Experimental Investigation on Flow Behaviors in a Novel In Situ

Article pubs.acs.org/IECR

Experimental Investigation on Flow Behaviors in a Novel In Situ Gasification Chemical Looping Combustion Apparatus Xiaojia Wang,†,‡ Baosheng Jin,*,† Xianli Liu,† Yong Zhang,† and Hao Liu*,‡ †

Key Laboratory of Energy Thermal Conversion and Control of Ministry of Education, School of Energy & Environment, Southeast University, Nanjing 210096, People’s Republic of China ‡ Faculty of Engineering, University of Nottingham, Nottingham NG7 2RD, United Kingdom ABSTRACT: A novel cold reactor apparatus for in situ gasification chemical looping combustion (iG-CLC) is proposed and investigated. It is mainly comprised of a circulating fluidized bed (CFB) riser as the fuel reactor and a cross-flow moving bed as the air reactor. The direct hydraulic link between the two reactors brings inherent simplicity and stabilization of the whole system. The CFB fuel reactor provides favorable gas−solids contacts over the whole reactor height. The realization of high solids flux operation conditions greatly enhances the solids holdups, and the gas−solids contacts and reactions in the riser. The moving bed air reactor has advantages in terms of having a low pressure drop, continuous solids flow, and large gas−solids contact area except for the risk of plugging and particle leakage caused by excessive high cross-flow gas velocity. Independent pressure adjustments between the two reactors could control the gas flow direction and restrain the gas bypassing to ensure high CO2 concentration with little nitrogen dilution. The flexible adjustments of flow parameters (e.g., gas−solids residence time, solids holdups, and solids inventory) have been experimentally achieved with the cold model experimental system. Valuable data and operational experience for the further hot experimental system have also been obtained. The design process for the future hot experimental system has also been briefly discussed in this paper.

1. INTRODUCTION It is generally accepted that the increasing emission of carbon dioxide (CO2) may greatly affect the global climate. Nowadays, there is a worldwide interest in capturing and sequestering CO2 generated from large-scale fossil fuel combustion plants. However, considerable extra energy is needed to separate and collect CO2 from conventional combustion systems, as CO2 is greatly diluted by N2 from air. Hence, there is an urgent need to develop new clean combustion methods that can greatly reduce the energy penalty of CO2 capture. Chemical looping combustion (CLC), which provides an inherent feature of isolating CO2, has been revealed as an attractive combustion technology with a low cost of CO2 separation.1,2 As shown in Figure 1, a conventional CLC system is mainly comprised of an air reactor and a fuel reactor. The fuel introduced into the fuel reactor is oxidized to CO2 and H2O by a solid oxygen carrier (OC). The reduced oxygen carrier particles are transferred to the air reactor where they are regenerated by contacting with air. Thus, in the whole combustion process, the fuel is not mixed with air. With a complete conversion of the fuel, the flue gas leaving the fuel reactor would only contain CO2 and H2O. Therefore, after the condensation of H2O, almost pure CO2 can be obtained with a small energy penalty.3−5 Studies on CLC of gaseous fuels have been extensively conducted since the introduction of CLC.4−6 Over the past years, attention has been paid to the coal-fuelled CLC due to the much more abundant storage of coal in comparison with that of gaseous fuels. As shown in Figure 1, for the implementation of coal-fuelled CLC, one of the options named in situ gasification chemical looping combustion (iGCLC) is to introduce coal into the fuel reactor directly. Thus, © 2013 American Chemical Society

Figure 1. Schematic diagram of in situ gasification chemical looping combustion (iG-CLC).

