ML21265A552
| ML21265A552 | |
| Person / Time | |
|---|---|
| Site: | University of California-Davis |
| Issue date: | 09/22/2021 |
| From: | McClellan Nuclear Research Center |
| To: | Office of Nuclear Reactor Regulation |
| Shared Package | |
| ML21265A540 | List: |
| References | |
| Download: ML21265A552 (29) | |
Text
Question 4-33
Equations (18) and (19) (page 4-81) show (, ) both on the left-hand-side and right-hand side. Equation (20) does not show how the peaking factor is calculated, since right-hand-side variables are continuous and not known.
Equations (18) and (19) represent a discrete averaging of an approximated continuous function over the bounded range i and i+1. The left-hand side of equations (18) and (19) should reflect this cell-centered average position to be i+1/2 as noted below.
(18)
(19)
Question 4-34
There are several references to tables in Section 4.7, Thermal and Hydraulic Design, that do not list table numbers (such as page 4-80) where is shows Table 4.X where X is missing.
PG 4-81: The following paragraph has been revised to note the accurate table - Table 4.7.
Table 4.7: Hot channel power summary Hot Channel Power Summary Hot Hot Channel Hot Channel Effective Hot Hot Channel Hot Channel Fuel Axial Fuel Radial Channel Core Configuration Channel Thermal Peak Factor Peak Factor Peak Factor Peak Factor Location Power [Pmax/Pavg] [Pmax/Pavg] [Pmax/Pavg] [Pmax/Pavg]
[kW]
OCC BOL Core I6 14.81 1.511 1.230 1.314 2.442 LCC BOL Core I6 17.69 1.804 1.218 1.681 3.694 LCC EOL Core I6 17.59 1.794 1.218 1.667 3.643
The hot channel thermal power, axial and radial power profiles were obtained from the MCNP5 analysis.
The hot channel power factor was obtained by taking the ratio of the hottest fuel element thermal power to the average fuel element thermal power in each core configuration. It is important to note that the hot channel thermal power found in Table 4.7 represents the thermal power after applying the hot channel peak factor. Similarly, the axial peak factor was obtained from referring to the axial power distribution procured from the MCNP5 analysis by taking the ratio of the hottest axial nodalized thermal power value to the average axial nodalized thermal power value. The radial peak factor was calculated by normalizing the thermal power in cylindrical coordinates as follows; all three power profiles are graphically presented below in Figure 4.44.
Page 4-89: The following paragraph has been revised to note the accurate table references
Past studies have been performed toward quantifying TRIGA core form losses of different lattice configurations, the results from these studies represented large variance in their results, and were not able to correlate a definite form loss value for each lattice configuration [4.55]. Previous studies performed by TRIGA Reactors during license efforts developed and presented a methodology for calculating each effective subchannel form loss rather than local form losses within the core [4.48].
These coefficients as well as a summary of the thermal hydraulic parameters found in the MNRC core hot channel are presented in Table 4.13. A brief description of how these form losses were calculated is presented herein.
Page 4-91: The following paragraph has been revised to note the accurate table references.
4.13
The initial outlet coolant temperature and absolute pressure are defined respectively as 45.0 °C (113.0 °F) and 1.58145E5 E5 Pa (22.9370 psia). These boundary input parameters can be seen in Table.
4.8. The geometric parameters for volume 104 are presented in Table 4.8. It is important to note that the boundary conditions in the coolant sink do not convect back into the solution domain.
Question 4-35
Provide the RELAP5 input and outputs used in Section 4.7 (page 4-77), Thermal and Hydraulic Design.
A reference of all inputs used to develop the RELAP5 model geometric construct and boundary conditions are detailed in section 4.7.1 - Description of the RELAP5-3D Model, including explicit identification of each quantity used as input in tables contained within that section (Table 4.8, Table 4.9, Table 4.10, Table 4.11, Table 4.12, Table 4.13, and Table 4.14).
All outputs utilized to support the analysis are detailed in section 4.7.2 - Steady State Results. Including.
Question 4-36
On page 4-83, it states, In these 20 axial nodal locations a modified cosine axial heat distribution has been applied, based on the results produced from the MCNP5 model.
How is the cosine distribution modified to reflect the power distribution tallied in MCNP?
The description of modified cosine axial heat distribution in the description was intended to reflect similar descriptions used in reactor physics text books1. The reference of modified cosine axial heat distribution references a non-uniform cosine heat distribution with tails at the top and bottom of the fuel element resulting from neutron reflection and down-scatter in energy. No modification of the heat distribution was conducted in process relative to that output by the MCNP5 analysis.
