Chapter 3 – REFERENCE PLANT DESCRIPTION AND ROD
EJECTION ACCIDENT
3.1.
REFERENCE PLANT DESCRIPTION
3.1.1. TERMAL-HYDRAULIC DATA
The NPP that has been chosen for simulating a reactivity insertion accident was the Three Mile Island Unit 1, a PWR-type plant, supplied by Babcock & Wilcox, with nominal power of 2772 MWt (850 MWe), still in operation in Pennsylvania (USA), which was object of an OECD
benchmark for the study of a Main Steam Line Break (MSLB) accident [5]. This plant started operation on April 1974 and is currently owned by Exelon Corporation and operated by AmerGen Energy Company. Located on the same site, there was another nuclear reactor, the Three Mile Island Unit 2, that was shut down after a serious accident occurred in 1979, which resulted in the melting of the reactor core.
Detailed information, concerning thermal-hydraulic and neutronic data related to the reference plant is presented in [5]. Most relevant information, useful for the comprehension and development of the present work, is summarised in this chapter.
The NPP is equipped with 2 Once-Through Steam Generators (OTSG) and composed by two loops, each one containing one hot leg (HL) and two cold legs (CL). The reactor is assumed to be at the end of cycle (EOC), with 650 effective full power days, boron concentration of 5 ppm, and equilibrium Xenon and Samarium concentration.
The OTSG are characterised by providing superheating at the outlet (in the steam lines) and bypass recirculation occurring through holes between riser and downcomer below feedwater entrance nozzle.
Tables 3.1 to 3.6 contain many relevant data related to the reference plant and Table 3.7 presents initial spatial power distribution.
A general arrangement of the reactor cooling system and OTSG is showed in Fig. 3.1 and 3.2.
ITEM DATA
Operating temperature (°C) 318
Overall height of vessel and closure head (m) 12.42 Straight shell minimum thickness (cm) 21.43 Water volume (core and internals in place) (m3) 113.6
Thickness of insulation (cm) 10.2
Flange ID (cm) 425.50
Shell ID (cm) 434
Inlet nozzle ID (cm) 71.1
Outlet nozzle ID (cm) 91.4
Core flooding water nozzle ID (cm) 29.21
Coolant operating temperature – inlet (°C) 291 Coolant operating temperature – outlet (°C) 318
Reactor coolant flow (kg/hr) 63.37 x 106
Closure head minimum thickness (cm) 16.83
Lower head minimum thickness (cm) 12.7
Control rod drive nozzles ID (cm) 7.01
Axial power shaping rod drive nozzles ID (cm) 7.01
Table 3.1 – Reactor vessel design data
DESCRIPTION VALUE Lower plenum (m3) 8.27 Core (m3) 20.4 Down comer (m3) 34.69 Upper plenum (m3) 21.97 Upper head (m3) 14.4
DESCRIPTION VALUE
Total core power output (MWt) 2 772
Design core flow available for heat transfer (106 kg/hr) 58.74 Core flow area available for heat transfer (m2) 4.57
Core pressure drop (kPa) 200
Reactor coolant system pressure drop (kPa) 738
Unrecoverable core pressure drop (kPa) 129
Average core coolant velocity (m/sec) 5.03
Cold leg coolant velocity (m/sec) 14.69
Hot leg coolant velocity (m/sec) 19.45
Table 3.3 – Reactor coolant system steady-state parameters
DESCRIPTION VALUE
Pressurizer (at 558.8 cm water level) Water volume (m3)
Steam volume (m3)
22.7 19.8
Cold leg – each (m3) 6.73
Hot leg – each (m3) 13.3
Reactor coolant pumps (m3) 1.59
Surge line (m3) 0.566 Containment Free volume (m3) Sprayed volume (m3) 5.99 x 104 4.61 x 104 Table 3.4 – Reactor coolant system volume data
DESCRIPTION VALUE
Total reactor flow (kg/hr) 63.4 x 106
Reactor coolant pumps 4 pumps (kg/sec/pump)
5 668 High pressure injection
3 pumps (kg/sec/pump)
18.14 Low pressure injection
2 pumps (kg/sec/pump)
181.4
DESCRIPTION VALUE Volume Lower plenum (m3) Upper plenum (m3) Secondary side (m3) 7.84 7.96 96.62 Steam conditions at full load, outlet nozzles
Steam flow (kg/hr) Steam temperature (°C) Steam pressure (MPa)
2.74 x106 299 (20°C superheat)
6.41
Feedwater temperature (°C) 238
Reactor coolant flow (kg/hr) 31.68 x106
Reactor coolant side
Operating pressure (MPa) Operating temperature (°C) Outlet
Inlet
Coolant volume – hot (m3)
14.96 318 291 57.12 Secondary side
Operating pressure (MPa) Net volume (m3)
6.41 96.62 Dimensions
Tubes – OD/min. wall (cm) Number of tubes
Overall height – including skirt (m) Shell – OD (cm)
Shell minimum thickness – at tube sheets and feedwater connect (m)
Shell minimum thickness (cm) Tube sheet – thickness (cm) Dry weight (kg)
Exposed tube length (ft/m)
1.59/0.086 15 500 22.314 383.858 1539.9 10.636 60.96 519 136 16.269
3.1.2. CORE AND NEUTRONIC DATA
The radial geometry of the core is presented in Fig. 3.3, in which it is divided into cells 21.811 cm wide, each corresponding to fuel assemblies and reflector assemblies (shaded area). It corresponds to an incomplete 17 x 17 matrix, totalizing 241 assemblies (177 FA + 64 reflector assemblies). Main data related to fuel assemblies are presented in Table 3.8.
