A paper in the proceedings of a conference on fluidization 325 16–17 November 2011, Johannesburg, South Africa

IFSA 2011, Industrial Fluidization South Africa: 325–338. Edited by A. Luckos & P. den Hoed Johannesburg: Southern African Institute of Mining and Metallurgy, 2011

Practical experiences with spouted-bed chemical vapour deposition

R. Cromarty Department of Material Science and Metallurgical Engineering,

University of Pretoria, South Africa

Keywords: spouted bed, chemical vapour deposition, CVD, silicon carbide, TRISO, coated fuel particle

Abstract—Chemical vapour deposition (CVD) has become an important process in a wide variety of industries, most notably the semiconductor industry. For most of these applications the products to be coated are large enough to be coated individually or in small batches. Few CVD processes are used to coat powders or granular materials. A notable exception to this is the process used to coat nuclear fuel particles for use in high-temperature, gas-cooled reactors. Coating of the fuel particles presents a challenge in that a fine granular material needs to be uniformly coated with a high degree of confidence in very low defect rates. The technique of choice for these coatings has been spouted-bed CVD coaters.

Aspects of operating a spouted-bed CVD coater using simulated TRISO fuel particles was investigated with a 50-mm ID coater and a 2D model of the coater. The effect of gas flow rate and bed depth was investigated using the coater and model.

This paper will consider some of the theoretical aspects of spouted beds, but will focus on results obtained from test work. Practical aspects of the coater operations will also be discussed, together with a brief summary of the coating results obtained. Finally some full-scale implementations of the CVD coating systems will be discussed along with some alternative applications.

INTRODUCTION Chemical vapour deposition (CVD) has developed into a technologically important process that is widely used. Typical applications include deposition of various dielectric, semiconducting and metallic films for the semiconductor-manufacturing industry, hard coatings for cutting tools, protective coatings for oxidation and corrosion prevention, and thermal barrier coatings for high-temperature applications. For the majority of these applications the substrate is fairly large and may be coated individually or in batches. Very few applications of CVD processes to coating of powdered or granular materials have been reported [1]. One unique application of CVD processes is the coating of fuel particles for high-temperature, gas-cooled nuclear reactors. In this application fuel kernels containing fissile material are coated with layers of carbon and silicon carbide. This coating system, referred to as a Tri Isotropic, or TRISO, coating, is depicted in Figure 1. Starting from the fuel kernels, one deposits sequentially the pyrocarbon and silicon carbide layers in a spouted bed coater. Typical coating conditions for each of the layers are given in Table 1. Total gas flow rates depend on the exact size and geometry of the coater used. A key concern when coating TRISO particles is to ensure that all particles are uniformly coated and that the coating layers are free of defects. Defect densities of less than 1 in 105 are expected for a good coating process. Defective particles can lead to the release of radioactive fission products into the reactor coolant.

Each of the layers plays a role in ensuring the integrity of the coating system during fabrication and operating; however, the silicon carbide layer is regarded as being the most

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important layer. Silicon carbide provides mechanical strength to the coating and acts as a barrier to diffusion, preventing the release of fission products at high temperatures.

Figure 1. Typical dimensions of a TRISO-coated fuel particle

Table 1. Typical TRISO-particle coating conditions

Layer Temperature (°C)

Gas mixture (volume %)

Time (minute)

Buffer 1300 argon/acetylene (33/67)

3

Inner pyrocarbon Outer pyrocarbon

1250 argon/acetylene/propylene (71/16/13)

15

Silicon carbide 1510 hydrogen/methyltrichlorosilane (98.5/1.5)

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Typically production batches of up to approximately 5 kg of fuel kernels are coated in a batch process. All layers are coated sequentially without interruption of the coating process. Fluidized beds offer an ideal way of uniformly coating the large number of particles in a batch. However, owing to the size and density of the particles, they do not fluidize easily, so the coaters are operated in a spouting mode.