1 J.J. Duderstadt, L.J. Hamilton, Nuclear Reactor Analysis, John Wiley & Sons Inc, New York 1976.
Question 4-37
Section 4.7.1 (page 4-82), Description of the RELAP5-3D Model, states This model is based on a single-channel analysis assumed to represent the hottest channel via combination of smallest hydraulic geometry and highest-power element in the core.,
however, Section 4.7.3 (page 4-94), Steady State Results, states The predicted steady state thermal-hydraulic performance of the MNRC OCC and LCC core configurations is determined for the reactor operating at 1.0 MWth How was a core configuration operating at 1.0 MWth simulated with a single channel model? Why was 1.0 MW used when the scram setpoint from Technical Specification 3.1.1 is 1.02 MW?
The analysis approach utilized in support of the thermal hydraulic characterization of the MNRC was to focus on a single channels thermal and hydraulic characteristics. Specifically, the hot channel was chosen in which this channel exhibits the most conservative boundary conditions within the core geometry (i.e. the highest single powered element, the largest axial power distribution, the largest radial power distribution, the smallest fuel-element pitch, etc.). These characteristics were implemented in the RELAP5-3D model using inputs detailed in section 4.7.1.1. These characteristics were operated by considering a single fuel elements total power contribution (the highest-powered fuel element) while the core is operating at an integral power level of 1.0 MWth. A single-channel approach to modeling thermal hydraulic characteristics of TRIGA reactors has proven successful while also producing conservative results in recent past studies2,3,4.
A 1.0 MWth integral core power level was chosen for the total core power given a desire to comprehensively characterize the relevant characteristics of the MNRC in its intended steady state operational mode. A substantial safety margin is shown in all results to demonstrate a safe power SCRAM setpoint of 1.02 MWth.
2 Oregon State University, Safety Analysis Report for the Conversion of the Oregon State TRIGA Reactor from HEU to LEU Fuel. 2007 3 Marcum, W.R., et al., Steady-State Thermal-Hydraulic Analysis of the Oregon State University TRIGA Reactor Using RELAP5-3D. Nuclear Science and Engineering, 2009. 162(3): p. 261-274.
4 Marcum, W.R., B.G. Woods, and R. S.R., Experimental and Theoretical Comparison of Fuel Temperature and Bulk Coolant Characteristics in the Oregon State TRIGA Reactor during Steady State Operation. Nuclear Engineering and Design, 2010. 240: p. 151-159.
Question 4-38
In Section 4.7.1.2 (page 4-84), Coolant Source (100), it states that The MNRCs technical specification requires a minimum water column height above the top of the core to be 7.01 meters (23 feet), however, this technical specification actually refers to the height of water in the tank, not above the core. What is the actual distance from the top of the core to the point where the water level is 23 feet above the bottom of the tank?
This distance is going to be much smaller than the 23 feet of water assumed to be above the core in the calculations. The result will end up lowering the pressure at the core outlet and change all the MDNBR results. Provide drawings (similar to Figure 4.2 on page 4-4) that show dimensions. The core itself is modeled as being 0.711 meters high.
The description of the safety plate in Section 4.2.6.3 (page 4-20), Safety Plate, suggests that there is at least an additional 18.25 inches (0.4636 m) below the core, implying that the actual height of the water column over the core should be at most 5.84 meters.
Based on description of the central thimble in Section 10.4.1 (page 10 -12), Central Irradiation Facility, there are less than 55 inches from the bottom of the tank to the top of the core which would result in 5.61 meters of water over the top of the core, at most.
An evaluation of the geometries within the MNRC primary was performed and found that a normal tank level to the top of the active fuel is 20.5 feet. For conservatism the core analysis was reevaluated at an assumed height of 19.0 related to the LSSS. This corresponds to an absolute pressure of 1.58145E5 E5 Pa (22.9370 psia). The result of adjusting the tank water height, in combination with the form losses (detailed in Question 4-42) requires the complete revision of all results. All results are presented hereinafter for each of the three core configurations, taking into account all items denoted in response to Question 4-38 and Question 4-42.
1.1.1 OCC BOL Core
The OCC core contains fuel elements that are geometrically similar and therefore the hot channel geometric parameters (i.e. hydraulic diameter, length, etc.) do not change. The hot channel power summary in terms of parameters and results are given in Table 2.1 and Table 2.2, respectively.