Figure 3.3 – Core radial geometry
PARAMETER VALUE
Pellet diameter (mm) 9.391
Clad diameter (outside) (mm) 10.928
Clad wall thickness (mm) 0.673
Fuel rod pitch (mm) 14.427
Guide tube diameter (outside) (mm) 13.462
Guide tube diameter (inside) (mm) 12.650
Geometry 15 x 15
Number of fuel pins 208
Number of guide tubes 16
Axially, the active core is divided in 24 layers with varying heights, adding up to a total core height of 357.12 cm. In addition, both the upper and lower axial reflectors have a thickness of 21.811 cm.
Fuels assemblies with different U235 enrichment and varying number of burnable absorber rods are present in the core. Relevant data concerning the thirty different assembly types present in the core are showed in Table 3.9.
Table 3.9 – Definition of assembly types
The radial arrangement of control assemblies is presented in Fig. 3.4. There are 61 Control Assemblies (CA), divided in 7 groups. These rods contain a strong neutron absorber over a length that spans most of the active core region. The total CA length, which coincides with the absorber length, is 342.7055 cm. No tip of control rods is defined.. Measured in units of steps, complete insertion and withdrawal of a CA correspond to 0 and 971 steps, respectively. Each step is 0.3531 cm.
In addition, eight of the CA (Group 8) consist of part-length control rods (axial power shaping rods, or APSR) whose presence is accounted for in the cross-section tables, not needing to be considered as an active bank when modelling the core neutronics.
Figure 3.4 – Arrangement of control rods
Two prompt and six delayed neutron groups are modelled. Table 3.10 shows the time constants and fractions of delayed neutrons.
It was recommended that ANS-79 be used as a decay heat standard model. In total 71 decay-heat groups are used: 69 groups are used for the three isotopes 235U, 239Pu and 238U with the decay-heat constants defined in the 1979 ANS standard.
GROUP DECAY CONSTANT (S-1) RELATIVE FRACTION OF DELAYED NEUTRONS IN % 1 0.012818 0.0153 2 0.031430 0.1086 3 0.125062 0.0965 4 0.329776 0.2019 5 1.414748 0.0791 6 3.822362 0.0197
Total fraction of delayed neutrons: 0.5211%
Thirty assembly types are contained within the core geometry. There are 438 unrodded and 195 rodded (CR inserted in FA) compositions. The corresponding sets of cross-sections are provided. Each composition is defined by material properties (due to changes in the fuel design) and burn-up. The burn-up dependence is a three-component vector variables: exposure (GWd/t), spectral
history (Tmod) and burnable poison (BP) history.
The radial distribution of these assemblies within the reactor geometry is shown in Table 3.11. The 2D assembly type map is shown in a one-eighth core symmetry sector together with the assembly exposure values at EOC. The axial locations of compositions for each assembly are shown in Table 3.12.
A complete set of diffusion coefficients and macroscopic cross-sections for scattering, absorption, and fission as a function of the moderator density and fuel temperature is defined for each composition. The assembly discontinuity factors (ADFs) are taken into consideration implicitly by incorporating them into the sections in order to minimise the size of the cross-section tables.