A laboratory-scale coater was built in the Department of Material Science and Metallurgical Engineering, University of Pretoria, specifically to investigate the influence of process variables on the properties of the silicon carbide layer. Details of the coater are illustrated in Figure 2. From a fluidization perspective the most important components of the coater are the graphite reaction tube, gas inlet and gas feed line. The reaction tube consisted of a 50-mm ID, 70-mm OD graphite tube with a 60° conical base. Gas was fed into the tube through a graphite nozzle with a 6 mm hole. Two different gas-feed systems were investigated; one consisting of a plain alumina tube, the other of a water-cooled copper tube. Details of the two systems can be seen in Figure 3. For all test runs a zirconia (ZrO2) surrogate particle was used in place of urania (UO2) kernels. The starting material for the tests had previously been coated with a buffer carbon and inner pyrolytic carbon coating. No outer pyrocarbon coating was deposited after silicon carbide deposition.

Although the main aim of the project was to investigate the relationship between deposition conditions and silicon carbide properties this paper will focus on aspects of fluidization behaviour and some related practical problems experienced with operating the experimental coater.

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Figure 2. Main features of laboratory scale coater.

Dimension of outer vessel are ~59 0mm long × 285 mm diameter. Details of the power supply, gas supply, MTS bubbler, control system and exhaust gas cleaning system are not shown.

Figure 3. Comparison of the gas feed arrangements for the non-cooled and cooled gas feed

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TRISO PARTICLE FLUIDIZATION A feature of a TRISO coating is the high density of the fuel kernels, urania having a density of 10.4 g/cm3. By the time the particles are fully coated the density has decreased to approximately 3.3 g/cm3, whereas the diameter has increased from 500 µm to 920 µm. Another factor is the changing composition of the process gas: heating and cooling steps are carried out in argon while the composition of the gas during deposition varies as indicated in Table 1. As a result of these changes one can expect the fluidization properties of the particles to change throughout the coating process. Using the Geldart classification [2], we see that the particles would be classified as type D during the whole coating process. Even with zirconia (density of ZrO2 5.8 g/cm3) surrogate kernels the particles would be still be classified as type D; however, they lies closer to the B-D boundary. The position of the particles on a Geldart plot at various stages of processing is shown in Figure 4.

A limitation of the original Geldart plots is that they were determined for air under standard conditions. Yang [3] proposed a modification to the charts whereby dimensionless density (equation 1) is plotted against Archimedes number (equation 2). This allows for the fluidization behaviour of particles in any gas under non-standard conditions to be determined.

Dimensionless density = f

fs

ρ

ρρ )( − (1)

Ar = gd 3 ρf (ρs – ρf)/µ2 (2)

where ρf and ρs are fluid and solids densities (kg/m3); Ar is the Archimedes number; g is the gravitational acceleration (9.8 m/s2); d is the particle diameter (m); and μ is dynamic viscosity (kg/ms)

Using a modified Geldart plot, we can show particle behaviour during coating (see Figure 5). We see that with the exception of the buffer deposition step the actual coating has relatively little effect on the fluidization behaviour. Changing the composition of the gas mixtures has a large impact, especially around the silicon carbide deposition step, where hydrogen is used as a carrier gas. From Figure 5, it appears as if the particles would be classified as type B during all steps, except that of silicon carbide deposition. From practical experience we know that the particles spout, implying that they behave as a type-D powder. Two possible explanations for this discrepancy are—

• The B-D boundary, as proposed by Yang [3] is not correct. Yang focuses on the A-B boundary and did not present any data confirming the correctness of the B-D boundary using the plot of dimensionless density versus Archimedes number. For both the A-B and B-D boundary he used the original boundaries, as proposed by Geldart, but recalculated and extrapolated so as to fit the modified plots. All data presented in support of the proposed modifications were for type-A particles or type-B particles close to the A-B boundary. Alternative B-D boundaries based on Archimedes number only have been proposed. Grace [4] proposed a boundary of Ar = 145,000, whereas Goossens [5] suggested Ar = 176,900. Other values quoted by Grace ranged from Ar = 20,000 (Chen and Pei) to Ar = 100 000 (Baeyens and Geldart). From these differences it is apparent that there is considerable uncertainty regarding the B-D boundary.