Figure 2.1 through Figure 2.5 graphically illustrate the results of the analysis on the OCC core.
Each parameter in Figure 2.1 is taken at a different elevation in the hot subchannel. The fuel centerline temperature is shown at the axial nodal location which produces the maximum f uel centerline temperature (fuel axial center). The outer cladding temperature is shown at the axial location which produces the maximum outer cladding temperature (slightly above the axial fuel centerline). The bulk coolant temperature is shown at the location which produces the maximum bulk coolant temperature (highest vertical subchannel node). Coolant mass flux is shown as the mass flux which corresponds to the maximum bulk coolant temperature (highest vertical subchannel node). Each location was selected to show the most limiting value of the associated parameter.
Table 2.1: Steady state results for the OCC BOL Core at 1.0 MWth
Parameter Value Flow rate for hottest rod [kg/s] 0.0848 Maximum wall heat flux [kW/m2] 421.931 Maximum fuel centerline temperature [C] 387.64 Maximum clad temperature [C] 129.92 Exit clad temperature [C] 126.93 Exit bulk coolant temperature [C] 86.68 Groeneveld 1986 8.61 MDNBR Groeneveld 1995 6.96 Groeneveld 2006 6.40 Bernath 3.16
Figure 2.1: Hot channel characteristics (OCC BOL Core)
Figure 2.2: Hot channel MDNBR (OCC BOL core)
Figure 2.3: Axial temperature distribution at 14.81 kWth (OCC BOL Core)
Figure 2.4: Hot channel axial DNBR at 14.81 kWth (OCC BOL Core)
Figure 2.5: Radial temperature distribution at 14.81 kWth (OCC BOL Core)
Table 2.2: Calculated fuel temperatures for various channel powers (OCC BOL Core)
P (kW) Calculated [C]
hot-channel T T T T max 0.3 clad coolant 14.81 386.53 370.94 129.82 86.84 15.00 389.86 374.06 129.94 87.16 20.00 476.31 455.25 132.88 95.10 25.00 561.15 534.82 135.31 99.58 30.00 644.71 613.12 137.41 101.84
The OCC BOL Core steady state results shown above provide a succinct summary of the thermal safety related characteristics while operating at the license limit power, lowest water level, and hottest water temperature in the hot-channel (most limiting location within the core).
Figure 2.2 shows that the MDNBR in the hot channel will reach a value of 2.00 at approximately 21.0 kWth hot channel steady state power. This is 141.8% of the 14.81 KWth produced in the hot channel of the OCC BOL Core operating at 1.0 MWth. Using either the Bernath or the Groeneveld 2006 correlations, the OCC BOL Core is operating at power well below that required for departure from nucleate boiling.
1.1.2 LCC BOL Core
The LCC BOL Core contain fuel elements that are geometrically similar and therefore the hot channel geometric parameters (i.e. hydraulic diameter, length, etc.) do not change. The hot channel power summary in terms of parameters and results are given in Table 2.3 and Table 2.4, respectively.
Figure 2.6 through Figure 2.10 graphically illustrate the results of the analysis on the LCC BOL Core.
Each parameter in Figure 2.6 is taken at a different elevation in the hot subchannel. The fuel centerline temperature is shown at the axial nodal location which produces the maximum fuel centerline temperature (fuel axial center). The outer cladding temperature is shown at the axial location which produces the maximum outer cladding temperature (slightly above the axial fuel centerline). The bulk coolant temperature is shown at the location which produces the maximum bulk coolant temperature (highest vertical subchannel node). Coolant mass flux is shown as the mass flux which corresponds to the maximum bulk coolant temperature (highest vertical subchannel node). Each location was selected to show the most limiting value of the associated parameter.