Received: Revised: Accepted: Published: 14208

September 4, 2012 August 19, 2013 September 9, 2013 September 9, 2013 dx.doi.org/10.1021/ie3023884 | Ind. Eng. Chem. Res. 2013, 52, 14208−14218

Industrial & Engineering Chemistry Research

Article

Figure 2. Schematic diagram of the novel iG-CLC experimental system.

the fuel is mixed with the oxygen carrier, and they are fluidized together by a gasification agent (steam or CO2). The reaction steps in the fuel reactor include coal gasification and subsequent reactions of intermediate gasification products with the oxygen carrier. This technical approach has been investigated with a variety of laboratory experiments7−11 that have shown some promising potentials of the iG-CLC technology. The oxygen carrier is a core issue of coal-fuelled CLC systems. As oxygen carriers may suffer from deactivation by sulfur and loss with coal ash due to agglomeration or elutriation, using low-cost and long-lasting material (such as natural minerals) as the oxygen carrier is of paramount importance for an iG-CLC system.12 In addition, a significant amount of oxygen carrier is needed for an industrial scale iGCLC system; hence, the use of cheap oxygen carrier materials would be economically advantageous. Compared with the reactions in the fuel reactor for gaseous fuels, the reactions for coal using inexpensive oxygen carriers are much more difficult because of the more complicated reaction steps and the lower reactivity of the oxygen carriers.

This means the fuel reactor for iG-CLC needs to be designed with extensive gas−solids contacts to enhance the reactions. In this aspect, a bubbling fluidized bed (BFB), which has been widely applied to the gas-fuelled CLC system, may not be a good option for the fuel reactor of an iG-CLC system. This is because the bubbling fluidized bed fuel reactors suffer from the risk of fuel gas bypassing through the bubble phase. Moreover, few gas−solids reactions can be expected to be happening in the freeboard of BFBs due to the insufficient solids holdups in this region.13,14 To some extent, these shortcomings of BFBs can be overcome by increasing the bed height or decreasing the superficial gas velocity. However, these measures will result in a larger sized system and higher solid inventories.15 Compared with BFB fuel reactors, a circulating fluidized bed (CFB) riser could provide sufficient gas−solids contacts over the whole or majority of the reactor height, which ensures that the gas−solids reactions are happening in a more homogeneous and favorable environment. At the same time, the adoption of CFB could reduce the system dimensions and allow the operation to be feasible with lower solid inventories. Hence, on 14209

dx.doi.org/10.1021/ie3023884 | Ind. Eng. Chem. Res. 2013, 52, 14208−14218


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the basis of the previous research,16−22 we selected CFB riser as the fuel reactor for our iG-CLC. Further, in order to ensure high solids holdups, which are good for the gas−solids reactions in the fuel reactor, the high solids flux operation (Gp ≥ 200 kg/ (m2 s)) has been adopted. In order to make the whole system compact and more controllable, we have proposed the use of a cross-flow moving bed as the air reactor, which can be located in the middle of the downcomer. This kind of cross-flow moving bed has advantages in terms of low pressure drop, continuous solids flow, and large gas−solids contact area, except for the risk of plugging and particle leakage caused by excessive high cross-flow gas velocity. They have been widely used in catalyst regeneration, heat recovery, solid drying, and filtering processes.23−25 In this work, we have used a cold model experimental system to investigate the feasibility, controllability, adjustability, and practicality of our novel iG-CLC system, which is mainly comprised of a circulating fluidized bed (CFB) riser as the fuel reactor and a cross-flow moving bed as the air reactor. The feasibility study focuses on the whole system’s normal and steady operation, particularly the circulation of high solids flux particles that is the prerequisite for the realization of the novel iG-CLC. The investigation of controllability mainly concerns the control of the gas flow directions in the two reactors and the gas bypassing between the two reactors, which affects the CO2 concentration at the exit of the fuel reactor. The adjustability study looks at if the adjustments of some important flow parameters (e.g., gas−solids residence time, solids holdups, and solids inventory) are possible and flexible. Feasibility and good controllability and adjustability with the cold model experimental system need to be confirmed before a hot experimental system is set up for real iG-CLC testing. The practicality of the novel iG-CLC system is investigated by conducting a brief design process for a real hot iG-CLC system using the data collected with the cold model system where possible. In order to perform all of the above tasks effectively, we have selected wide ranges of operating conditions (e.g., solids mass flux, fluidizing numbers, and pressure ratio) for the cold model experiments.