• There is no clear boundary between type-B and D fluidization and that the appearance of a fountain above the bed of particles is more a function of the gas inlet geometry than the type of fluidization. This would imply that with suitable gas inlet geometry it may be possible to get type-B fluidization to appear to be type-D. Tests carried out in a Perspex model did appear to be operating in a bubbling mode at low gas flow rates and then transitioned to spouting when the gas flow was increased. This despite the fact that the particles would have been classified as type D under the test conditions. Zhou [6] investigated the spouting of type-B particles in an arrangement very similar to the coater tube. Grace [4] points out that the boundaries are not fixed limits and indicates that spouting can occur at Archimedes numbers below the limits he proposed as the B-D boundary.

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Further investigation into the transition between type-B and type-D spouting will be required before the exact nature of fluidization for these particles under typical coating conditions can be clarified.

Figure 4. Position of TRISO fuel particles on a Geldart plot.

Coating starts with particles of 500 µm diameter and 10.4 g/cm3 and ends with 920 µm particles of density 3.3 g/cm3—i.e., coating starts at the top left and ends at the bottom right of the

dashed line. For ZrO2 particles the line is shifted diagonally left and downward.

Archimedes Number

1e+1 1e+2 1e+3 1e+4 1e+5 1e+6

Dim

ensi

onle

ss D

ensi

ty (

(ps -

pf)/

p f )

1e+3

1e+4

1e+5

1e+6

Group A Group BGroup D

Process start

Process end

SiC

Buffer deposition

Inner pyrocarbonOuter pyrocarbon

Figure 5. Position of TRISO particles during coating on a Geldart plot as modified by Yang. Vertical dashed lines indicate alternative B-D boundaries as proposed by Chen (Ar = 20,000),

Geldart (Ar = 100,000), Grace (Ar = 145,000) and Goossens (Ar = 176,900).

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EXPERIMENTAL PROCEDURE Two sets of experiments were run to investigate the fluidization behaviour of the carbon-coated zirconia particles later used for coating test work. Initially tests were run using the coater; later tests were carried out using a Perspex two-dimensional model of the coater tube. Use of the Perspex model allowed a view of a cross section of the particle bed during fluidization. In both sets of tests the pressure drop across the gas-feed nozzle and particle bed was measured as a function of gas flow. All experiments were carried out at room temperature. It was originally intended to use the coater for tests at high temperature, but no meaningful results could be obtained, probably owing to interference from the induction heating system.

The Perspex model was intended for simulating a section of the coater tube so that flow within the particle bed could be seen. It was designed to simulate a 5-mm thick vertical section through the centre of the coater tube. The thickness of 5 mm was chosen as a compromise between wall friction effects and creating a two-dimensional bed. A bed thickness of 5 mm is equivalent to a thickness of at least 6.5 particle diameters. This exceeds the minimum thickness of 5 diameters quoted by Liu [7]. Various gas inlets could be inserted into the base of the model, allowing for different inlet geometries to be tested. Inlets of 3 mm, 6 mm and 4 mm with a 1 mm inlet on either side of the central inlet were used. The 4 mm + 2×1 mm inlet was intended for simulating a concentric inlet. By using only the 1mm inlets on the 4 mm + 2×1 mm inlet, we were able to test the impact of an annular gas inlet. Details of the model are illustrated in Figure 6.

For both sets of tests the gas flow rate was controlled by a mass-flow controller. Pressure was measured by means of an electronic pressure gauge and a data-capture system. The pressure sensor was calibrated against a water-filled manometer; the factory calibration of the mass-flow meters was accepted.

Each test was started using the maximum flow rate; this prevented particles from dropping out of the bed before the tests started. The flow was then decreased in steps down to the minimum flow rate, after which the flow was stepped back up to the maximum. After each step the flow was allowed to stabilize for 2 s before pressure readings were recorded at 0.5-s intervals. In the case of the coater tube the minimum flow rate was approximately 2.8 nL/min; this was the point where particles started to drop out of the bed. For the model the flow rate was taken all the way to 0 nL/min, which resulted in the inlet being filled with particles.