Table 2.3: Steady state results for the LCC BOL Core at 1.0 MWth
Parameter Value Flow rate for hottest rod [kg/s] 0.08531 Maximum wall heat flux [kW/m2] 492.10 Maximum fuel centerline temperature [C] 409.96 Maximum clad temperature [C] 131.46 Exit clad temperature [C] 130.71 Exit bulk coolant temperature [C] 91.49 Groeneveld 1986 6.85 MDNBR Groeneveld 1995 5.38 Groeneveld 2006 5.02 Bernath 2.26
Figure 2.6: Hot channel properties (LCC BOL Core)
Figure 2.7: Hot channel MDNBR (LCC BOL Core)
Figure 2.8: Axial temperature distribution at 17.69 kWth (LCC BOL Core)
Figure 2.9: Hot channel axial DNBR at 17.69 kWth (LCC BOL Core)
Figure 2.10: Radial temperature distribution at 17.69 kWth (LCC BOL Core)
Table 2.4: Calculated fuel temperatures for various channel powers (LCC BOL Core)
P (kW) Calculated [C]
hot-channel T T T T max 0.3 clad coolant 15.00 363.19 350.79 129.58 87.24 17.69 405.27 390.64 131.21 91.64 20.00 441.00 424.46 132.47 94.77 25.00 517.28 496.61 134.88 99.54 30.00 592.32 567.51 136.96 101.89
The LCC BOL Core steady state results shown above provide a succinct summary of the thermal safety related characteristics while operating at the license limit power, lowest water level, and hottest water temperature in the hot-channel (most limiting location within the core).
Figure 2.7 shows that the MDNBR in the hot channel wil l reach a value of 2.00 at approximately 19.10 kWth hot channel steady state power. This is 108 % of the 17.69 kWth produced in the hot channel of the LCC core operating at 1.0 MW th. Using either the Bernath or the Groeneveld 2006 correlations, the LCC core is operating at power well below that required for departure from nucleate boiling.
1.1.3 LCC EOL Core
The LCC EOL Core contain fuel elements that are geometrically similar and therefore the hot channel geometric parameters (i.e. hydraulic diameter, length, etc.) do not change. The hot channel power summary in terms of parameters and results are given in Table 2.5 and Table 2.6, respectively.
Figure 2.11 through Figure 2.15 graphically illustrate the results of the analysis on the LCC EOL Core.
Each parameter in Figure 2.11 is taken at a different elevation in the hot subchannel. The fuel centerline temperature is shown at the axial nodal location which produces the maximum fuel centerline temperature (fuel axial center). The outer cladding temperature is shown at the axial location which produces the maximum outer cladding temperature (slightly above the axial fuel centerline). The bulk coolant temperature is shown at the location which produces the maximum bulk coolant temperature (highest vertical subchannel node). Coolant mass flux is shown as the mass flux which corresponds to the maximum bulk coolant temperature (highest vertical subchannel node). Each location was selected to show the most limiting value of the associated parameter.
Table 2.5: Steady state results for the LCC EOL Core at 1.0 MWth
Parameter Value Flow rate for hottest rod [kg/s] 0.09085 Maximum wall heat flux [kW/m2] 489.32 Maximum fuel centerline temperature [C] 408.38 Maximum clad temperature [C] 131.40 Exit clad temperature [C] 130.65 Exit bulk coolant temperature [C] 91.33 Groeneveld 1986 6.90 MDNBR Groeneveld 1995 5.41 Groeneveld 2006 5.05 Bernath 2.28
Figure 2.11: Hot channel properties (LCC EOL Core)
Figure 2.12: Hot channel MDNBR (LCC EOL Core)
Figure 2.13: Axial temperature distribution at 17.59 kWth (LCC EOL Core)
Figure 2.14: Hot channel axial DNBR at 17.59 kWth (LCC EOL Core)
Figure 2.15: Radial temperature distribution at 17.59 kWth (LCC EOL Core)
Table 2.6: Calculated fuel temperatures for various channel powers (LCC EOL Core)
P (kW) Calculated [C]
hot-channel T T T T max 0.3 clad coolant 15.00 363.19 350.79 129.58 87.24 17.59 403.71 389.17 131.15 91.48 20.00 441.00 424.46 132.47 94.77 25.00 517.28 496.61 134.88 99.54 30.00 592.32 567.51 136.96 101.89
The LCC EOL Core steady state results shown above provide a succinct summary of the thermal safety related characteristics while operating at the license limit power, lowest water level, and hottest water temperature in the hot-channel (most limiting location within the core).
Figure 2.12 shows that the MDNBR in the hot channel will reach a value of 2.00 at approximately 19.10 kWth hot channel steady state power. This is 108.6 % of the 17.59 kW th produced in the hot channel of the LCC core operating at 1.0 MW th. Using either the Bernath or the Groeneveld 2006 correlations, the LCC core is operating at power well below that required for departure from nucleate boiling.
