Ind. Eng. Chem. Res. 2007, 46, 1843-1850
1843
Impact of the Circulating Fluidized Bed Riser on the Performance of a Loopseal Nonmechanical Valve Esmail R. Monazam,‡ Lawrence J. Shadle,*,† and Joseph S. Mei† National Energy Technology Laboratory, U.S. Department of Energy, 3610 Collins Ferry Road, Morgantown, West Virginia 26507-0880, and REM Engineering SerVices, PLLC, 3537 Collins Ferry Road, Morgantown, West Virginia 26505
Most advanced coal-fuel power systems require the transfer and control of solids between two or more vessels. In many instances, the key to a successful process operation is how well the solids transfer and control system has been designed. This is particularly true in a transport gasifier and circulating fluidized bed (CFB) combustors, which are dependent upon the rapid and reliable circulation of solids to maintain a constant solids concentration in the CFB. Proper design and operation of solids returning systems are essential to the performance and operation of CFB combustion systems. An experimental investigation was conducted at the National Energy Technology Laboratory (NETL) of the U.S. Department of Energy (DOE) to study the flow and control of a light material (cork), which has a particle density of 189 kg/m3 and a mean diameter of 812 µm, through a nonmechanical valve, or loopseal, in a 0.3 m diameter CFB cold model. Fluidizing this material in ambient air approximates the same gas:solids density ratio as coal and coal char in a pressurized gasifier. The loopseal is composed of the lower section of the standpipe, an upward-flowing fluidized-bed section, and a downwardly angled overflow tube which is connected to the desired return point at the bottom of the riser. In the nonmechanical valve, both the standpipe and the fluidized-bed up-flow section of the loopseal were aerated and fluidized with air, respectively. The objective of this study was to investigate the effects of standpipe aeration, loopseal aeration, solids inventory, and superficial gas velocity through the riser on the flow rate of circulating solids. A correlation that predicts the solids flow rate as a function of these variables was developed. Comparison of the correlation with the experimental data is discussed. Pressure drop across the fluidized-bed up-flow section of the loopseal was found to increase slightly with the solid flow rates. Introduction Circulating fluidized beds (CFB) have been commercially utilized in the petroleum fluidized catalytic cracking, coal combustion, coal gasification, and Synthol processes. In general, a CFB consists of a riser, cyclones, standpipe, and solids recycle and feed system. A standpipe is a vertical tube which transfers the solids from a region of low pressure (the cyclone dip leg) to a region of high pressure (the bottom of the riser). The standpipe also provides a seal against the back-flow of gas. The solids flow rate is usually controlled by the installation of either a mechanical or nonmechanical valve at the lower end of the standpipe. Mechanical valves such as screw feeders, rotary feeders, and slide valves cannot be easily employed under high temperature and pressure conditions due to the complex geometrical and moving parts inside the feeder. Therefore, nonmechanical valves (loopseal, L-, J-, and V-valves) are usually used to control the solids flow rate by aeration, especially for process applications under elevated temperature and pressure. Extensive work on characterizing nonmechanical valves such as the L-valve,1-3 J-valve,4 V-valve,5 and loopseal6 have been reported in the literature. However, there is very little data available in literature on the interactions between the riser and the nonmechanical valve and, in particular, how the nonmechanical valve parameters have an impact on the performance of the riser. Geldart and Jones1 carried out an extensive experimental study using three different types of sand particles, group B * To whom correspondence should be addressed. Tel.: (304) 2854647. Fax: (304) 285-4403. E-mail:
[email protected]. † U.S. Department of Energy. ‡ REM Engineering Services.
powders, in L-valves of up to 100 mm in diameter. They proposed correlations relating the solids flux to both the aeration gas and the pressure drop across the L-valve. Daous and AlZahrani2 carried out an experimental study using L-valves of three different diameters of 25, 36, and 50 mm, and accordingly correlated solids flow rate to the pressure drop across the L-valve, L-valve diameter, and gas aeration. Knowlton and Hirsan4 investigated the effects of several parameters such as the dimensions, particle size, and the location of the aeration tap on the performance of a J-valve. These authors concluded that the length of the downcomer affected the maximum solids flow rate. Leung et al.5 developed a quantitative procedure for design and analyzed the V-valve system. Kim and Kim6 studied the effects of particle size and densities on solids recycle characteristics of the loopseal. The pressure drop across the loopseal was correlated to solids flux, particle size, and the density ratio of gas and solids. In this study, a cold flow circulating fluidized bed was operated with a nonmechanical valve, a loopseal to control the solids flow rate at ambient temperature. The objective of this study was to determine the influence of standpipe aeration, loopseal aeration, solids inventory, gas density, and riser superficial gas velocity on the solids flow rate. A correlation that predicts the solids flow rate as a function of these variables is proposed. Experimental Section The experimental setup used in this study is shown in Figure 1. It consists of a 0.3 m diameter, 15.45 m high riser and 0.25 m diameter, 11.4 m tall standpipe, which connects the primary cyclone to the riser through a nonmechanical valve, the loopseal.
