Thermal Analysis of the SAFKEG HS DesignThe heat transfer coefficients used to model convective heat transfer in the heating phase of the fire accident are described in Section 5

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Thermal Analysis of the SAFKEG HS Design

To: Croft Associates Ltd

Date: September 2012

From: AMEC

Your Reference: Croft Project No 06/08/10

Our Reference: AMEC/6335/001 Issue 1

AMEC/6335/001 Issue 1 Page 2

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Title Thermal Analysis of the SAFKEG HS Design

Prepared for

Croft Associates Ltd

Your Reference

Croft Project No 06/08/10

Our Reference

AMEC/6335/001 Issue 1

Confidentiality, copyright & reproduction

This report is submitted by Energy, Safety and Risk Consultants (UK) Limited (hereafter referred to as AMEC) in connection with a contract to supply goods and/or services and is submitted only on the basis of strict confidentiality. The contents must not be disclosed to third parties other than in accordance with the terms of the contract.

Contact Details AMEC Kimmeridge House Dorset Green Technology Park Winfrith Newburgh Dorchester Dorset DT2 8ZB United Kingdom Tel +44 (0) 1305 851100 Fax +44 (0) 1305 851105 amec.com

Name Signature Date

Author(s) N Butler & J R Martin

Reviewed by C J Fry

Approved by C J Fry

Transport Flask Photograph courtesy of Magnox Electric Ltd Submarine Photograph by: Mez Merrill; © Crown Copyright/MOD, image from www.photos.mod.uk. Reproduced with the permission of the Controller of Her Majesty’s Stationery Office


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Executive Summary The thermal performance of the SAFKEG HS design of transport container has been analysed using an axi-symmetric finite element model. The model was first validated by comparison against experimental data from a self heating test. Good agreement was obtained with the results of the self heating test, although there may have been significant heat loss through the base of the keg despite it being stood on an insulating surface. The temperature of the container has been predicted under normal conditions of transport, for heat loads of 30, 5 and 0W, both with and without solar insolation. Without solar insolation the temperature at the inner containment vessel lid seal is 135, 56 and 38ºC, respectively. For all internal heat loads, the effect of the solar insolation is to increase the temperature at the inner containment vessel lid seal; by about 15ºC in the case of the 30W heat load, and by about 20°C for both the 5 and 0W heat loads. The temperature of the container has been predicted during the thermal (fire) test. At a heat load of 30W, the inner containment vessel lid seal is predicted to reach a maximum temperature of 196°C. At a heat load of 5W, it is predicted to reach 130°C and at zero heat load it is predicted to reach 115°C. The main effect of having the various air gaps in the model filled with cork rather than air is slightly to enhance the efficiency of heat transfer from the containment vessel to the keg’s outside surface in the case of the NCT simulation, and in the opposite direction in the case of the HAC fire test simulation.


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Contents

1 Introduction 5

2 Description of the Model 5

2.1 Geometry 5 2.2 Materials 6 2.3 Boundary Conditions 7

3 Validation of the Model 8

3.1 Normal Conditions of Transport 8 3.2 Fire Accident 9

4 Calculation of Temperatures During Normal Transport 9

4.1 No Insolation 9 4.2 With Insolation 10

5 Calculation of Temperatures During the Fire Accident 11

6 Air Gap Sensitivity Analysis 13

7 Conclusions 13

8 References 14

Appendix 1 Drawings used in model generation 49 Appendix 2 Natural Convection Heat Transfer Correlations 51


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1 Introduction The SAFKEG HS design is a general purpose container for transporting modest quantities of radioactive material. It consists of an outer keg and an inner containment vessel, as illustrated in Figure 1. The outer keg is ‘beer barrel’ in shape, ~0.6m high and ~0.4m diameter and consists basically of a stainless steel shell filled with cork which provides protection from impacts and fire. The containment vessel is cylindrical in shape, ~0.3m high and ~0.2m diameter. It is also constructed from stainless steel and has a lid at the top. Both the body and lid of the containment vessel contain depleted uranium shielding. An elastomeric seal is used to provide sealing between the inner containment vessel body and lid. The inner containment vessel, when placed inside the keg, is surrounded by a further layer of cork. The SAFKEG HS is designed as a type B(U) package and to meet all the requirements of a B(U) package as specified in the 10CFR71 Transport Regulations [1] with an internal heat load up to 30W. This report presents a thermal assessment of the SAFKEG HS design which has been performed in order to demonstrate that the package satisfies all the thermal performance requirements of the 10CFR71 Regulations. The thermal assessment is based upon finite element modelling. The model is described in Section 2. The model has been validated by comparison against an experimental self heating test (representing normal conditions of transport). This validation is described in Section 3. The model is then used to determine temperatures both under normal conditions of transport and during the fire accident, as specified in the 10CFR71 Regulations [1]. These are described in Sections 4 and 5 respectively.

2 Description of the Model

2.1 Geometry The SAFKEG HS package design is almost axi-symmetric. The main features which are not axi-symmetric are:

� The bolts fastening the keg lid � The handle on the top of the keg lid � Finger holes in the top cork � The bolts in the containment vessel lid � Location and lifting points in the containment vessel lid None of these features was considered to be significant with respect to heat transfer. An axi-symmetric model has therefore been used for the thermal assessment. The axi-symmetric model is shown in Figure 2. The model explicitly represents:

� The keg outer skin � The keg inner liner � The cork filling the keg (between the inner liner and outer skin) � The top cork within the keg cavity � The side cork within the keg cavity � The inner containment vessel � The depleted uranium shielding in the inner containment vessel A 3-dimensional, cut-away, view of what the model represents is shown in Figure 3. The model, which contains 7841 nodes and 5022 elements, was generated using the Abaqus code, version 6.12.1 [2]. Each of the components listed above is generated separately and joined, thermally, using tied constraints or interactions (representing narrow air gaps). The thin outer skin


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of the keg was modelled using ‘shell’ elements while all the other components were modelled using solid elements. The drawings on which the model is based are listed in Appendix 1.