It was found that when the gas flow was being increased the mass-flow controller did not respond correctly at flow rates lower than 1 nL/min. Pressure and flow reading for flows lower than 1 nL/min were recorded, but not included in the analysis. This was only a problem when using the model, as the flow rate with the coater tube was always greater than approximately 2.8 nL/min.

For tests with the coater tube, in addition to measuring pressure drop across the particle bed, we recorded the gas flow required for the following events:

1. Particles start to drop out of the bottom of the tube. For the coater tube this represented the minimum gas flow for each test

2. There are the first signs of movement of the particle bed 3. The particle bed starts to circulate. At this point the fountain has not yet formed, but

there is enough movement of the bed to circulate slowly 4. A fountain forms. At this stage the bed is behaving as a spouted bed with a fountain of

particles above the bed

As the bed had to be observed from the top looking down the coater tube, it was not always possible to determine these events accurately. Occasionally some of these events were not seen at all. When running tests using the model the behaviour of the bed could be filmed. This made assessment of the fluidization behaviour much easier.

Various particle loadings were tested for each set up. In the case of the coater tube the load was varied by weight, as it was difficult to measure bed height accurately, and bed weight rather than bed depth was later to be used as an experimental variable. A load of approximately 40 g filled the conical section of the tube. For the model the fill height was varied; a fill height of 40 mm filled the conical section of the model (i.e., a 40 mm fill height in the model was equivalent to 40 g in the coater). To prevent particles dropping out through the

331

gas inlet particles were loaded while gas was flowing. The pressure drop across the gas-feed line and inlet nozzle were measured by conducting tests without any particles.

Figure 6. The coater model used to investigate particle flow patterns

The geometry of the gas-feed nozzle could be altered by using various inserts as shown on the right. For the insert used for testing concentric gas feed, tests were carried out using the 4-mm together

with the two 1-mm inlets, as well as with only the two 1-mm inlets.

COATER-TUBE TEST RESULTS A typical plot of pressure versus gas flow rate is shown in Figure 7. Gas flows marked as “first movement”, “bed circulates” and “fountain forms” were determined by visual observation of the top of the particle bed and describe bed behaviour for increasing gas flow. Flow rates for these points were similar for increasing and decreasing gas flow. At flow rates greater than the minimum required for spouting there is no difference in pressure drop for increasing and decreasing flow rates. The flow rate at the point where the pressure drop for increasing and decreasing flow begins to differ (e.g., 15 nL/min in Figure 7) probably corresponds to the point where an internal spout forms. Several features of the plots are surprising:

1. Decreasing flow results in a single jump in pressure, but not at the flows where it would be expected. For decreasing flow it would be expected that there would be an increase in pressure at the point where spouting stopped. This should be at a flow slightly lower than that required for the fountain to form with increasing flow. Comparing Figure 8A and 8B, we can see that the pressure jump occurs at lower flows than expected.

2. For all particle loads, except 5 g and 10 g, increasing flow results in two distinct steps in the pressure-versus-flow curve (see Figure 8A). We expected pressure to increase with increasing flow and then have a single step at the point where spouting started. It is assumed that the pressure drop at lower flows is associated with the formation of an internal spout.

3. For increasing flow the pressure drop at lower flows is larger than the pressure drop associated with the onset of spouting. We would expect the onset of spouting to result in a more significant pressure drop than other events. Also, the pressure at a low-flow pressure drop appears to be unstable, jumping between a higher and lower level, as we see from the highest and lowest pressures in Figure 7.

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Flow Rate (slm)

0 2 4 6 8 10 12 14 16 18 20 22

Pre

ssur

e D

rop

(Pa)

0

500

1000

1500

2000

2500

3000

3500

Decreasing FlowIncreasing FlowNo Load

First movement

Bed circulates

Fountain forms

Figure 7. Typical trace of pressure drop across gas inlet and particle bed for decreasing and increasing

flow rates compared with the pressure drop across the gas inlet with no load Gas flow required for particle movement, bed circulation, and the formation of a fountain of

particles is based on visual observation of the top of the particle bed.