10.1021/ie0606486 CCC: $37.00 © 2007 American Chemical Society Published on Web 02/10/2007
1844
Ind. Eng. Chem. Res., Vol. 46, No. 6, 2007 Table 1. Bed Material (Cork) Properties Fs, kg/m3 Fb, kg/m3 b dp50, µm dsv, µm Ut, µm/s Umf, µm/s emf φ
Figure 1. Schematic of CFB test unit.
189 95 0.45 1280 1007 1.02 0.15 0.50 0.69
velocities, and bed heights indicated that this constant voidage estimate is reasonable over the range of operating conditions reported here. An aeration port located near the base of the standpipe approximately 0.4 m above the base of the loopseal was used to control the solids flow rate. The aeration air was commonly referred to as the “move” air because it was found to directly impact the solids flow rate. Riser velocities were corrected for temperature and pressure as measured at the base of the riser. The relative humidity was maintained between 40% and 60% to minimize the effect of static charge buildup on the solids. In this study, a Geldart type B granulated cork bed material was used. Cork material was selected to generate data relevant for advanced high-pressure coal conversion processes. According to Buckingham-Pi analysis of the riser in a CFB the density ratio of gas to solids is a critical factor, important in scaling from a model, such as these cold flow tests, to a prototype application, such as a high pressure and high temperature transport reactor.8 Cork offers an excellent bed material which when tested at ambient conditions in air yields a density ratio similar to that of coal converted at 10-20 atm and 1000 °C. The material properties are presented in Table 1. The particle density was measured using water displacement, taking care to wet the surface completely. The cork surface is sufficiently hydrophobic to avoid filling any porosity with water. The particle size was measured using standard sieve analysis. The minimum fluidization velocity was measured in a small 0.064 m diameter, 0.203 m high fluid bed by increasing the gas velocity. The terminal velocity was calculated from drag laws9 using the measured solids density and particle size, and an estimated sphericity. Results and Discussion
Figure 2. Schematic of the loopseal.
The schematic diagram of the loopseal is shown in Figure 2. It is composed of a downcomer (A-B), a horizontal section (B-C), and a fluidized-bed section of the loopseal (C-D) that is 1.5 m high and has a diameter of 0.25 m. On-line measurement of solids circulation rates was determined continuously using a twisted fiber glass vane, the spiral, which was located in the packed-bed region of the standpipe 2.85 m above the outlet leading to the riser. A detailed discussion of the development, physical description, specification, bench calibration, and operation of the spiral is given elsewhere.7 This calibrated volumetric measurement was converted to a mass flux using the measured packed-bed density presented in Table 1, with the assumption that the packed-bed void fraction at the point of measurement is constant (i.e., b ) 0.45). Analysis of the standpipe pressure profile, estimated relative gas-solids
A set of statistically designed experiments was conducted in the CFB cold model in order to evaluate the interactions of operating parameters between the CFB and the nonmechanical valve. For these tests, four independent variables were examined: superficial gas velocity (Ug), move aeration at the base of the standpipe (Umove), loopseal fluidization velocity (Ulpsl), and total solids inventory in the system (M). The tests were conducted using a central composite statistical design testing the four independent variables over five levels each. The tests were conducted over a 2 day time period, and the date was the blocked variable. This matrix was fully randomized with two center points taken each day to provide sufficient degrees of freedom in order to estimate the experimental error. The independent variables for the experimental design are listed in Table 2. The pressure drop across the riser and loopseal, the standpipe height, and the circulation rate were the dependent variables. The levels selected for the superficial gas velocity were chosen to stay above the upper transport velocity as defined by Monazam et al.10 In this operating regime the riser is characterized as either dilute or core annular flow depending upon the