2.2 Materials The following materials are represented in the model:

� Stainless steel

� Cork

� Depleted uranium The thermal properties for each of these materials used in the model are described in the following sections.

2.2.1 Stainless steel

Type 304L stainless steel is used to construct the outer and inner skins and lid of the keg and the inner containment vessel and lid. The thermal properties used for type 304L stainless steel, and their sources, are shown in Table 1. Radiation is an important heat transfer mechanism. The emissivity of the stainless steel surfaces, under different conditions, assumed in the model is shown in Table 2. The emissivity of stainless steel can vary significantly depending upon surface finish and level of oxidation. Good agreement with the measured temperatures in the steady state heating test was obtained with an assumed emissivity of the external keg surface of 0.25 (Section 3.1), which is a typical reference value for stainless steel exposed to the atmosphere.

2.2.2 Cork

Resin bonded cork surrounds the inner containment vessel. Croft have had measurements made at temperatures up to 108°C and the results of these are shown in Figure 4. During the fire the cork will experience temperatures up to ~800°C. No measurements of cork properties at high temperatures are available. However, a furnace test has been performed on the SAFKEG LS container design, which uses the same cork specification as the HS design. This test has been simulated using the LS model [3] in order to validate the model and, in particular, demonstrate the acceptability of the thermal properties assumed for the cork. It was found in [3] that, in order to obtain agreement with the measured temperatures, the thermal conductivity of the cork needed to be increased by 50%. It should be noted that these thermal properties, validated against the furnace test, are ‘effective’ properties that include any effects of charring and shrinkage of the cork. A self heating test performed on the SAFKEG HS design has also been simulated using the model. This is described in Section 3.1. As with the corresponding test on the LS design, it was found that, to produce the best agreement with the measured temperatures, the thermal conductivity of the cork needed to be reduced by 15% (as described in Section 3.1 and shown on Figure 4). Because cork is a natural material, this degree of variation in conductivity may well be possible. To ensure that all the calculations performed with the model are pessimistic, the lower, fitted conductivity has been assumed for the calculations of temperature during normal transport and the higher thermal conductivity assumed for the calculations of temperature during the fire accident. These are the same properties as used for the SAFKEG LS modelling [3].

2.2.3 Depleted Uranium

The shielding inside the containment vessel body and lid is provided by depleted uranium (DU). The thermal properties used for this material, and their sources, are shown in Table 1.


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2.3 Boundary Conditions

2.3.1 Natural Convection from Exterior Surfaces

Well established correlations for natural convection heat transfer coefficient have been applied at the outer surface of the keg. Several separate correlations were used depending upon the surface being considered and the container orientation: � Outside of a vertical cylinder (the side of the keg when vertical) � Horizontal plate (top of the keg when vertical) � Vertical plate (top and bottom of the keg when horizontal) � Outside of a horizontal cylinder (the side of the keg when horizontal) Details of the correlation used are given in Appendix 2. The heat transfer coefficients used to model convective heat transfer in the heating phase of the fire accident are described in Section 5.

2.3.2 Radiation from Exterior Surfaces

Heat loss by thermal radiation was modelled from all the outer surfaces. The emissivity assumed for the surface of the keg is shown in Table 2. The top of the keg and the inner surfaces of the top skirt form an open cavity. In the calculations simulating normal transport, radiation exchange between these surfaces, and to ambient, was modelled (including the reflection of radiation). With the keg vertical, radiation exchange between the bottom of the keg and the skirt was again modelled but there was no net loss of heat as the keg was assumed to be sat upon an insulating surface. In the calculations simulating the fire accident, the flames were pessimistically assumed to fill the cavity inside the skirt at either end of the keg and all surfaces were assumed to receive radiation directly from the flames. The flames were assumed to have an emissivity of 1.0 which satisfies the ‘minimum average flame emissivity coefficient of 0.9’ specified in the 10CFR71 Regulations [1].

2.3.3 Internal Heat Load

The heat generated by the package contents (30W) was represented in the model as a uniform heat flux applied over the side, top and bottom of the cavity inside the containment vessel. The package contents themselves were not represented in the NCT and fire accident models since: � The assessment is intended to be general and not specific to one type and geometry of

contents � Omitting the contents from the fire accident calculation is pessimistic as it minimises the

thermal capacity of the package. In the calculations used to validate the model against the self heating test, changes were made to the model to more accurately represent the contents (described in Section 3).

2.3.4 Narrow Gaps

Narrow air gaps are present in several locations in the model: � Between the keg body and the cork � Between the keg lid and the cork � Between the keg inner liner and the cork � Between the keg inner liner and the keg lid � Between the cork and the containment vessel The thickness of these gaps was obtained from the drawings. Even when surfaces are in contact (particularly a rough surface such as that of cork) a contact resistance will exist which can be


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represented as an air gap. Justification for the assumed gaps was obtained through validation of the model against test data (described in section 3). Across each air gap heat transfer by both conduction and thermal radiation was represented in the model.

2.3.5 Radiation across Open Spaces

The package design creates a number of small cavities across which radiation heat transfer will occur (in such small cavities heat transfer by conduction and convection is expected to be negligible). In these cavities radiation exchange between all the surfaces is modelled. The view factor from each element to each other element in the cavity is determined and radiation heat transfer, including the effect of reflection, is then calculated.

2.3.6 Solar Insolation

The 10CFR71 Regulations [1] require that the effect of solar insolation is considered and specify the incident heat fluxes on various orientations of surfaces which should be assumed for a Type B(U) package. The regulations specify that the solar insolation should be considered for 12 hours each day. In order to determine the maximum temperature of the container under normal conditions of transport, a transient calculation was therefore performed with the heat flux from solar insolation turned on or off every 12 hours. The insolation flux incident on each surface is summarised in Table 3. The Regulations specify the incident radiation flux. In order to determine the absorbed heat flux from solar insolation a surface absorptivity must also be defined. No data is readily available on the absorptivity of bead-blasted type 304 stainless steel to short wavelength radiation. A pessimistically high absorptivity to solar radiation of 0.8 was therefore assumed.