A

0 2 4 6 8 10 12 14 16 18 20 22

Pres

sure

Dro

p (P

a)

0

500

1000

1500

2000

2500

10g20g 30g 40g0 g

B

Flow Rate (slm)

0 2 4 6 8 10 12 14 16 18 20 22

Pre

ssur

e D

rop

(Pa)

0

500

1000

1500

2000

2500

10g20g30g40g0g

Figure 8. Pressure difference versus flow for (A) decreasing flow and (B) increasing flow

for particle beds of 10 g, 20 g, 30 g and 40 g.

333

Pressure and flow measurements required for internal spout formation as a function of particle load are shown in Figure 9A. Flow rates required for the onset of movement of the particle bed surface, circulation of the particle bed, and the formation of a fountain as a function of particle load are shown in Figure 9B. For the range of particle loads tested there is a good linear fit (R2 ≥ 0.98) for all these parameters.

In Figure 9 B values marked as “drop” are the minimum flows required to prevent particles from dropping out of the bed. At flows slightly below this minimum the particles tend to drop individually rather than rapidly stream out. The minimum flow required to prevent particles from dropping out of the bed is not dependent on the particle load. This is to be expected as it depends on the gas velocity in the feed tube rather than on the flow through the particle bed.

A

0 10 20 30 40 50

Pre

ssur

e D

rop

(Pa)

0

200

400

600

800

1000

1200

1400

1600

Flow

Rat

e (s

lm)

4

6

8

10

12

14

16

B

Load (g)

0 5 10 15 20 25 30 35 40

Flow

Rat

e (s

lm)

2

4

6

8

10

12

14

16

18

20DropMovement Circulation Fountain

Figure 9. (A)—Flow rate and pressure drop required to form an internal spout as a function of particle

load. Values read from pressure-versus-flow plots for each load tested. (B)—Flow rate required for onset of spouting as indicated by movement of the surface of the particle bed,

circulation of the particle bed and the formation of a fountain of particles. Also indicated is the minimum flow required to prevent particles from dropping out of the

coater, marked as “drop”.

MODEL TEST RESULTS Pressure drop versus gas flow relationships for the model differed significantly from those measured using the coater tube. A typical example of data obtained using the model is shown in Figure 10. Compared with the coater tube the model had two notable differences. First, with increasing gas flow there is only one point where the pressure drop significantly decreases with increasing flow. Secondly, no increase in pressure drop was seen with decreasing flow—in all tests the pressure drop steadily decreased with decreasing flow.

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In the case of increasing flow, the flow at the decrease in pressure drop (i.e., the first point where the pressure values for increasing and decreasing flow are similar) corresponds to the start of bubbling. As the gas flow increases beyond this point the rate of bubbling increases. As a single stream of rapidly rising bubbles forms above the gas inlet it is sometimes difficult to distinguish between bubbling and spouting. While the particle bed is bubbling the fountain consists of a rapid series of bubbles bursting out the top of the bed; as a result the fountain height is very irregular. This variation decreases as the operating mode tends to true spouting. For bed depths of 50 mm and 60 mm bubbling clearly continues up to the highest flow rates tested. For bed depths of 20 mm and 30 mm a more uniform fountain forms.

At the point where the bed starts bubbling, or spouting, particles are circulated around the spout but not all the way to the outer edge of the particle bed. This results in dead zones with little particle movement on the outer edge of the bed. As the gas flow increases the dead zones shrink until the entire bed is circulated. At higher flow rates the shallow beds tend to become unstable and oscillate from side to side.

Flow Rate (slm)

0 1 2 3 4 5 6

Pre

ssur

e D

rop

(Pa)

0

1000

2000

3000

4000

5000

6000

Decreasing flowIncreasing flowNo load

Figure 10. Pressure drop versus gas flow measured in the model

Fill height, 40 mm; inlet, 3 mm.