Ind. Eng. Chem. Res., Vol. 46, No. 6, 2007 1845 Table 2. Statistically Designed Experiment for the Investigation of a Loopseal independent parameters
dependent variables
run
block
M, kg
Umove, m/s
Ug, m/s
Ulpsl, m/s
1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28
1 1 1 1 1 1 1 1 1 1 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2
70.3 70.3 59.0 70.3 81.6 70.3 70.3 70.3 70.3 70.3 64.9 64.9 75.7 70.3 75.7 64.9 70.3 75.7 75.7 64.9 75.7 64.9 64.9 75.7 75.7 64.9 64.9 75.7
0.14 0.14 0.15 0.15 0.14 0.15 0.15 0.07 0.14 0.22 0.18 0.11 0.18 0.15 0.18 0.11 0.14 0.18 0.11 0.11 0.18 0.18 0.18 0.11 0.11 0.18 0.11 0.11
4.71 4.72 4.73 4.73 4.71 3.80 4.73 4.78 5.60 4.65 5.11 5.16 5.10 4.71 5.08 4.32 4.69 4.23 4.31 5.15 4.25 4.24 5.08 4.29 5.16 4.26 4.31 5.16
0.34 0.08 0.21 0.21 0.21 0.21 0.21 0.22 0.21 0.21 0.15 0.28 0.15 0.21 0.27 0.15 0.21 0.27 0.15 0.15 0.15 0.27 0.27 0.28 0.15 0.15 0.28 0.28
solids circulation rate. The range was chosen to vary the velocity over as wide a range as possible. The levels used for the move aeration were chosen to span a wide range of superficial velocities across the standpipe from 50% to 150% of the minimum fluidization velocity. In this way the standpipe was operated from clearly defined packed regime up to a rapidly moving bed. This spans as wide a range as possible in circulation rates without creating unpredictable bubbling or expansion in the standpipe. The loopseal aeration was varied from 50% to 200% of the superficial minimum fluidization velocity within its lift leg. This range of aeration is generally considered a safe range to maintain fluidization while also trying to minimize the aeration. In a gasifier loopseal aeration may be a diluent such as an inert gas that is to be avoided. During operations this flow through the loopseal is augmented by aeration being carried with the solids. For that reason it was decided to also investigate the behavior below minimum fluidization. The total solids inventory was chosen to provide a standpipe bed height that would generate sufficient pressure head for high circulation rate without reducing the bed height enough to affect the solids circulation rate measurement. As a result, the solids bed level was maintained at a level well above the spiral. Over this range of conditions the standpipe bed height always operated within the linear region of the Ergun equation and never approached the pressure differential of a fluidized bed. The analysis of variance (ANOVA) was conducted using a statistical model including the four independent variables as the main factors, the quadratic terms for each, the six two-way interactions for each of these independent parameters, and the blocked variable to estimate the dependent parameters. This test design allowed 12 degrees of freedom to estimate the error, and as is commonly done, higher level interactions (three- and fourway) were confounded with this error estimate. Typical results of this analysis for solids circulation rate are shown in Table 3. The configuration and geometry of the CFB and the bed materials will undoubtedly affect the coefficients in an empirical
Gs,
kg/(s‚m2) 10.7 8.0 8.7 9.4 9.9 7.6 9.1 3.5 9.4 14.5 11.8 6.9 12.6 9.2 13.7 5.6 9.4 12.4 6.2 5.8 10.9 10.4 12.7 7.0 6.5 9.9 6.1 7.2
∆Plpsl, kPa
∆Pr, kPa
Hstp, m
1.36 1.59 1.49 1.52 1.52 1.37 1.53 1.46 1.56 1.60 1.54 1.43 1.57 1.50 1.52 1.54 1.52 1.50 1.55 1.58 1.60 1.50 1.55 1.47 1.59 1.59 1.47 1.48
1.81 1.25 1.39 1.57 1.71 2.43 1.54 0.32 1.15 2.64 1.71 0.94 1.88 1.61 2.10 1.09 1.67 2.84 1.26 0.76 2.52 2.39 1.91 1.50 0.81 1.94 1.14 0.97
10.1 9.1 11.9 9.6 11.9 7.5 9.6 10.6 10.1 8.0 7.7 9.3 10.5 9.6 11.0 7.2 9.6 8.9 9.4 9.3 8.4 7.2 8.2 9.9 11.5 7.7 7.7 12.0
