control schemes [1]. They are predetermined open-loop control, reactive feedforward open-loop control, and reactive feedback closed-loop control. The physical arrangement of sensors and actuators depends on the nature of control schemes and flow characteristics. The criteria for selection of appropriate actuators have been discussed in next section. It should be noted that the actuators require parasitic power supply for its operation. This fact should be taken into consideration for overall performance analysis of actuators. 

Depending on the core conditions (power level), the reactivity feedback coefficient due to the coolant density may either be negative or slightly positive. Ingress of gas into the core is likely during injection by the hydroaccumulators. This also adds to the coolant density variations, which are bound to affect the reactivity. 

Large excursions. We next mention briefly the third class of reactor kinetics problem referred to earlier, i.e., the large excursion problem involved in safety analyses. We illustrate the problem by means of a very oversimplified example. We let w(t) represent the cumulative energy release in the reactor, and assume that the reactivity feedback is just proportional to this energy release, i.e.,... 

In these equations n(r) is the neutron density, c (r) the density of precursors, d the fractional production of precursors per neutron produced in the core. The symbols G and P are linear integral operators, the operator G denoting absorption and transport processes while P denotes production by fission. We will assume that all the reactivity feedback takes place through the operator G, although this assumption is not essential. This assumption may be expressed in the form ... 

It would thus be necessary to measure reactivity feedback for all possible pairs of frequencies o>i = — isi, o>2 = 52, including both sum and difference components. This represents a formidable experimental problem because the effect due to feedback alone would have to be separated from other sum and difference frequencies in the reactor response arising from terms of the form kc(t)(f> t) in equation (14). Thus the simple separation of the feedback by means of equation (13) would no longer be possible. This problem has not been fully investigated. 

This equation completely determines the reactivity feedback in terms of the... 

The general expression for the disassembly reactivity feedback derived in Appendix Al, Eq. (A 1.23), is... 

The disassembly reactivity feedback for a spherical reactor core is derived in Appendix Al, Eq. (A1.35). However, this expression presents some problems when trying to analyze a pancaked cylindrical core, since it is not clear what the one-group buckling, j8, should be in the equivalent spherical core. This difficulty is circumvented by assuming V Fin Eq. (A1.23) is given by... 

This approach assures a correct value (within the accuracy of one-group perturbation theory) for near the center of the core, where the reactivity feedback due to core motion occurs in accidents of interest in our studies, i.e., Doppler-controlled accidents. A value for the shape factor c in Eq. (A3.8) can be determined with sufficient accuracy (since it only weakly influences the result) by a simple curve fit of V F obtained from a one-group perturbation calculation for an approximately equivalent spherical core. The value for c is typically near unity for cores that have been studied. 

This higher-order power series, not feasible in the cylindrical version because of the additional complexity of the space integration of Eq. (A2.2I), allows more flexibility in curve-fitting the equation of state data presented in Appendix B. The final expression for the disassembly reactivity feedback is obtained by substituting Eqs. (A3.9) and (A3.10) into the general expression given by Eq. (A 1.23). 

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