3 Validation of the Model

3.1 Normal Conditions of Transport Measurements of container temperatures under normal conditions of transport with no solar insolation have been made by Croft [4]. The heat source in this test, 31W, was provided by cartridge heaters inside an aluminium block located inside the cavity inside the inner containment vessel. The container was placed vertically upon a wooden board covered in aluminium foil (approximating an insulating surface) in a room where the ambient temperature was relatively constant. The test ran for around seven days with the temperatures being measured every minute. The measured temperatures over the final 24 hours have been averaged to obtain the ‘steady state’ temperatures. Further measurements were subsequently made with the container horizontal, but these were not used for the validation excercise. Seven thermocouples were used to monitor the temperature of the heater block, the outside of the containment vessel and the inside and outside of the keg. An additional thermocouple was used to measure the ambient temperature in the room. Once equilibrium conditions had been achieved, additional measurements were made of the temperature at various locations on the outside of the keg. Full details of the test are given in [4]. The finite element model was modified to represent the heating test. The required modification involved: � Addition of the aluminium heater block � Representation of heat into the heater block � Changing the ambient temperature to that measured in the test (21.7°C) The heater block is fitted inside the containment vessel with just a small air gap around it. Heat transfer by conduction and radiation was simulated in this air gap.


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The predicted temperatures on the outside of the keg were in reasonable agreement with the measured values, except near the base (which is discussed below), but the temperature of the containment vessel was initially underestimated. As with the LS model, the calculation showed that the majority of the temperature difference between the containment vessel and the outer surface of the keg results from heat transfer through the cork. It was therefore concluded that the low predicted containment vessel temperature was probably due to the thermal conductivity of the cork being lower than assumed in the model. The calculation was therefore repeated with the thermal conductivity of the cork reduced by 15% (as was done for the LS model). Such a variation in thermal conductivity is considered possible in a natural material such as cork. The predicted internal temperatures in the repeat calculation are in good agreement with those measured in the test. The predicted temperature profile is shown in Figure 5 and it can be seen that, as expected, the highest temperatures occur in the heater block and inner containment vessel and high temperature gradients are generated in the cork. The predicted temperatures are compared against those measured in the test in Figure 6. The temperatures initially predicted by the model are also shown on this Figure. The predicted internal temperatures agree with the measured temperatures to within 0.8°C, except for thermocouple T4 which is on the keg liner between the inner and outer cork. T4 is over-predicted by 3.5°C, which may be due to the air gaps in this region being different to the nominal values. The predicted temperatures on the external surfaces of the keg agree reasonably well with measured values except near the base, where they are around 2.5° – 7.6°C higher. A similar, though less pronounced effect was found with the LS model [3] near the base. As before, it is thought to be due to the board on which the container stood being modelled as perfectly insulating whereas, in practice, there is some heat loss through the board. The HS temperature discrepancy compared with the LS (around 2°C at the bottom of the keg), is roughly in accordance with the relative heat load, which is three times higher in the HS. It is therefore consistent with interpreting both these results as due to heat loss though the base region. As the predicted temperatures on the outer surface of the keg away from the base are in reasonable agreement with measured values, no change in the external surface emissivity was required. Therefore the reference value of 0.25 was used for the NCT conditions.

3.2 Fire Accident No specific validation of the fire test modelling was done for the SAFKEG HS design. A furnace test was carried out for validation of the LS model [3] and, as there are only minor differences between the two designs, it is considered that the LS validation gives sufficient confidence in the fire modelling of the HS container.

4 Calculation of Temperatures During Normal Transport

4.1 No Insolation The finite element model has been used to determine the temperature of the container under normal conditions of transport in the absence of solar insolation. Steady state calculations were performed which represented the container, stood vertically on an insulating surface, with internal heat loads of 30, 5 and 0W, respectively, and an ambient temperature of 38°C. Heat loss to ambient by radiation and natural convection from the sides and top of the keg was simulated. These boundary conditions (apart from the ambient temperature) were the same as used to obtain good agreement with the self heating test (see Section 3.1).


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In the validation of the model against the self heating test (Section 3.1) it was found that the best agreement was obtained with the thermal conductivity of the cork reduced by 15% compared to the measured values. To ensure that the temperatures predicted under normal conditions of transport are pessimistic, the lower, adjusted, thermal conductivity value has been used. Other properties, e.g. emissivities, were taken to be the same as for the self-heating test. The predicted temperature profile at a heat load of 30W is shown in Figure 7. The temperature at the inner containment vessel lid seal is 135°C, although the maximum temperature reached by the containment vessel is 148°C. The maximum accessible surface temperature is 46.5°C and occurs on the top of the keg lid. The base of the keg is predicted to reach 57°C but this surface is not readily accessible. The temperature on the outside of the keg, at mid-height, is 47°C. The predicted temperature profile at a heat load of 5W is shown in Figure 8. The temperatures are appreciably lower than for the 30W case, with that at the inner containment vessel lid seal being 56°C, and the maximum containment vessel temperature being 59°C. The maximum keg lid surface temperature is 40°C, and the keg bottom 42°C. The temperature on the outside of the keg, at mid-height, is 40°C. For zero internal heat load the keg’s temperature was predicted to be the ambient temperature of 38°C throughout, as confirmed by Figure 9.