correlation. Various characteristic fluid properties and dimensionless parameters were measured and are presented in an attempt to allow one to scale these findings. Unfortunately, the varying of all of the relevant diameters, heights, or particle and fluid properties was beyond the scope of this study. Certainly the use of overall inventory is one particular property whose influence will vary depending upon the relative volumes of the riser, nonmechanical valve, and standpipe. For example, the influence of the solids inventory can be related to the height of the solids column above the loopseal valve, which in turn affects the pressure at the valve. A larger relative volume in the standpipe will reduce the influence of inventory, while a smaller relative volume increases it. Solids Circulation. Solids flow through a nonmechanical valve is usually controlled by aeration air (Umove) and the geometrical shape of the pipe. Before the aeration air is turned on, the solids form a packed bed in the standpipe and no solids circulation occurs. When the aeration flow is added, gas flows through the particles and relative gas-solids movement produces a drag force on the particles in the direction of flow. When this drag force exceeds the force required to overcome the resistance to the solids moving through the constricting bend, gravity of the particles, and the pressure in the riser, the solids begin moving through the nonmechanical valve.11 The solids do not begin to flow immediately upon injecting aeration gas into the bottom of the standpipe. There is a threshold aeration rate at which there is enough gas to produce a drag force sufficient to initiate solids flow. At this point, increasing the aeration rate causes the solids flow rate to increase and decreasing the aeration rate causes the solids flow rate to decrease. The threshold volumetric aeration rate was determined from a series of standpipe aeration ramps conducted for different riser superficial gas velocities (Ug) at a fixed inventory of cork material (Figure 3). Figure 4 is a typical plot showing how the pressure drop at the bottom of the riser (∆P873) varied with a monotonic increase in standpipe aeration for different riser gas
1846
Ind. Eng. Chem. Res., Vol. 46, No. 6, 2007
Table 3. ANOVA for Solids Circulation Ratea
a
source
type I sum of squares
df
mean square
F
significant
corrected model intercept block Ug Umove Ulpsl M UgUg UmoveUmove UlpslUlpsl MM UgUmove UgUlpsl UgM UmoveUlpsl UmoveM UlpslM error total corrected total
65 956 598.6 780 827 537.3 6720.7 2 064 756.7 59133 775.5 2 139 490.2 1 324 248.2 309 585.2 41 394.6 1650.2 42.4 712 758.1 5513.1 66 951.6 14 101.6 114 075.1 21 535.6 311 908.1 847 096 044.0 66 268 506.7
15 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 12 8 7
4 397 107 780 827 537 6721 2 064 757 59 133 775 2 139 490 1 324 248 309 585 41 395 1650 42 712 758 5513 669 52 141 02 114 075 21 536 25 992
169.17 3004 0.26 79.44 2275. 82.31 50.95 11.91 1.59 0.06 0.00 27.42 0.21 2.58 0.54 4.39 0.83
0.000 0.00 0.620 0.000 0.000 0.000 0.000 0.005 0.231 0.805 0.968 0.000 0.653 0.134 0.476 0.058 0.381
Computed using R ) 0.05, resulting R2 ) 0.995 (adjusted R2 ) 0.989).
Figure 5. Solid flux as function of the ratio of Umove/Umf. Figure 3. Linear increase of the threshold volumetric aeration rate with increasing riser gas velocity for the light material, cork.
and higher resistance of both gas and solids flow into the riser. This demonstrates that there were factors other than standpipe aeration, in this case superficial riser velocity, influencing solids flow characteristics out of the loopseal. The relationship between the mass flux of solids through the loopseal and the ratio of Umove/Umf is shown in Figure 5. The solids mass flux (Gs,v) through the loopseal increased linearly with the increasing ratio of Umove/Umf. The increase in mass flux with the ratio of Umove/Umf can be attributed to higher gas flow in the standpipe entraining and moving more solids into the nonmechanical valve. Based on these data and using a similar form used in past studies,1,3,5 the solid mass flux (Gs,v) can be correlated as
Gs,v ) 13.1 Figure 4. Pressure drop at the bottom of the riser as a function of aeration rate for different gas velocities.