Increasing the bed depth resulted in an increase in the flow required to initiate spouting as well as an increase in the maximum pressure drop across the bed. Pressure drop values for the 3 mm gas inlet are given in Figure 11. From Figure 12 it is seen that, over the range of bed depths tested, the flow required to initiate spouting varies approximately linearly with bed depth for all gas inlet geometries tested.

Direct comparison of the fluidization behaviour of the various inlets is complicated by the differences in pressure drop across the inlets under no-load conditions. To overcome this, the difference between various bed heights and no-load conditions were compared. A typical example is shown in Figure 13. The significant difference between the plain inlets (3 mm and 6 mm) and the other two inlet types was seen for all bed heights except 20 mm, in which case only the 3 mm and 6 mm inlet showed a clear transition.

GAS INLET COOLING Initially the coater was designed without any cooling of the gas inlet. This was deemed to be a much simpler approach than using a cooled inlet. During operation it was found that the inlet would very quickly clog owing to the build up of deposits in the gas feed tube and inlet. This problem was reduced by making use of a shorter, smaller diameter inlet as depicted in Figure 3. Even with the modified design it was found that deposits still built up within the feed tube and inlet but not to the extent that the inlet became completely blocked before the end of a test run. To completely eliminated deposition within the inlet it was found necessary to make use of a

335

water-cooled gas feed tube and inlet as shown in Figure 3. Investigation of the impact of gas feed temperature ultimately became an important aspect of this study.

Flow Rate (slm)

0 1 2 3 4 5

Pre

ssur

e dr

op (P

a)

0

2000

4000

6000

8000

10000No load20mm30mm40mm50mm60mm

Figure 11. Variation in pressure drop with increasing bed depth for 3 mm inlet

Maximum values of 10,000 Pa may actually be higher as the pressure sensor was set up for a maximum pressure of 10,000Pa.

Fill Height (mm)

20 30 40 50 60

Flow

(slm

)

1.5

2.0

2.5

3.0

3.5

4.0

3mm 6mm Concentric 2X1mm

Figure 12. Flow required to cause a decrease in pressure for increasing flow

for various gas inlet geometries. Fill height, 40 mm

For both the cooled and the non-cooled options the same gas inlet was used. Owing to the build up of deposits within the non-cooled inlet it was decided to make use of a single 6-mm-diameter inlet rather than multiple smaller or concentric inlets. This, rather than any fluidization behaviour, was ultimately what decided the final configuration of the gas inlet. It was also decided to use the same inlet geometry for the cooled and non-cooled gas feed so as to facilitate direct comparison of the two options.

Although analysis of the test results has not yet been completed it appears that use of a cooled gas inlet had virtually no influence on the deposition rate and deposition efficiency. Crush strength of the particles coated using a cooled gas inlet is higher than those coated with no cooling of the inlet. As can be seen in Figure 14 this effect is more pronounced at higher temperatures.

336

Flow Rate (slm)

0 1 2 3 4 5

Pre

ssur

e di

ffere

nce

(Pa)

0

1000

2000

3000

40003mm 6mm Concentric2 X 1mm

Figure 13. Impact of gas-inlet geometry on pressure drop versus flow

Pressure difference values are the differences between the pressure drops measured with a bed depth of 40 mm and no load.

Temperature (°C)

1200 1250 1300 1350 1400 1450 1500 1550 1600

Cru

sh L

oad

(g)

5500

6000

6500

7000

7500

8000Cooled inletHot inlet

Figure 14. Crush load versus deposition temperature for hot and cooled inlet

Crush load values corrected to a silicon carbide thickness of 35 µm. Values are the average crush load for all tests at each temperature; other variables were MTS concentration

and total gas flow. Crush tests were conducted with soft aluminium anvils.

OTHER APPLICATIONS Spouted-bed CVD applications for TRISO particle coating have been widely investigated. For a number of smaller research coaters simple gas inlet arrangements similar to those used in this study were used [8],[9]. Other researchers have used more complex gas inlet arrangements with multiple inlets [10] or even porous plates instead of inlet ports [11]. For large-scale production coaters it is common to make use of gas inlets with multiple ports [12].