Figure 3. Schematic drawing of the cross-flow moving bed air reactor.

cuboid-shaped downcomer (16). Finally, the solids returned into the riser via the J-valve (17). The air stream leaving the inertial separator passed through the bag filter (20) before being discharged to the atmosphere. Another air stream from the 18 kW air compressor (26) was fed into the cross-flow moving bed (12). The gas (air) passed horizontally through the solids layer while solid particles moved downward as shown in Figure 3. Gas flow rates in the riser, J-valve, and cross-flow moving bed were controlled by control valves and measured by flow meters. All gas flow rates with the subscript “sta” in this paper are normalized to the standard state (101.325 kPa for pressure and 20 °C for temperature). The pressures were measured by pressure manometers and controlled by the back pressure regulators. Pressure drops were measured by multichannel differential pressure transducers and continuously data logged by a computer. A digital camera was employed to photograph the flow regimes through the transparent walls during the experiments. The solids circulation fluxes were measured by the butterfly valve (11) located in the upper part of the downcomer. The rotary blade in the butterfly valve was drilled and covered with fine wire mesh. When the butterfly valve was closed, the amount of reduced solids in the lower sections of the downcomer was recorded for a given time period; thus, the corresponding solids circulation flux could be calculated. High purity CO (99.99%) was selected as the tracer gas (27). When the gas−solids flow became stable, the tracer gas I was added to the riser. One minute later, the gas sampling began at the outlets of the inertial separator and cross-flow moving bed. After the gas sampling, the tracer gas I was closed, and the

2. EXPERIMENTAL SECTION 2.1. Cold Model Experimental System Setup. Figure 2 schematically shows the cold model experimental system of the proposed iG-CLC system. It is mainly comprised of a vertical riser (0.06 m I.D. × 5.8 m height) (6), an inertial separator (9), a cylinder-shaped downcomer (0.1 m I.D. × 1.7 m height) (10), a cross-flow moving bed air reactor (12), a cuboid-shaped downcomer (0.1 m × 0.1 m × 2.0 m height) (16), a J-valve (17) and a bag filter (20). For the purpose of visualization, some sections of the riser and the cylinder-shaped downcomer are made of plexiglas. The cross-flow moving bed air reactor (12) is further illustrated in Figure 3.The channel of the moving bed is 700 mm in height, 100 mm in width, and 100 mm in depth. The angle, length, and spacing of louver are 75°, 130 mm, and 60 mm, respectively. The solids inside the moving bed channel are sandwiched between two transparent panels made of plexiglas. Air from the 90 kW air compressor (25) was fed into the riser through a gas distributor. From the riser (6), the suspension of solids was directed into the inertial separator (9). The solids were separated from the air stream and fell into the cylinder-shaped downcomer (10). Then they moved downward through the cross-flow moving bed (12) and 14210



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smaller the value of f I, the higher CO2 capture efficiency will be achieved. 2.2.4. Fluidizing Numbers in the Two Reactors. The fluidizing number is usually used to define how a bed of solids is fluidized.29,30 In this study, we used two fluidizing numbers to characterize the fluidizing gas flows in the two reactors. Nf represents the fluidizing number in the fuel reactor, which is defined as the ratio between the superficial gas velocity of the fuel reactor and the minimum fluidizing velocity under the operating conditions of the fuel reactor

tracer gas II was added to the inlet of moving bed. Then similar gas sampling was carried out 1 min later. 2.2. Data Evaluation. 2.2.1. Apparent Cross-Sectional Average Solids Holdup εp. The cross-sectional average solids holdup (εp) in the riser was estimated from the local pressure drop. According to the previous research,26−28 when the wall friction or solids acceleration is negligible compared with the static head of the solids, the apparent cross-sectional average solids holdup (εp) can be estimated as ΔPZ/ΔZ ≈ [ρp εp + ρg (1 − εp)]g

(1)

Nf =

where ΔPZ is the local pressure drop at two adjacent riser elevations. 2.2.2. Gas Flow Rate. The flow rates of the system were measured by flow meters and were normalized to the standard state. The total inlet flow rate of the system could be calculated as

(2)

Na =

where Q1,sta, Q2,sta, Q3,sta, Q4,sta, and Q5,sta represent the inlet air flow of the riser, fluidizing air flow of the J-valve, aeration air flow of the J-valve, aeration air flow of the makeup material tank, and inlet air flow of the moving bed, respectively. Qg,sta is used to denote the sum of Q1,sta, Q2,sta, Q3,sta, and Q4,sta. According to the mass balance on air, the total outlet flow rate of the system could be calculated as Q out,sta = Q in,sta