4.2 With Insolation The finite element model has been used to determine the temperature of the container under normal conditions of transport and subject to solar insolation. In the SAFKEG LS assessment, both a vertical orientation of the container and a horizontal orientation of the container were studied. The most pessimistic orientation was found to be vertical (i.e. it gave the highest internal temperatures). Therefore only the vertical orientation was modelled for the HS design. As before, the keg is modelled as standing on an insulating surface. The model used was the same as that used for the case of normal transport without insolation (described above in Section 4.1) except that solar insolation was added as a heat flux incident on the top and side external surfaces for 12 hours each day. Transient calculations were performed covering a period of 4½ days, with 12 hours of cooling between each insolation period (and with an ambient temperature of 38°C), and for internal heat loads of 30, 5 and 0W. The temperatures at the end of the calculation were used as the starting point for the fire test calculation. The results for the internal heat load of 30 W are illustrated in Figures 10 – 12. Figure 10 shows the transient temperature at various locations on the outer surface of the keg. The highest temperatures occur on the top of the container because the insolation flux is greater on the top than on the side. The maximum predicted temperature, which occurs on the keg lid, is 102°C. The bottom of keg is hotter than the side even though it does not receive any heat from solar insolation. This results from there being much less heat loss from the bottom of the keg than from the side. Figure 11 shows the transient temperature at the inner containment vessel lid seal. It can be seen that the maximum temperature has effectively been reached after about 2 days. The maximum seal temperature is predicted to be 151°C. Figure 12 shows the predicted temperature profile at the end of the transient calculation. The maximum temperature reached by the containment vessel was 163°C. The results for the internal heat load of 5W are illustrated in Figures 13 – 15. The transient temperatures on the outer surface of the keg are shown in Figure 13. They show a similar pattern to those for the 30W heat load (see Figure 10), though with lower overall temperatures. The maximum predicted temperature is 98°C and occurs on the keg lid. The keg bottom has a maximum temperature of 72°C, whereas that of the keg side is 66°C. Figure 14 shows the transient temperature at the inner containment vessel lid seal. Once again, the


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maximum temperature is reached after a couple of days. The maximum seal temperature is predicted to be 76°C. Figure 15 shows the predicted temperature profile at the end of the transient calculation. The maximum temperature reached by the containment vessel was 78°C. The results for the internal heat load of 0W are illustrated in Figures 16 – 18. The temperatures on the outside of the keg predicted for the 0W heat load are shown in Figure 16 and very similar to those for the heat load of 5W (see Figure13). The keg lid reaches a maximum temperature of 97.5°C, the keg bottom reaches 70°C, and the keg side reaches 65°C. Figure 17 shows the transient temperature at the inner containment vessel lid seal. Once again, the maximum temperature is reached after a couple of days. The maximum seal temperature is predicted to be 60°C. Figure 18 shows the predicted temperature profile at the end of the transient calculation. The maximum temperature reached by the containment vessel was 59°C. The maximum temperatures reached by the containment vessel under normal conditions of transport, both with and without insolation, are summarised in Table 4, together with those reached under hypothetical accident conditions (see Section 5 below). The maximum temperatures reached by the various locations on the keg under normal transport conditions, both with and without insolation, are summarised in Table 5.

5 Calculation of Temperatures During the Fire Accident The finite element model has been used to determine the temperature of the container during the fire accident specified in the 10CFR71 Regulations [1]. A 30 minute, 800°C fire was simulated followed by a 12 hour cooling period. The model used for the heating phase was derived from the one used for Normal Conditions of Transport, with the following changes:

� A transient calculation was made with heating from a constant 800°C fire for 30 minutes followed by cooling for 12 hours.

� The calculation started from the temperature profile obtained for normal conditions of transport with insolation (see Section 4.2).

� No solar insolation was applied in the transient calculation during the heating phase of the fire test; however, insolation was pessimistically applied during the 12 hour cooling phase following the fire.

� During the heating phase of the fire test all exterior surfaces of the keg were assumed to receive heat by forced convection and radiation from the fire. A convection coefficient of 15W/m2/K was assumed (the value suggested in the Advisory Material for the IAEA Regulations [5]). The absorptivity of the surface of the keg was assumed to be 0.8, which is the value specified in the 10CFR71 Regulations [1].

� During the cooling phase, heat was modelled as being lost from all exterior surfaces of the keg by radiation and natural convection. The emissivity of the surface of the keg was assumed to remain at 0.8, the value used during the heating phase (pictures of the LS container after the furnace test show the surface to be blackened and oxidised). Established correlations for natural convection were again used to derive the appropriate convection coefficient (see Appendix 2).

The 10CFR71 Regulations [1] require the thermal test to be performed upon a container which has already been subjected to the regulatory impact tests. A series of impact tests, as specified by the 10CFR71 Regulations [1] have been performed upon the SAFKEG LS design. The only significant distortion or damage that occurred was to the ‘skirt’ at either end of the keg. These ‘skirts’ are not significant to the thermal performance and it is judged that the damaged ‘skirt’ would provide greater protection in a fire than an undamaged ‘skirt’ (since, when bent over, it will provide


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shielding of the top and bottom of the keg from the fire). Given the similarity between the LS and HS designs, a similar result is expected from the SAFKEG HS impact tests; the finite element mesh used to model the fire accident for the HS design was therefore unchanged from that used to model normal conditions of transport. The maximum temperatures reached at key locations on the package are summarised in Table 6. The results for the internal heat load of 30 W are illustrated in Figures 19 – 21. Figure 19 shows the predicted temperature profile at the end of the heating phase. The external surface of the keg is, as expected, close to the temperature of the fire (800°C). At this time the temperature of the inner containment vessel lid seal is around 150°C and has hardly changed from that predicted under normal conditions of transport. Figure 20 shows the predicted temperature, for the same case, on the exterior surface of the keg. As found with the SAFKEG LS, the outer skin of the keg heats up and cools down rapidly because it is insulated from the inner containment vessel by the cork. The temperature of the keg side changes more quickly than that of the lid or base because the outer shell is thinner than the lid or base and therefore has a lower thermal capacity. Figure 21 shows the predicted temperature of the inner containment vessel lid seal. The lid seal reaches a maximum temperature of 196°C after about 3 hours. Melting of the containment vessel shielding is not an issue for the SAFKEG HS because the depleted uranium used for this design has a melting point of around 1120°C. Twelve hours after the fire test, the lid seal region has cooled to 164°C. The maximum temperature experienced in the containment vessel cavity is 208°C. The results for the internal heat load of 5W are illustrated in Figures 22 – 24. Figure 22 shows the predicted temperature profile at the end of the heating phase. The external surface of the keg is close to the temperature of the fire (800°C). At this time the temperature of the inner containment vessel lid seal is around 76°C and has hardly changed from that predicted under normal conditions of transport. Figure 23 shows the predicted temperature, for the same case, on the exterior surface of the keg. The results are very similar to those from the 30W case. Figure 24 shows the predicted temperature of the inner containment vessel lid seal. The lid seal reaches a maximum temperature of 130°C after about 4 hours. Twelve hours after the fire test, the lid seal region has cooled to 108°C. The maximum temperature experienced in the containment vessel cavity is 132°C. The results for the internal heat load of 0W are illustrated in Figures 25 – 27. Figure 25 shows the predicted temperature profile at the end of the heating phase. The external surface of the keg is, as expected, close to the temperature of the fire (800°C). At this time the temperature of the inner containment vessel lid seal is around 60°C and has hardly changed from that predicted under normal conditions of transport. Figure 26 shows the predicted temperature, for the same case, on the exterior surface of the keg. The results are very similar to those from the 30W case. Figure 27 shows the predicted temperature of the inner containment vessel lid seal. The lid seal reaches a maximum temperature of 115°C after about 4¼ hours. Twelve hours after the fire test, the lid seal region has cooled to 95°C. The maximum temperature experienced in the containment vessel cavity is 115°C.