2 CLOSING
A comprehensive thermal hydraulic analysis was performed of the UCD MNRC using RELAP5-3D while applying best practice approaches as they related to previous applications reviewed and approved by the U.S. Nuclear Regulatory Commission in support of TRIGA reactor licensing. Three core configurations were considered as a part of this analysis including the Operational Core Configuration and Beginning of Life, as well as Limiting Core Configuration at both Beginning and End of Life. It was found that the most limiting core configuration was the LCC BOL Core producing a predicted minimum departure from nucleate boiling ratio of 2.26 at full power steady state operations with the most conservative predictive correlation applied (the Bernath Correlation). All analyses demonstrated that no safety limit is credibly compromised during normal operations with sufficient margin provided to ensure safe and reliable operations of the MNRC.
Question 4-39
On page 4-86, the pitch is listed as 1.60 inches, which is inconsistent with the pitch listed on page 4-17, of 1.714 inches. Which is the correct pitch?
The reference to hexagonal pitch of 1.714 inches on page 4-17 is the nominal average pitch between fuel elements as fabricated into the upper grid plate. The RELAP5-3D model was developed to include the most limiting geometric conditions including those factors resulting from epistemic uncertainty.
Including the fabrication uncertainties when machining the upper grid plate, a minimum pitch between two fuel elements may result in 1.60 inches. This was chosen to be implemented within the RELAP5-3D model to ensure that the most conservative outputs and limiting conditions were produced in support of the safety analysis.
Question 4-40
In Table 4.13 on page 4-89, the flow area is listed as 3.80E-4 m2, which contradicts the value of 3.3E-4 m2 listed on page 4-86. What flow area was used in the RELAP5 calculations used to obtain the results shown in this section?
The flow area used for both the cold-leg and the hot-channel was 3.80E-4 m2 (0.589 inch2). This was accurately represented in the RELAP5-3D model, however was inconsistently detailed int the document for the cold-leg.
Question 4-41
In Table 4.13 on page 4-89, why do Heated Diameter and Hydraulic Diameter have different values? The wetted perimeter is the same as the heated perimeter in the subchannel shown in Figure 4.47 on page 4-86.
The heated diameter is reflected by the outer diameter of the fuel element. Whereas the hydraulic diameter is acquire through the relation:
This this case Afl is the hexagonal subchannel area (3.805E-4 m2) and P is the wetted perimeter (0.117 m).
Question 4-42
In Table 4.13 on page 4-89, how are inlet and exit pressure loss coefficients derived from the values given in the text (on pages 4-88 and 4-89)?
The process by which the inlet and exit pressure loss coefficients were computed was baselined against the calculation performed by General Atomics in 20085 whereby the regions were tabulated and the sum of the form loss coefficients added. These regions comprised the pool to grid-hole, grid-hold to above grid-hole, and above grid-hold to fuel element. This method resulted in the calculation of an inlet pressure loss coefficient of 0.587 and an exit pressure loss coefficient of 0.539 as noted in the following table.
Parameter Value Flow area [m2] 3.805E-4 Fuel Element Pitch [m] 0.04064 Wetted perimeter [m] 0.117 Hydraulic diameter [m] 1.301E-2 Heated diameter [m] 3.734E-2 Fuel element heated length [m] 0.381 Fuel element surface area [m2] 4.469E-2 Fuel element surface roughness [m] 2.134E-6 Inlet pressure loss coefficient 0.587 Exit pressure loss coefficient 0.539 Absolute pressure at the top of the core [Pa] 1.70068E5
5 J. Bolin, TRIGA Reactor Thermal Hydraulic Study, TRD 070.01006.04 Rev A. TRIGA Reactors Division, General Atomics-ESI, San Diego, California, 2008.
Question 4-43
On page 4-91, it states The fuel to clad contact gap that is created by material surface roughness is originally hydrided during manufacturing of TRIGA fuel. As the U-ZrH fuel is burnt through its lifetime fission product gasses are released and migrate from the fuel lattice structure into the gap. Is this the same process referred to on page 4-44, which states Measurements of fuel temperatures as a function of steady-state power level provide evidence that after operating at high fuel temperatures, a permanent gap is produced between the fuel body and the clad.?
That is correct, these two references are articulating the same phenomena. The former details the process by which fission product gas migrates to the fuel-gap region; the latter details the empirical method by which fuel temperature is monitored to confirm this phenomenon does not compromise the safety of the reactor.
Question 4-44
On page 4-92, it states The outer gap coordinate (Node 22) is varied during this study.
What gap thickness was used for the results shown in Section 4.7.3 (page 4-94), Steady State Results? Are the results sensitive to this parameter? It appears from the data in Table 4.14 (page 4-92), Heat structure radial node lengths, that the fuel cladding thickness was changed to vary the gap width. Why was this done as opposed to varying the outer fuel radius?