Thus, the outlet flow of the inertial separator Qa,sta could be calculated as (4)

where Qb,sta is the outlet air flow rate of the moving bed. 2.2.3. Tracer Gas. Through the use of tracer gas I, we investigated the distribution of the exhaust gas from the riser (fuel reactor) outlet. f I represents the fraction of the riser exit gas bypassing into the air reactor fI =

Q b,staxb,CO Q a,staxa,CO + Q b,staxb,CO

′ Q a,staxa,CO ′ + Q b,staxb,CO ′ Q a,staxa,CO

(8)

3. RESULTS AND DISCUSSION The bed materials were low-grade iron ore particles with a mean diameter (dp) of 0.71 mm, real density (ρp) of 3505 kg/ m3, and bulk density of 1528 kg/m3. The minimum fluidization gas velocity under the cold condition (Umin) is 0.45 m/s. Experiments were carried out with the fluidizing number in the fuel reactor (Nf) ranging from 6.0 to 32.0, and the fluidizing number in the air reactor (Na) ranging from 0.015 to 0.140. The solids mass flux (Gp) was in the range of 120−360 kg/(m2 s). In order to search for the best pressure match relation between the two reactors, we adjusted the pressure ratio Pb/Pa between 1.0 and 2.0 using the back pressure regulators at the outlets of the inertial separator and the cross-flow moving bed (air reactor). Pa and Pb represent the pressures at the outlets of the inertial separator and the cross-flow moving bed, respectively. Table 1 summarizes the adopted experimental conditions. 3.1. System Pressure Profiles. Figure 4 shows the pressure profiles of the whole iG-CLC cold system under the operating condition of Gp = 220 kg/(m2 s), Pb/Pa = 1.3, Nf = 22.9, and Na = 0.042, which is defined as the reference condition for analysis in this paper. It can be seen that there was an obvious decay of pressure along the riser (fuel reactor)

× 100% (5)

where xa,CO, xb,CO are the concentrations of tracer gas I measured at the outlets of the inertial separator and the moving bed (air reactor), respectively. Similarly, we also investigated the distribution of the gas from the moving bed (air reactor) inlet using tracer gas II. f II represents the fraction of the gas from the moving bed inlet bypassing into the inertial separator outlet fII =

Ua Umin

where Ua is the superficial gas velocity of the air reactor, that is, the ratio between the gas flow from the air reactor inlet and the cross-sectional area. In order to make sure the gas flows in the riser can drive the particles for circulation, there is a minimum requirement for the fluidizing number in the fuel reactor, which is denoted as the minimum fluidizing number in the fuel reactor Nf,min. Similarly, in order to protect the steady vertical flow of particles in the moving bed from the problems of plugging and particle leakage, there is a upper limit of the fluidizing number for the air reactor, which is denoted as the maximum fluidizing number in the air reactor Na,max. These two values, measured from the cold experimental system, are two of the crucial parameters that need to be taken into consideration when designing the hot system.

(3)

Q a,sta = Q out,sta − Q b,sta = Q g,sta + Q 5,sta − Q b,sta

(7)

where Uf and Umin are the superficial gas velocity of the fuel reactor and the minimum fluidizing velocity of the bed materials, respectively. Similarly, Na represents the fluidizing number in the air reactor, which is defined as the ratio between the superficial gas velocity of the air reactor and the minimum fluidizing velocity of the bed materials

Q in,sta = Q 1,sta + Q 2,sta + Q 3,sta + Q 4,sta + Q 5,sta = Q g ,sta + Q 5,sta

Uf Umin

× 100% (6)

where x′a,CO, x′b,CO are the concentrations of tracer gas II measured at the outlets of the inertial separator and moving bed (air reactor), respectively. In order to achieve high CO2 concentration at the exit of the fuel reactor, the contact of N2 from the air reactor with the exhaust gas from the fuel reactor should be minimized, which means fΠ should be or approach zero. Although f I does not affect the CO2 concentration at the exit of the fuel reactor, the 14211