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6 Air Gap Sensitivity Analysis A short additional study was conducted to investigate the influence of the air gaps on the numerical predictions. In particular, the NCT and HAC fire test calculations, for an internal heat load of 30W, were re-run with cork assumed to have swollen to such an extent that all of the air gaps around the various pieces of cork were filled with cork rather than air. This was done by modifying the model so that heat transfer across each gap was represented by conduction through cork rather than by conduction and radiation through air. Table 7 compares the predicted maximum temperatures from the NCT calculations for the air and cork-filled gaps for both the steady state (i.e., no insolation) and transient (i.e., with insolation) calculations. At locations on the outside of the keg (e.g., the keg lid, bottom and mid height on its outer surface) the temperatures are very similar, though with the keg lid and side being marginally hotter if the gaps are cork-filled and the keg bottom about 1ºC cooler. At locations deeper within the keg and closer to the internal heat source, however, such as the containment vessel cavity and lid seal, the model with cork-filled gaps predicts maximum temperatures 5 - 6ºC cooler than the original model with air-filled gaps. This is true for both the steady state and transient simulations. This suggests that the higher conductivity cork-filled gaps have the effect of allowing the heat from the internal source to flow away from the keg centre more efficiently than when the gaps are filled with air, resulting in lower temperatures in the vicinity of the containment vessel. Figure 27 shows how the lid seal temperatures for the cork and air-filled gap models vary with time during the NCT transient simulations. The lid seal temperatures for the cork-filled gap model are consistently 5 – 6 ºC cooler throughout the entire simulation. Table 8 shows the maximum temperatures at various locations for cork and air-filled gap versions of the model for the HAC fire test simulation. At most locations the predicted maximum temperatures are the same to with a few tenths of a degree; only at the containment vessel cavity and lid seal and in the depleted uranium shielding are the temperature differences greater, with the cork-filled gap temperatures being approximately 2ºC cooler. Figure 28 shows how the containment vessel lid seal temperatures for the cork and air-filled gap models vary with time during the HAC fire test simulations. The lid seal temperature for the cork-filled gap model increases at a slightly greater rate, and reaches its peak temperature earlier, than in the air-filled gap model, indicating that if the gaps are filled with cork the efficiency of heat transfer between the outside of the cask and the containment vessel is enhanced.

7 Conclusions The thermal performance of the SAFKEG HS design of transport container has been analysed using an axi-symmetric finite element model. The model is similar to the LS design previously assessed [3]. It was validated by comparison against experimental data from a self heating test. Good agreement was obtained with the results of the self heating test but only after the thermal conductivity of the cork had been decreased by 15% compared to its measured value. Using the lower, fitted, value of the cork conductivity ensures that the model is pessimistic when calculating temperatures for normal conditions of transport. The temperature of the container has been predicted under normal conditions of transport, for heat loads of 30, 5 and 0W, both with and without solar insolation. Without solar insolation the temperature at the inner containment vessel lid seal is 135, 56 and 38ºC, respectively. For all internal heat loads, the effect of the solar insolation is to increase the temperature at the inner containment vessel lid seal; by about 15ºC in the case of the 30W heat load, and by about 20°C for both the 5 and 0W heat loads.


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The temperature of the container has been predicted during the thermal (fire) test. At a heat load of 30W, the inner containment vessel lid seal is predicted to reach a maximum temperature of 196°C. At a heat load of 5W, it is predicted to reach 130°C and at zero heat load it is predicted to reach 115°C. The main effect of having the various air gaps in the model filled with cork rather than air is slightly to enhance the efficiency of heat transfer from the containment vessel to the keg’s outside surface in the case of the NCT simulation, and in the opposite direction in the case of the HAC fire test simulation.

8 References 1. Title 10, Code of Federal Regulations, Part 71, Office of the Federal Register, Washington, DC,

2009. 2. Abaqus version 6.12-1, Dassault Systemes Simulia Corp. 3. ‘Thermal Analysis of the SAFKEG LS Design ‘, SERCO/TAS/5388/001 Issue 2, Croft Project

No 06/08/10, July 2009. 4. Croft report CTR 2010/02, Issue A Prototype SAFKEG HS 3977A/0002 NCT and HAC

Regulatory Test Report. 5. ‘Advisory Material for the IAEA Regulations for the Safe Transport of Radioactive Material’,

2005 Edition, IAEA Safety Guide No. TS-G-1.1 (ST2), 2002.


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Material Property Temperature