velocities. The air flow rate was ramped in the standpipe from 0 to 7.0 m3/h at a rate of 0.56 m3/h per minute. From inspection of Figure 4 a threshold volumetric aeration rate was identified that increased linearly with increasing riser superficial gas velocity. Higher riser gas velocities created a higher pressure drop in the riser, with the crossover piping and cyclone requiring additional gas to offset the resulting gas compression in the standpipe. As a consequence, a higher standpipe aeration rate was required to overcome the gas compression in the standpipe
( ) Umove Umf
1.26
(1)
The mean square and the mean relative errors between the measured solids mass flux and the predicted values calculated from eq 1 are 1.3 kg/(m2 s) and (8.0%. The mean relative errors give an indication of the percentage deviation between measured and calculated solids mass flux. This correlation explains 89% of the variance in the data set; however, from an inspection of Figure 5 it was apparent that there was a lot of variability not captured by this simple expression. Near the center of the data matrix a given aeration rate must predict the circulation rate that experimentally varied from 10.9 to 15.4 kg/(m2 s). The relationship between the mass flux of solids through the loopseal was improved by adding the ratio of Ulpsl/Ut to eq 1
Ind. Eng. Chem. Res., Vol. 46, No. 6, 2007 1847
Figure 6. Solid flux as function of the ratios of Umove/Umf and Ulpsl/Ut.
Figure 8. Linear dependence of solids flux on mass inventory within the CFB loop at fixed levels of Umove, Ulpsl, and Ug. The vertical bars reflect the range of values around the average responses depicted by lines for different levels of Umove/Umf.
Figure 7. Interaction of standpipe move aeration with riser gas velocity on solids flux at a different inventory levels, M. The data on Ulpsl variation are included but unlabeled, appearing as duplicates for the M ) 65 and 75 lb cases.
(Figure 6) using the following expression:
Gs,v ) 17.4
( ) ( ) Umove Umf
1.26
Ulpsl Ut
0.17
(2)
Equation 2 explains 92% of the variance in the data set. The analysis of variance with respect to the circulation rate showed that all the main effects were significant to the 95% confidence limit including the superficial gas velocity, standpipe aeration, loopseal fluidizing air flow, and total solids inventory in the CFB. The influence of riser flow on the circulation rate was nonlinear, and its effect was dependent upon the amount of standpipe aeration (Figure 7). The quadratic term of superficial riser gas velocity and the interaction between the standpipe aeration with the riser superficial gas flow were both statistically significant. This nonlinear dependence was not observed with any of the other main factors. For example, the solids flux was linearly dependent on total solids inventory, increasing with increasing solids load in the CFB as shown in Figure 8. The magnitude and nature of the interactions between riser gas velocity and move aeration can also be discerned from Figure 7. As the standpipe aeration increased, the gain in circulation rate increased per unit riser flow. At low standpipe aeration, below minimum fluidization, the impact of riser flow was negligible. At higher standpipe aeration, the solid circulation rate increased with riser flow. However, as the riser flow
increased to higher levels, the influence on circulation rate diminished. As expected, the blocked variable (date of testing) was not significant. Thus a correlation was desired that included each of these independent variables. The influence of riser flow and standpipe aeration may be expected to vary depending upon the relative volumes of the riser and standpipe because these are influenced by the shifting of inventory from the riser to the standpipe as discussed above. The relative effects of the standpipe move aeration, solids mass, and loopseal aeration on the solids flux are presented in Figure 8. As described above, the solids flux increased with increasing CFB mass inventory in a simple linear manner. The rate of this solids flux increase with inventory was found to increase with standpipe move aeration (as represented by the three pairs of gradually increasing sloped lines). The lines in Figure 8 represent the averages of different standpipe aeration rates and thereby reflect all of the three- and four-way interactions involving riser and loopseal gas velocities with Umove and M. As noted above, this two-way interaction between Umove and M was only significant at the 90% confidence limit. The linear effect of total solids inventory in the CFB is consistent with the nondimensional form of solids circulation rate as developed by Kehlenbeck et al.12 for their study on scaleup of circulating fluidized beds. In Figure 8 the solids flux is plotted against the solids inventory while all of the aeration rates were held constant. This influence was sustained regardless of the levels of aeration rates. This could be expected since the increased solids load, M, in the CFB resulted in greater bed height in the standpipe, which in turn generated higher pressure gain at a given distributed, standpipe “move” aeration rate. The scatter in the central mass value is representative of the experimental variability. This represents the replicated center point in the central composite design. Additional pairs of data exhibited parallel dependence on solids inventory but for different levels of aeration. Likewise the higher standpipe and loopseal fluidization aeration rates resulted in proportionally greater solids fluxes. Regardless of inventory, the solids flux had very little dependence on gas velocity at low standpipe aeration rates, while at high standpipe move aeration rates a higher riser gas velocity resulted in higher solids fluxes. The magnitude of the effect of gas velocity on solids flux was comparable to that of increased solids inventory. This effect is likely dependent on the relative
1848
Ind. Eng. Chem. Res., Vol. 46, No. 6, 2007
Figure 9. Comparison of experimental results with predicted msDr/ MUt from eq 3.