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of the pressure drop in the cross-flow moving bed are negligible compared with those in the riser because of the steady flow characteristic. 3.2. Flow Patterns. Figure 6a shows the axial distributions of the apparent solids holdup along the riser (fuel reactor). The solids holdup generally decreases approximately exponentially along with the bed height, except there is a slight increase near the top of riser due to the nearly L-shaped exit configuration.31 Meanwhile, the change of solids holdup is much more intense at the bottom of the riser than that in the upper region, indicating a more uniform distribution of particles in the upper part. The riser can be divided into three sections along the height: the bottom high-density section (εp ≥ 0.1), the middle transition section, and the upper dilute section. Typical near-wall flow patterns of the three visible sections of the riser are presented in Figure 6b−d. Generally speaking, efficient and wild mixing between the gas phase and solid phase could be seen within all of the three visible sections. From Figure 6b, we could observe the dense suspension upflow (DSU) structure (defined as solids mass flux Gp > 200 kg/(m2 s), solids holdup εp ≥ 0.1, and almost no net solids downward motion at the wall)32 at the bottom of the riser. However, as shown in Figure 6c,d, the two higher sections of the riser were still in the dilute/transition phase, which is consistent with Figure 6a. Moreover, some small clusters could also be observed in the upper regions. Admittedly, due to the wall effect, the solids holdups within each section will be somewhat higher near the wall than those in the central region. Figure 7 displays the solids flow patterns at the visible sections of the cylinder-shaped downcomer and the cross-flow moving bed. We observed a steady interface in the downcomer, indicating a favorable solids circulation around the whole system. Meanwhile, a smooth solids layer was observed in the cross-flow moving bed while particles moved downward. The phenomenon of plugging did not appear, indicating favorable gas−solids flows and reactions in the cross-flow moving bed could be achieved. 3.3. Effect of Pressure Ratio on Gas Distribution. In order to achieve the pressure independence of the two reactors, we adopted two air compressors to supply gases to the riser (fuel reactor) and cross-flow moving bed (air reactor), respectively. Thus, the pressures of the two reactors could be controlled by adjusting the back pressure regulators. The pressure ratio between the two reactors (Pb/Pa) can have an impact on the gas distribution and system balance, and this needs to be investigated. Table 2 lists the main parameters measured in the experiment for the analysis of gas flow directions. Figure 8a shows the two possible exhaust paths for the flue gas of riser (fuel reactor): the outlet of the inertial separator and the vent of the cross-flow moving bed (air reactor). As we all know, the more flue gas of the fuel reactor gets into the air reactor, the lower the efficiency of CO2 capture even though the CO2 concentration at the outlet of the fuel reactor (i.e., the outlet of the inertial separator) will not be affected. Hence, the part of the flue gas passing into the air reactor should be limited to an acceptable low level. Here we used a parameter f I to represent the fraction of the fuel reactor flue gas passing into the air reactor. As shown in the eq 5, f I could be expressed in more detail as

Table 1. Summary of the Experimental Conditions descriptions solids mass flux Gp (kg/(m2 s)) pressure ratio Pb/ Pa (unitless) fluidizing number in the fuel reactor Nf (unitless) fluidizing number in the air reactor Na (unitless)

range of values 120−360 1.0−2.0 6.0−32.0

0.015−0.140

main research aims testing the effects on the solids holdups and residence time testing the pressure match relation of the two serial reactors testing the minimum fluidizing number in the fuel reactor Nf,min and the effects on the solids holdup and inventory testing the maximum fluidizing number in the air reactor Na,max

Figure 4. Whole-system pressure profiles at the reference condition.

height. Meanwhile, as expected, a low pressure drop was seen within the cross-flow moving bed (air reactor). The total pressure drop in the riser was as high as 8.4 kPa, but it was only 0.778 kPa within the cross-flow moving bed. The pressure difference (3.05 kPa) in the J-valve represents the pressure loss due to the flow of solids. Figure 5 presents the pressure drop fluctuations with time at different sections of the iG-CLC system. ΔP1 and ΔPb

Figure 5. Differential pressure fluctuations with time at different sections: (a) ΔP1 = P0 − P1; (b) ΔP2 = P0 − P2; (c) ΔPb = Pc − Pb (reference condition).