(ºC) Value Reference

21 14.9 W/m/K

38 15.0 W/m/K

93 16.1 W/m/K

149 16.9 W/m/K

205 18.0 W/m/K

260 18.9 W/m/K

316 19.5 W/m/K

371 20.4 W/m/K

427 21.1 W/m/K

482 22.0 W/m/K

538 22.8 W/m/K

593 23.5 W/m/K

649 24.2 W/m/K

705 25.1 W/m/K

760 25.8 W/m/K

Conductivity

816 26.5 W/m/K

1

Density - 7,900 kg/m3

2

21 483 J/kg/K

38 486 J/kg/K

93 506 J/kg/K

149 520 J/kg/K

205 535 J/kg/K

260 544 J/kg/K

316 551 J/kg/K

371 559 J/kg/K

427 562 J/kg/K

482 570 J/kg/K

538 577 J/kg/K

593 583 J/kg/K

649 585 J/kg/K

705 591 J/kg/K

760 596 J/kg/K

304 Stainless Steel

Specific Heat

816 601 J/kg/K

1

0 23.1 W/m/K Conductivity

400 32.5 W/m/K 3

Depleted Uranium with 2% Mo

Density - 18,650 kg/m3

4


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Material Property Temperature

(ºC) Value Reference

0 117.5 J/kg/K Specific heat

300 142.0 J/kg/K 3

Conductivity - See Figure 4 5

Density - 290 kg/m3

5 Cork

Specific heat - 1,650 J/kg/K 5

Conductivity - 167 W/m/K 6

Density - 2,700 kg/m3 6

Aluminium

Specific Heat - 896 J/kg/K 6

0 0.0243 W/m/K

200 0.0314 W/m/K Air Conductivity

400 0.0515 W/m/K

Table 1 – Material Properties used in the Model Table 1 References

1. ASME Section II (2001), Part D, Subpart 2, Table TCD, Group J 2. ‘Design Manual for Structural Stainless Steel (Second edition)’, The Steel Construction

Institute, Building series, Vol 3. 3. ‘For Computer Heat-Conduction Calculations: a Compilation of Thermal Properties Data’,

UCRL-50585,1969 (THERM database). 4. ‘MIDUS Transportation Package Safety Analysis Report’, Revision 1, Document No.

TYC01-1600, Docket No. 71-9320, EnergySolutions Spent Fuel Division Inc, Campbell, CA, March 2007.

5. ‘Summary of the Physical Properties and Composition of Resin Bonded Cork’, CTR 2001/11, Issue D, 2002.

6. Values for Aluminium grade 6061-T6 from www.matweb.com

Material Condition Value Reference

Internal surfaces 0.2 1

External surface – heating test & NCT 0.25 1 304 Stainless Steel

External surface – fire test 0.8 2

Cork All conditions 0.95 3

Depleted Uranium

Internal surfaces (un-oxidised) 0.31 4

Table 2 – Emissivities used in the Model Table 2 References

1. Touloukian & DeWitt, "Thermal Radiative Properties – Metallic elements and alloys", Thermophysical properties of matter, Vol 7, Pub IFI/PLENUM, 1970.

2. Title 10, Code of Federal Regulations, Part 71, Office of the Federal Register, Washington, DC, 2009.


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3. ‘The Emissivity of Various Materials Commonly Encountered in Industry’, Land pyrometers Technical Note 101.

4. ‘Heat Release Rate from the Combustion of Uranium’, Charles W. Solbrig, Argonne National Laboratory, P.O. Box 2528, Idaho.

Container orientation

Surface Insolation heat

flux (W/m2)

Lid end 800

Side 200 Vertical

Bottom end None

Lid end 200

Side 400 Horizontal

Bottom end 200

Table 3 – Solar Insolation Fluxes Applied to the Model Notes Insolation values obtained from 10CFR71 Transport Regulations [1] Insolation flux applied for 12 hours in each day The ‘side’ of the keg was taken to be the entire height of the keg plus the inside of the skirts (except the inside the bottom skirt when the container is oriented vertically)

Max CV Temperatures (ºC)

Condition 0W Heat Load 5W heat Load 30W Heat Load

Normal transport - no insolation

38.0 58.9 148.4

Normal transport – with insolation

58.8 78.1 163.2

Fire Accident 115.4 132.0 208.0

Table 4 – Predicted Maximum Containment Vessel Temperatures (ºC) for Internal Heat Loads of 0, 5 and 30W


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0W Heat Load 5W Heat Load 30W Heat Load

Location Max T (ºC)

No Insolation

Max T (ºC) with

Insolation

Max T (ºC)

No Insolation

Max T (ºC) with

Insolation

Max T (ºC)

No Insolation

Max T (ºC) with

Insolation

CV cavity 38.0 58.8 58.9 78.1 148.4 163.2

CV lid seal 38.0 59.5 56.2 76.4 135.0 151.1

Keg lid 38.0 97.5 39.7 98.4 46.5 102.4

Keg bottom 38.0 69.9 41.5 72.0 56.8 84.5

Mid height on keg surface

38.0 65.4 39.8 66.4 46.8 71.2

Table 5 – Summary of Calculated Temperatures under Normal Conditions of Transport

0W Heat Load 5W Heat Load 30W Heat Load


Time After Fire Start

(mins)

Max T (ºC)


(mins)

Max T (ºC)


(mins)

CV cavity 115.4 210 132.0 210 208.0 180

CV lid seal 115.3 254 130.1 244 196.3 210

Cork 787.4 30 787.6 30 788.2 30

DU shielding 115.3 210 130.3 210 198.2 180

Keg lid 784.3 30 784.4 30 785.0 30

Keg bottom 788.6 30 788.7 30 789.3 30


786.6 30 786.6 30 786.1 30

Table 6 – Summary of Calculated Temperatures under Hypothetical Accident Conditions


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Gaps Filled with Cork Gaps Filled with Air


No Insolation

Max T (ºC) with

Insolation

Max T (ºC)

No Insolation

Max T (ºC) with

Insolation

CV cavity 142.9 158.3 148.4 163.2

CV lid seal 129.2 145.9 135.0 151.1

Keg lid 46.8 102.3 46.5 102.4

Keg bottom 55.8 83.6 56.8 84.5


46.9 71.2 46.8 71.2

Table 7 – Comparison of Temperatures Calculated under NCT (30W) for Gaps filled with Cork and Air