Although direct application of this technology to nuclear-fuel production still remains uncertain, other applications have been reported. An application that has reached commercial scale production is the production of solar-grade polysilicon, where a spouted-bed CVD is used in place of the tradition Siemens process.

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CONCLUSIONS Pressure drop versus flow measurement on the coater yielded unexpected results, in that a double drop in pressure was observed when flow was increased. But, for low flow rates, the pressure drop versus flow rate curve for increasing and decreasing flow was as predicted. No fluidization was observed at low flow rates. Spouting only started at flow rates significantly higher than the measured minimum fluidization flow rates—close to the flow rates where the second pressure drop was measured. It is assumed that the pressure drop observed at low flow rates was as a result of the formation of an internal spout, or cavity, within the particle bed. This effect was observed when using the 2D coater model; however, the double pressure drop was not measured on the model.

As expected, the flow rate required for spouting to start was dependent on the particle loading in the coater. This effect was measured on the coater tube and in the 2D model.

When testing the 2D model no minimum fluidization flow could be measured when flow rates were decreased. For increasing flow a single pressure drop was observed. The magnitude of the pressure drop was dependent on the type of gas inlet used. For 3-mm and 6-mm inlets the maximum pressure was substantially higher than for the 4+2×1-mm and 2×1-mm inlets. This may be related to the simple inlets getting filled with particles at low flow rates while few particles dropped into the 1mm inlets.

A flow rate substantially above the flow required for a fountain to form may be required to ensure that all the particles are circulated effectively through the bed. At low flow rates a stagnant layer of particles will form along the outer edge of the bed.

Pressure drop versus flow behaviour of the coater tube and the model were significantly different. In particular, the model did not exhibit the double drop in pressure observed when testing the coater. This brings into doubt the applicability of two dimensional models of three dimensional beds. Possible reasons for the differences between the coater and the model are the friction effects of the walls and the differences in inlet area and geometry.

Practical considerations beyond only the fluidization behaviour may govern the design of the coating system. In this particular case clogging of the gas inlet necessitated a redesign of the gas inlet. This may have an impact on the properties of the CVD deposits.

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May. 1973. 2. W.-C. Yang, “Modification and re-interpretation of Geldart’s classification of powders,”

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phase suspensions,” The Canadian Journal of Chemical Engineering, vol. 64, no. 3, pp. 353-363, Jun. 1986.

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5. J. Zhou, D. D. Bruns, C. E. A. Finney, C. S. Daw, S. Pannala, and D. l McCollum, “Hydrodynamic correlation with experimental results from cold mockup spouted beds for nuclear fuel particle coating,” in Annual Meeting - AACE International, 2005.

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7. B. J. Choi, D. W. Park, and D. R. Kim, “Chemical vapour deposition of silicon carbide by pyrolysis of methylchlorosilanes,” Journal of materials science letters, vol. 16, no. 1, pp. 33–36, 1997.

8. E. H. Voice and D. N. Lamb, DPR0677 The deposition and structure of pyrolytic silicon carbide, no. November. 1969.

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9. E. López-Honorato, P. Meadows, J. Tan, and P. Xiao, “Control of stoichiometry, microstructure, and mechanical properties in SiC coatings produced by fluidized bed chemical vapor deposition,” Journal of Materials Research, vol. 23, no. 6, pp. 1785–1796, 2008.

10. D. P. Stinton and W. J. Lackey, “Effect of deposition conditions on the properties of pyrolytic SiC coatings for HTGR fuel particles,” American Ceramic Society Bulletin, vol. 57, no. 6, pp. 568-573, 1978.

11. J. T. Keeley, B. L. Tomlin, W. C. Richardson, C. Barnes, and D. Marshall, “Development of a continuous process for TRISO coating of AGR fuel,” in 4th International Topical Meeting on High Temperature Reactor Technology HTR2008, 2008, pp. 1-6.