A gap distance of 2.54 µm (0.1 mil) was chosen based on the result of the sensitivity study performed.
This is consistent with other analyses performed in support of recent TRIGA reactor thermal hydraulic studies.
The gap region was adjusted by thinning the cladding as necessary. This was done as apposed to adjusting the outer fuel radius for conservatism. The fuel is cast as a part of the fabrication process and the geometry is highly controlled as a part of this process. If the fuel were to vary in dimension after fabrication it would only increase in radius due to fracture or swelling. In doing this would increase the outer radius of the fuel only and reduce the power density, inherently producing lower local temperatures. The conservative nature of this study resulted in the decision to decrease cladding thickness to ensure that the most conservative results were produced in support of the thermal hydraulic analysis.
Question 4-45
Table 4.14 (page 4-92 and 4-93), Heat structure radial node lengths, appears to have inconsistent information relative to other portions of the document. The table shows a cladding outer radius of 0.7375 in, however, it is shown as 1.5 in / 2 = 0.75 in on page 4-8, while Table 4.1 (page 4-10),
SUMMARY
OF FUEL ELEMENT SPECIFICATIONS, shows it as 1.47 in / 2 = 0.735 in. The table shows a fuel outer radius of 0.70275 in, however, Table 4.1 on page 4-10 shows a value of 1.43 in / 2 = 0.715 in. Table 4.1 on page 4-10 shows a cladding thickness of 0.02 in, while Table 4.14 shows values of either 0.03465 of 0.03445 (depending on the gap thickness).
The thermal hydraulic analysis was performed consistent with that documented in Table 4.1 located on page 4-10 as detailed with the outside diameter fuel cladding of 1.47 inches. The outside stainless-steel cladding radius was transcribed incorrectly and is corrected below to reflect what was modeled in RELAP5-3D.
Heat Structure Radial Node Lengths Nodal Description Node Number Node Coordinate [m] (in)
Outer Zirconium Pin 01 0.00000 (0.00000) 02 0.00318 (0.12500) 03 0.00355 (0.13976) 04 0.00430 (0.16929) 05 0.00506 (0.19921) 06 0.00581 (0.22874) 07 0.00656 (0.25827) 08 0.00731 (0.28779) 09 0.00807 (0.31772) 10 0.00882 (0.34724) 11 0.00957 (0.37677)
Fuel 12 0.01032 (0.40630) 13 0.01108 (0.43622) 14 0.01183 (0.46575) 15 0.01258 (0.49527) 16 0.01333 (0.52480) 17 0.01409 (0.55472) 18 0.01484 (0.58425) 19 0.01559 (0.61378) 20 0.01634 (0.64331) 21 0.01710 (0.67323) 22 0.01785 (0.70275)
Outer Gap 23 0.01785-0.01786 (0.70285-0.70305)
Outer Stainless Steel Clad 24 0.018161 (0.7150)
Question 4-46
Section 4.7.3 (page 4-94), Steady State Results, states Groeneveld 1986, 1995, and 2006
[4.62] critical heat flux tables were used as the primary means for predicting margin to departure from nucleate boiling with the he Bernath correlation [4.61] provided as a qualitative reference for historic purposes. Were the results for DNB computed directly by RELAP5, or were RELAP5 output parameters used to separately compute DNB?
RELAP5-3D presently computers Groeneveld 1986 directly, all others were computed using the output quantities resulting from RELAP5-3D external to the software. A verification was performed by comparing the Groeneveld 1986 DNB values computed directly by RELAP5-3D and compared against that of the same calculation external to the software while using the results from RELAP5-3D. The verification exercise resulted in identical predicted values.
Question 4-47
The figures shown on pages 4-95 through 4-97 have an inconsistent numbering scheme (the sequence goes 4.51, 4.52, 4.51, 4.52, 4.53).
The figures presented within pages 4-95 through 4-97 appear in sequential order with the version of record presently in hand. Starting at 4.51 and continuing through 4.55.
Question 4-48
On page 4-95, the text describes coolant mass flux as being shown in Figure 4.52.
Coolant mass flux is not shown in either figure labeled Figure 4.52. Coolant mass flow rate is shown in Figure 4.51 on page 4-95, is this what the text is referring to?
The accurate reference is of Figure 4.51 on page 4-95 in which coolant mass flow is denoted.