volumes in the riser and standpipe and on the relationship between gas velocity and solids fraction in the riser. Under the conditions for this particular system the higher gas velocity reduced the solids holdup in the riser by an amount similar to that which was varied in the test matrix, about 10 kg. At the lower standpipe move aeration rates, the riser was very dilute and the amount of change in the solids inventory in the riser caused by a change in riser flow resulted in a negligible or insignificant change in the standpipe height. The variability or spread in the solids flux for each group of move aeration velocities in Figure 8 is represented by the range (not error) bars on each of the move velocity lines. These ranges were greater at the higher move aeration rates than the lower ones. In addition, the solids flux increased with higher loopseal aeration in each case where everything else was kept constant. The increase was greater at higher levels of M, Ug, and Umove due to the independent main factor effects. None of the interactions with Ulpsl was sufficient to be significant at the 95% confidence limit. Based upon the above analyses, an empirical equation was developed using the Kehlenbeck et al.12 form for nondimensional solids circulation rate which correlated both the CFB and loopseal operating parameters. The linear dependence on solids inventory was incorporated in the dependent variable, the dimensionless solids circulation rate term. This empirical correlation, which captured the interactions among these operating parameters, is presented in the following equation.
( ) ( )( )
m˘ sDr Ug2 ) 0.00192 MUt gDr
0.323
Umove Umf
1.27
Ulpsl Ut
0.18
(3)
In this expression each of the factors determined to be statistically significant was incorporated in dimensionless form. From the values of the exponents the relative contributions of each of these factors can be readily determined. The dominant factor was the aeration at the base of the standpipe, Umove/Umf. The riser superficial velocity was surprisingly important. This can be thought to be mainly a result of the impact that increased riser velocity has on reducing riser pressure drop in this core annular regime. The loopseal aeration exhibited relatively minor impact on solids flux even though the aeration rate was varied from more than twice minimum fluidization to well below minimum fluidization. In Figure 9 the experimental data for cork were plotted against the predicted values, m˘ sDr/MUt, calculated from the above empirical correlation. The mean square and the mean relative errors between the measured solids mass flux and the predicted values calculated from eq 3 are 0.2 kg/(m2 s) and (4.5%. This explains 97% of the variance in the
Figure 10. Pressure drop across the vertical section of the loopseal.
data. As can be seen, the correlation provided excellent agreement with the experimental data for this light material under the ranges of the operating conditions. Pressure Drop across the Loopseal. The pressure drop across the loopseal is important for solids flow control, yet excessive pressure drop can limit the circulation rate and associated mixing desired in many energy-conversion applications. In Figure 10, pressure drop across the lift-leg section of the loopseal is plotted against the solids mass flux (Gs,v). As can be seen in this figure, the pressure drop was essentially constant for the lower circulation rates. Since the principal component in the loopseal is a bubbling fluidized bed, the pressure drop is equivalent to the pressure drop across the bubbling fluidized bed, which remains constant once it reaches the point of minimum fluidization. As the solid mass flux (Gs,v) increased above a threshold solids velocity, the pressure drop across the valve began to increase. This corresponded to an increase in loopseal aeration above minimum fluidization velocity. This may have been due to either an increase in bed density or an increase in bed height in the loopseal lift leg. A similar increase in bed density was observed for incremental pressure drops across a smaller subsection of the lift leg of the loopseal. This supports the concept that the increased circulation rate was primarily due to an increase in the bed density. With this in mind, the relationship between solids flux and pressure drop across the lift leg of the loopseal was quantified in the following correlation:
∆P865 ) 1.17 + L
0.12 Gs,v 1+ 16.3
( )
-7.25
(4)
The intercept represents the mean initial bed density across the loopseal lift leg while the numerator in the first term represents the level of increase in bed density at higher flux levels. The exponential term and denominator under the solids flux reflect the gain factor or rate of change of loopseal pressure drop per unit length. The pressure drop across the entire loopseal (from bottom of the standpipe to bottom of the riser, ∆P801) is illustrated by Figure 11. The analysis of variance for the pressure drop across the loopseal displayed a significant lack of fit. This may be attributed to hysteresis effects due to the potential for gas bypassing in the lift leg of the loopseal under certain operating conditions. In support of this, a block effect was exhibited for loopseal pressure drops indicating that the results on the first day of testing were significantly different from the same matrix
Ind. Eng. Chem. Res., Vol. 46, No. 6, 2007 1849
Figure 11. Pressure drop across the entire loopseal. Figure 13. Effects of CFB inventory, standpipe move aeration, and riser gas velocity on the pressure drop across the riser.