represent the pressure drops of the riser distributor and the cross-flow moving bed, respectively; ΔP2 refers to the differential pressure between the dense-phase fluidized region and the gas plenum. It can be seen that the amplitudes of the pressure fluctuations in dense-phase fluidized region are much higher than those at the distributor due to the intense irregular gas−solids interactions in the bed. As expected, the fluctuations 14212



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Figure 6. Axial profiles of the apparent solids holdup along the riser (a) and typical near-wall flow patterns observed with the three visible sections: (b) 0.7−0.9, (c) 2−2.2, and (d) 4.2−4.4 m (reference condition).

Table 2. Main Parameters Measured in the Experiment for the Analysis of Gas Flow Directions

= =

n f,2 nf

× 100%

n f,2 n f,1 + n f,2

× 100%

Q b,staxb,CO Q a,staxa,CO + Q b,staxb,CO

Qg,sta (m3/h)

Q5,sta (m3/h)

Qb,sta (m3/h)

xa,CO (ppm)

xb,CO (ppm)

x′a,CO (ppm)

x′b,CO (ppm)

1.0 1.3 1.6 2.0

125.0 125.0 125.0 125.0

5.0 5.0 5.0 5.0

12.5 11.5 10 8

1687 1450 1532 1548

405 330 322 304

0 0 28 53

1687 2149 2259 2563

where nf, nf,1, and nf,2 represent the molar flow rates of the total fuel reactor flue gas, the part of the total fuel reactor flue gas exiting from the separator outlet, and the part of the total fuel reactor flue gas passing into the air reactor, respectively. Figure 8b displays the influence of the pressure ratio Pb/Pa on f I. It can be seen that f I maintained low values within the range of the pressure ratios tested and decreased with an increase in the pressure ratio. When Pb/Pa was increased from 1.0 to 2.0, f I decreased from 2.49% to 1.27%, indicating a positive effect of higher Pb/Pa on restraining the fuel reactor flue gas passing into the air reactor. Figure 9a presents the two possible paths for the inlet gas of the cross-flow moving bed (air reactor): the vent of the crossflow moving bed (air reactor) and the outlet of the inertial separator. If the gas from the air reactor inlet (containing N2) gets into the separator and mixes with the fuel reactor flue gas, the dry basis concentration of CO2 will be reduced. Therefore, the fraction of the gas from the air reactor inlet passing into the separator should be minimized. Here, we used a parameter fΠ to represent the fraction of gas from the air reactor inlet passing into the separator. As shown in eq 6, fΠ could be expressed in more detail as na,2 fII = × 100% na na,2 = × 100% na,1 + na,2

Figure 7. Solids flow patterns at the visible sections of the cylindershaped downcomer and the cross-flow moving bed at the reference condition.

fI =

Pb/ Pa

=

′ Q a,staxa,CO ′ + Q b,staxb,CO ′ Q a,staxa,CO

× 100% (10)

where na, na,1, and na,2 represent the molar flow rates of the total gas flow at the air reactor inlet, the part of the total gas flow from the air reactor inlet exiting from the air reactor outlet, and

× 100% (9)

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Figure 8. Effect of Pb/Pa on the fraction of the riser exit gas passing into the moving bed ( f I).

Figure 9. Effect of Pb/Pa on the fraction of the moving bed inlet gas passing into the inertial separator outlet (fΠ).

Figure 10. Axial profiles of the apparent solids holdups along the height of the riser at various solids mass fluxes Gp (Pb/Pa = 1.3, Nf = 19.1 and Na = 0.120).