Gaps Filled with Cork Gaps Filled with Air



(mins)

Max T (ºC)


(mins)

CV cavity 206.7 150 208.0 180

CV lid seal 194.1 180 196.3 210

Cork 788.3 30 788.2 30

DU shielding 196.1 150 198.2 180

Keg lid 784.8 30 785.0 30

Keg bottom 789.4 30 789.3 30


787.4 30 786.1 30

Table 8 – Comparison of Temperatures Calculated under HAC (30W) for Gaps filled with Cork and Air


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Figure 1 – The SAFKEG HS Container Design

Diagram extracted from drawing OC-5900 Issue A


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Figure 2 – The Model of the SAFKEG HS Container

Materials Grey – stainless steel Brown – cork Blue – depleted uranium


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Figure 3 – 3-Dimensional View of the Axi-Symmetric Model

Materials Grey – stainless steel Brown – cork Blue – depleted uranium


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Fig

ure

4 –

The

rmal C

onductivity o

f C

ork

0

0.0

5

0.1

0.1

5

0.2

0.2

5

0.3

0.3

5

0.4

0100

200

300

400

500

600

700

800

Tem

pera

ture

(°

C)

Thermal conductivity (W/m/K)

Measure

d v

alu

es

Used in n

orm

al tr

an

sport

calc

ula

tions

Used in f

ire a

ccid

ent calc

ula

tions


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Figure 5 – Temperature Profile in the Steady-State Heating Test

The red dots indicate the monitoring points used in Figure 6

Temp (°C)


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Figure 6 – Comparison of Measured and Predicted Temperatures in the Steady-State Heating Test

Key nn – Predicted (adjusted properties) (nn) – Predicted (reference properties) (nn) – Measured (fixed thermocouple) (nn) – Measured (surface measurement)

24 (24) (27)

31 (31) (32)

31 (31) (32)

101 (93) (100)

125 (116) (125)

124 (115) (124)

86 (79) (82)

31 (31) (31)

124 (115) (123)

42 (42) (34)

28 (28) (25)

34 (34)

(29)

31 (31) (29)

31 (31) (31)

Heat load 31W Ambient temperature 21.7°C

28 (28) (29)

30 (30) (31)

All temperatures in °C


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Figure 7 – Predicted Temperature Profile under NCT without Solar Insolation – Internal Heat Load of 30W

Temp (°C)


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Temp (°C)


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Temp (°C)



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30

40

50

60

70

80

90

100

110

0 1 2 3 4 5

Time (days)

Tem

pera

ture

(C

)

Side

Bottom end

Lid end

Figure 10– Predicted Temperature on the Outside of the Keg during NCT – Internal Heat Load of 30W


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80

90

100

110

120

130

140

150

160

0 1 2 3 4 5

Time (days)

Tem

pera

ture

(C

)

Figure 11 – Predicted temperature at the Containment Vessel Lid Seal during NCT with Insolation – Internal heat Load of 30W


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Figure 12 – Predicted Temperature Profile under NCT with Insolation – Internal Heat Load of 30W

Temp (°C)


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30

40

50

60

70

80

90

100

110

0 1 2 3 4 5

Time (days)

Tem

pera

ture

(C

)

Side

Bottom end

Lid end

Figure 13– Predicted Temperature on the Outside of the Keg during NCT –

Internal Heat Load of 5W


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0

10

20

30

40

50

60

70

80

90

0 1 2 3 4 5

Time (days)

Tem

pera

ture

(C

)

Figure 14 – Predicted temperature at the Containment Vessel Lid Seal during NCT with

Insolation – Internal heat Load of 5W


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Temp (°C)


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30

40

50

60

70

80

90

100

110

0 1 2 3 4 5

Time (days)

Tem

pera

ture

(C

)

Side

Bottom end

Lid end

Figure 16– Predicted Temperature on the Outside of the Keg During NCT – Internal Heat Load of 0W


AMEC Commercial

0

10

20

30

40

50

60

70

0 1 2 3 4 5

Time (days)

Tem

pera

ture

(C

)

Figure 17 – Predicted temperature at the Containment Vessel Lid Seal During NCT with Insolation – Internal heat Load of 0W


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Temp (°C)


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Figure 19 – Predicted Temperatures at the end of the Heating Phase of the Fire Accident – Internal Heat Load of 30W

Temp (°C)


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0

100

200

300

400

500

600

700

800

900

0 60 120 180 240 300 360

Time (minutes)

Tem

pera

ture

(C

)

Side

Bottom end

Lid end

Figure 20 – Predicted Temperature on the Outside of the Keg during the Fire Accident – Internal Heat Load of 30W


AMEC Commercial

0

20

40

60

80

100

120

140

160

180

200

0 60 120 180 240 300 360

Time (minutes)

Tem

pera

ture

(C

)

Figure 21 – Predicted Temperature of the Containment Vessel Lid Seal during the Fire Accident – Internal Heat Load of 30W


AMEC Commercial


Temp (°C)


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0

100

200

300

400

500

600

700

800

900

0 60 120 180 240 300 360

Time (minutes)

Tem

pera

ture

(C

)

Side

Bottom end

Lid end

Figure 23 – Predicted Temperature on the Outside of the Keg during the Fire

Accident – Internal Heat Load of 5W


AMEC Commercial

0

20

40

60

80

100

120

140

0 60 120 180 240 300 360

Time (minutes)

Tem

pera

ture

(C

)

Figure 24 – Predicted Temperature of the Containment Vessel Lid Seal during the

Fire Accident – Internal Heat Load of 5W


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Temp (°C)


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0

100

200

300

400

500

600

700

800

900

0 60 120 180 240 300 360

Time (minutes)

Tem

pera

ture

(C

)

Side

Bottom end

Lid end

Figure 26 – Predicted Temperature on the Outside of the Keg during the Fire

Accident – Internal Heat Load of 0W


AMEC Commercial

0

20

40

60

80

100

120

140

0 60 120 180 240 300 360

Time (minutes)