The group inverse neutron velocities are also provided for each composition. Dependence of the 15 cross-sections on the above variables is specified through a two-dimensional table look-up. Each composition is assigned to a cross-section set containing separate tables for the diffusion coefficients and cross-sections, with each point in the table representing a possible core state. The expected range of the transient is covered by the selection of an adequate range for the independent variables shown in Table 3.13. A linear interpolation scheme is used to obtain the appropriate total cross-sections from the tabulated ones based on the reactor conditions being modelled.
Table 3.14 shows the macroscopic cross-section table structure for one cross-section set. All cross-section sets are assembled into a cross-section library. The cross-sections are provided in separate libraries for rodded and unrodded compositions. All cross-section data was supplied by the benchmark.
3.2.
CONTROL ROD EJECTION ACCIDENT
The Rod Ejection Accident (REA) is defined as the loss of integrity of the CR drive housing with rapid expulsion of a control assembly from the core due to the differential pressure between primary coolant and the containment. It is a postulated accident, or a deviation from the normal operations of the plant of safety concern, with a probability of occurring less than 10-2 /year [8]. This event leads to a rapid reactivity insertion, causing a RIA, and to a SBLOCA at the same time. The consequences of this accident is a rapid reactivity insertion together with a power burst and an adverse core power distribution. If this event were to happen, a fuel rod thermal transient which could cause DNB may occur together with limited fuel damage.
Safety aspects leading to the challenging of acceptance criteria are as follows:
1. Rapid reactor power increase resulting in a fuel temperature rise and in a reduction of DNB Ratio; hence a reduction of heat removal and potential for consequential fuel rod damage and radioactivity release;
2. Primary coolant pressure increase as a consequence of the power increase as well as of a turbine trip. This depends on the actual break size, on whether the pressure increase will in fact occur and also on whether there will be a need for actuation of the ECCS due to a loss of primary coolant;
3. Containment pressure and differential pressures increase, leading to pressure loading of the containment walls; owing to the smaller break size, this aspect is usually much less important; and
4. Radiological consequences due to a loss of primary coolant, potentially also due to a loss of cladding integrity or fuel disintegration.
In the present work, safety evaluation will be focused on the first aspect.
The amount of fuel damage that can result form such an accident will be governed mainly by the worth of the ejected rod and the power distribution attained with remaining control rod pattern [16]. Therefore the fuel suffers a rapid overheating that can lead to the partial melting of the pellets and to their fragmentation. The DNB that could happen can also lead to a clad rupture with consequently fuel dispersal into the coolant. The transient is terminated by the Doppler reactivity effect of the increased fuel temperature, by the negative moderator temperature and eventually by a reactor scram caused by high neutron flux signal.
Figure 3.5 – Rod ejection accident scheme
The break in the pressure housing mechanism create also a hole in the top of the reactor vessel, producing a SBLOCA. This scenario would initiate a depressurization of the RCS. The effects of this scenario are studied in the SBLOCA analyses; in the present work, it will be assumed that the hole is plugged by the ejected rod. As a consequence, a pressure increase may occur in the RCS until pressure-relieving devices act to reduce it. Therefore, this event could challenge the faulted condition stress limits of the RCS [16].
Feature in the design preclude the possibility of a REA, or limit the consequences of the event if it were to occur. These include a thorough quality control program during assembly, and a nuclear design which lessens the potential ejected worth of the control rods. The consequences of this accident are limited by the presence of control rod insertion limits and the limited reactivity worth of the individual control rods. The control rod insertion limits, which vary as function of power level, are present to limit the reactivity worth of the rod inserted in the core when the core is critical. In addition, by ensuring that adequate shutdown margin is available throughout the cycle at all power levels, the rod insertion limits also constrain the amount of reactivity associated with the ejection of a control rods as well as the resulting core peaking factors following the ejection.
Typically, the plant operates with the rods well above the insertion limits. Therefore, only a minor reactivity excursion would be expected following a REA. However, the plant may
occasionally operate with control rods near the insertion limits after changes in plant power level or during load follow operation.
For the purpose of the analysis, the accident will be simulated by linearly introducing reactivity up to the ejected rod worth within a sufficiently short time span (0.1 s).This linear reactivity addition is a calculation convenience rather than an attempt to simulate the actual ejection of a rod. Such an assumption is acceptable since the 0.1 s ejection time is rapid enough compared with the non-linear feedback effects, so that the calculation is not sensitive to the expected shape of the reactivity versus distance curve.
The value of the reactivity to be inserted during the accident will be calculated by means of 3D steady-state neutronic calculations with PARCS code, as will be presented in the next chapter. That incorporate all possible reactor states and all possible CR positions allowed by the operational limits and conditions of the NPP.