For ∆Pr the significant interactions included Ug × M and Umove × M. This can help us understand the role that solids inventory played in determining the operational performance of the riser in the CFB loop. The effect of the aeration flow, either to the riser or to the standpipe, on the riser pressure drop depended upon the total inventory in the CFB. At low inventories a change in the aeration to the riser or standpipe had the same effect independent of the level of the gas flows. However, at the higher inventory the increase in riser pressure drop was more pronounced at lower riser gas velocities and at higher move aeration rates. For Hsp there were no significant interactions, but there was also a significant nonlinear dependence on the inventory, M. Figure 12. Effects of loopseal and standpipe move aeration on the pressure drop across the loopseal. Data points represent the midlevel values for variables not shown. The symbol dark blue 2 represents Umove/Umf ) 1.2 and light blue 2 is 0.76, while red [ is Ulpsl/Ut ) 0.27 and pink [ is 0.15, with all points being averaged over riser gas velocity and total inventory.
center points tested on the second day of testing. As a result, the significance of each of the test factors was interpreted with skepticism. A measure of the radial flow distribution would be required to investigate this hypothesis as a cause for the lack of fit observed. When using a general linear model including main, quadratic, two-way interactions, and three-way interactions, the only significant factors influencing the loopseal pressure drop were the fluidization aeration rate under the loopseal and the move aeration rate. As anticipated, the higher the loopseal aeration the lower the ∆Plpsl as a result of reducing the bed density in the lift leg, while the higher the Umove the higher the ∆Plpsl as a result of higher solids flow through the loopseal (Figure 12). A closer inspection of the data suggests that the response of ∆Plpsl to loopseal aeration depended upon whether the move aeration was at a high level, where the solids flow dominated the response, or at a low level, where bed density was dominant. Pressure Drop across the Riser and Standpipe Height. The pressure drop across the riser reflects the bed density in the reacting region of a typical circulating fluidized bed reactor in energy-conversion applications. This pressure drop is coupled with the standpipe through the nonmechanical valve. In the cases of pressure drop across the riser, ∆Pr, and the height of bed material in the standpipe, Hsp, both exhibited nonlinear dependence on the riser gas velocity, Ug × Ug (see Figure 13 for example). The inventory in the riser and standpipe were inversely related to each other with the gas velocity in the riser serving as the conduit to move solids from one to the other.
Summary Standpipe and loopseal aeration, solids inventory, and riser superficial gas velocity all significantly influence the solids circulation rate of the loopseal within the National Energy Technology Laboratory’s 0.3 m diameter CFB cold model when operated with a light Geldart type B bed material, cork. Standpipe aeration flow ramp experiments were conducted, demonstrating that the threshold aeration velocity in the standpipe was markedly influenced by the superficial riser velocity. A statistically designed composite test was then conducted in order to map the influence of each of these independent parameters on solids circulation rate and loopseal pressure drop. While the standpipe aeration was dominant factor, the riser gas velocity was also found to be a relatively large factor. Increasing the riser flow resulted in increased solids fluxes; however, the impact of riser flow was greater at lower riser flows and when standpipe aeration rates were relatively high. The solids circulation rate was linearly related to the inventory of solids in the CFB unit. Pressure drop across the fluidized-bed up-flow section of the loopseal was found to be relatively insensitive and increase slightly with the solid flow rates. These observations taken in conjunction with previous researcher findings were used to develop empirical correlations to predict the solids flow rate and pressure drop. Comparison of this correlation with the experimental data produced excellent agreement for this light material. These results confirm that solids inventory and riser gas flow affect the performance of a nonmechanical valve, specifically the loopseal, operating within an integrated circulating fluidized bed loop. The authors caution that the nonmechanical valve cannot be completely separated from the circulating fluidized
1850
Ind. Eng. Chem. Res., Vol. 46, No. 6, 2007
bed loop to quantitatively evaluate its performance. The riser gas velocity influenced each and every dependent parameter studied in a significant manner.