Tem

pera

ture

(C

)

Figure 27 – Predicted Temperature of the Containment Vessel Lid Seal during the Fire Accident – Internal Heat Load of 0W


AMEC Commercial

80

90

100

110

120

130

140

150

160

0 1 2 3 4 5

Time (days)

Tem

pera

ture

(C

)

Gaps filled with

air

Gaps filled with

cork

Figure 28 – Predicted Temperatures of the Containment Vessel Lid Seal under NCT for Gaps filled with Air and Cork

120

130

140

150

160

170

180

190

200

0 60 120 180 240 300 360

Time (minutes)

Tem

pera

ture

(C

)

Gaps filled with air

Gaps filled with cork

Figure 29 – Predicted Temperatures of the Containment Vessel Lid Seal for the HAC Fire Test for Gaps filled with Air and Cork


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Appendices

Contents

Appendix 1 Drawings used in model generation Appendix 1 Natural convection heat transfer correlations


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Appendix 1 Drawings used in model generation

Contents

Table of engineering drawings


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Table of Engineering Drawings The engineering drawings of the SAFKEG HS design which were used to generate the finite element model are listed in the following table:

Component Engineering Drawing Reference

Keg Body

Keg Liner

Keg Lid

0C-5901, Issue A – Keg Assembly Design No. 3977

1C-5905 Issue A – Keg Liner Assembly

3C-5917 Issue A – Keg Liner Disc

1C-5908 Issue A – Keg Top Flange

1C-5909 Issue A – Keg Base Plate

0C-5907 Issue A – Top and Bottom Rim

1C-5914 Issue A – Keg Lid

Cork

3C-5906 Issue A – Top Cork

1C-5910 Issue A – Outer Cork

1C-5903 Issue A – Inner Cork

Inner Container Body

CV Body Shielding

1C-5999, Issue A-P4 – Containment Vessel HS Body Construction

2C-5961 Issue A – CV Body Shielding

Inner Container Lid

CV Lid Shielding

1C-5997, Issue A-P3 – Containment Vessel HS Lid Construction

1C-5951 Issue A – CV Lid Machined

3C-5962 Issue A – CV Lid Shielding

Dimension check and

Air gaps 0C-5949 Issue A-P2 – SAFKEG HS Construction

Heating Block Figure 1 of CP 407 “Prototype SAFKEG-HS 3977A Procedure for the NCT Steady State Thermal Equilibrium Test”, Issue A (drawing assumed to be to scale)


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Appendix 2 Natural Convection Heat Transfer Correlations

Contents

Introduction Vertical plate Vertical cylinder Horizontal plate (upward facing) Horizontal cylinder References


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Introduction The calculations performed using the finite element model used standard correlations to determine the appropriate convection coefficient on the outside surfaces of the keg. The convection coefficients corresponding to four different geometries were required: � Vertical plate � Vertical cylinder � Horizontal plate � Horizontal cylinder The correlations for natural convection heat transfer coefficients are all taken from [1]. In each case the correlation is for the Nusselt number, and is in the form:

( ) mmturb

m

lam NuNuNu/1

+=

i.e. it combines a Nusselt number for fully laminar conditions with a Nusselt number for fully turbulent conditions using a power law rule with index m. The correlations for Nulam, Nuturb and the corresponding values for m are given below for the four geometries used in the calculations. Figure A2.1 compares the four correlations (expressed as heat transfer coefficients) for an ambient air temperature of 38°C.

Vertical plate Index m = 6 Laminar part:

)/8.21ln(

8.2

,lamP

lamNu

Nu+

=

9/416/9

4/1

,)Pr)/492.0(1(

503.0

3

4

+=

RaNu lamP

Turbulent part:

42.081.0

3/122.0

)Pr61.01(

Pr13.0

+=

RaNuturb

Vertical cylinder Index m = 6 Laminar part:

VP

lam

lam

lam

lam NuNu)9.01ln(

9.0

ξ

ξ

+=


AMEC Commercial

lamP

lamNu

DL

,

/2=ξ

)/8.21ln(

8.2

,lamP

VP

lamNu

Nu+

=

9/416/9

4/1

,)Pr)/492.0(1(

503.0

3

4

+=

RaNu lamP

Turbulent part:

)9.01ln(

9.0 ,

turb

turbPturb

turb

NuNu

ξ

ξ

+=

turbP

turbNu

DL

,

/2=ξ

42.081.0

3/122.0

,)Pr61.01(

Pr13.0

+=

RaNu turbP

Horizontal plate (upward facing) Index m = 10 Laminar part:

)/4.11ln(

4.1

,lamP

lamNu

Nu+

=

9/416/9

4/1

,)Pr)/492.0(1(

503.0

3

4835.0

+=

RaNu lamP

Turbulent part:

3/114.0 RaNu turb =

Horizontal cylinder Index m = 3.3 Laminar part:

)/21ln(

2

,lamP

lamNuf

fNu

+=

where f is a constant (set to 1)

9/416/9

4/1

,)Pr)/492.0(1(

503.0

3

4.772.0

+=

RaNu lamP

Turbulent part:


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3/1

)1002.0Pr0039.0( RaNuturb +=

References 1) W M Rohsenow et al., Handbook of Heat Transfer Fundamentals, second edition.


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Fig

ure

A2.1

– C

alc

ula

ted H

eat

Tra

nsfe

r C

oeffic

ients

for

va

rious G

eo

metr

ies a

nd O

rienta

tions

0123456789

10

0100

200

300

400

500

600

700

800

Su

rface t

em

pera

ture

(°

C)

Heat Transfer Coefficient (W/m2/K)

Vert

ical cyl

inder

Hori

zonta

l cyl

inder

Hori

zonta

l pla

te (

upw

ard

)

Vert

ical pla

te

Am

bie

nt te

mpera

ture

3

8°C

Documents

Thermal Analysis of the SAFKEG HS DesignThe heat transfer coefficients used to model convective heat transfer in the heating phase of the fire accident are described in Section 5