∆Pr ) pressure drop across the entire length of the CFB riser (kPa) Literature Cited
Acknowledgment The authors acknowledge the Department of Energy for funding the research through the Fossil Energy’s advanced research and gasification technology programs. Notation dp ) mean particle diameter Dr ) riser diameter (m) g ) acceleration due to gravity (m2/s) Gs ) solids flux (kg/(m2 s)) Gs,v ) solids flux through loopseal (kg/(m2 s)) m˘ s ) solid mass flow rate (kg/s) M ) inventory (kg) Ug ) superficial gas velocity (m/s) Ulpsl ) superficial velocity of the loopseal aeration over the cross-sectional area of the lift leg in the loopseal (m/s) Umf ) minimum fluidization velocity (m/s) Umove ) superficial velocity of the standpipe aeration over the standpipe cross-sectional area (m/s) Ut ) terminal velocity (m/s) Greek Symbols b ) bed voidage mf ) voidage at minimum fluidization Fb ) bed density (g/cm3) Fs ) solid density (g/cm3) φ ) shape factor ∆P865 ) pressure drop across the lift-leg riser section of the loopseal (kPa) ∆P873 ) pressure drop across bottom section of the riser (kPa) ∆Plpsl or ∆P801 ) pressure drop across the loopseal (kPa)
(1) Geldart, D.; Jones, P. The behavior of L-valves with granular powders. Powder Technol. 1991, 67, 163. (2) Daous, M. A.; Al-Zahrani, A. A. Modeling solids and gas flow through an L-valve. Powder Technol. 1998, 99, 86. (3) Knowlton, T. M.; Hirsan, I. L-valves characterized for solids flow. Hydrocarbon Process. 1978, 57, 149. (4) Knowlton, T. M.; Hirsan, I. The effect of aeration tap location on the performance of a J-valve. Proceedings of Second Engineering Foundation Conference on Fluidization; Cambridge, England, April 1978; p 128. (5) Leung, L. S.; Chong, Y. O.; Lottes, J. Operation of V-valves for gas-solid flow. Powder Technol. 1987, 49, 271. (6) Kim, S. W; Kim, S. D.: Effect of particle properties on solids recycle in loop-seal of a circulating fluidized bed. Powder Technol. 2002, 124, 76. (7) Ludlow, J. C.; Lawson, L.; Shadle, L.; Syamlal, M. Development of a spiral device for measuring the solids flow in a circulating fluidized bed. SeVenth International CFB Conference, Niagara Falls, Ontario, Canada; May 2002; p 513. (8) Ghordzoe, E.; Smith, P.; Vimalchand, P.; Lui, G.; Longanbach, J. Initial Operations of the PSDF Transport Gasifier. Proceedings of the 16th International Conference on Fluidized Bed Combustion; Geiling, D. W., Ed.; ASME: New York, 2001; FBC01-0065. (9) Wen, C. Y.; Yu, Y. H. Mechanics of Fluidization. Chem. Eng. Prog. Symp. Ser. 1966, 62, 100. (10) Monazam, E. R.; Shadle, L. J.; Mei, J. S.; Spenik, J. Identification and Characteristics of Different Flow Regimes in an Industrial Scale Circulating Fluidized Bed. Powder Technol. 2005, 155, 17. (11) Knowlton, T. M. Standpipe and return system. In Circulating Fluidized Beds, 1st ed.; Grace, J. R., Avidan, A. A., Knowlton, T. M., Eds.; Blackie: London, 1977; Chapter 7, p 240. (12) Kehlenbeck, R.;Yates, J.; Di Felice, R.; Hofbauer, H.; Rauch, R. Novel Scaling parameter for circulating fluidized beds. AIChE J. 2001, 47, 582.
ReceiVed for reView May 23, 2006 ReVised manuscript receiVed November 1, 2006 Accepted November 2, 2006 IE0606486