the part of the total gas flow from the air reactor inlet passing into the separator, respectively. Figure 9b displays the influence of the pressure ratio Pb/Pa on fΠ. It can be seen that fΠ stayed at zero when Pb/Pa was in the range of 1.0 and 1.3, and it increased to 12.9% and 24.0% when Pb/Pa was increased to 1.6 and 2.0, respectively. This indicates that having a higher Pb/Pa has a negative effect on limiting the gas from the air reactor inlet passing into the separator. By considering the results shown in both Figures 8 and 9, it can be concluded that Pb/Pa = 1.3 was the optimal value under the conditions of this study, as fΠ was successfully controlled to zero and f I was limited to a low value. A similar optimal value of Pb/Pa could be used with a future hot iG-CLC experimental system to ensure that little N2 from the air reactor is mixed with the CO2 formed in the fuel reactor and a high efficiency on CO2 capture. 3.4. Flexible Adjustments of the Flow Characteristics. Figure 10 shows the axial profiles of the apparent solids holdups along the height of the riser at various values of solids mass flux Gp. It can be seen that the solids holdup at a given height obviously increased with the increasing solids mass flux, implying that a higher solids flux has a positive effect on the gas−solids contact in the fuel reactor. It can also be seen that the height of the high-density flow region (εp ≥ 0.1) increases with the solids mass flux. This indicates that further increases in

the solids mass flux could lead to the realization of a fully highdensity CFB. Figure 11 shows the effect of the solids mass flux Gp on the solids residence time in the cross-flow moving bed. When the solids mass fluxes increased from 120 to 360 kg/(m2 s), the solids residence time decreased from about 31.5 to 12.3 s. This indicates the cross-flow moving bed air reactor designed could provide a relatively long solids residence time even under the conditions of high solids mass fluxes. When the original solids

Figure 11. Effect of solids mass flux Gp on the solids residence time in the cross-flow moving bed (Pb/Pa = 1.3, Nf = 19.1 and Na = 0.120). 14214



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Table 3. Main Properties of the Used Coala

residence time in the air reactor becomes insufficient, for example, in the case of a higher coal feeding rate, the solids residence time in the air reactor can be easily increased by decreasing the solids mass flux accordingly. Figure 12 exhibits the axial distributions of the apparent solids holdups along the riser height with two different values of

proximate analysis (ar,a wt %) moisture

a

fixed carbon

volatile

10

28.3 48.2 ultimate analysis (ar,a wt %)

ash 13.5

carbon

hydrogen

nitrogen

oxygen

sulfur

63.0

4.1

1.5

6.7

1.2

ar: as received basis.

Table 4. Main Operating Parameters Designed in the Hot Experiment description

value

temperature (K) pressure (bar) coal feed rate (kg/h) thermal power (kW) theoretical air requirement (kg/kg coal)

1223 4 6 ∼50 8.4

Table 5. Calculation Process of the Gas Flows in the Two Reactors Figure 12. Axial profiles of apparent solids holdups along the height of the riser at different values of fluidizing number Nf (Gp = 220 kg/(m2 s), Pb/Pa = 1.3, and Na = 0.120).

reactor type fuel reactor

the fluidizing number in the fuel reactor Nf. It can be seen that when the solids mass flux was fixed, the solids holdup in the riser increased with a decrease in the fluidizing number Nf. This means a lower fluidizing number in the fuel reactor would be helpful in achieving enhanced gas−solids contacts in the fuel reactor without affecting the solids residence time in the air reactor. In addition, decreasing the fluidizing number Nf could decrease the ratio of char/oxygen carrier in the freeboard, which would lead to an increase of the combustion efficiency in the fuel reactor.21 However, an increase in the fluidizing number Nf could promote the solids circulating rate and lower the solids inventory in the fuel reactor. Therefore, for the operation of a hot iG-CLC system, both the potential positive and negative effects of the solids mass flux and superficial gas velocity on the system performance should be considered simultaneously. In this case, achieving a high level of fuel conversion and a good degree of oxygen carrier regeneration would be of paramount importance and hence the operation conditions of the system should be optimized regarding of the solids mass flux and fluidizing number in the fuel reactor. 3.5. Dimensional Design of the Hot Experimental System. 3.5.1. Gas Volume Flows under Hot Conditions. One of the fuels to be tested with the hot experimental system is a bituminous coal from China. Table 3 lists the main properties of the coal. According to the proximate analysis and ultimate analysis, we could calculate that the theoretical air requirement is 8.4 kg/kg coal. Table 4 lists the main operating parameters to be used with the design of the hot experimental system. The temperature, pressure, and coal feeding rate are selected as 1223 K, 4 bar, and 6 kg/h, respectively. From these operating parameters we could further calculate the gas flows of the two reactors, which are shown in Table 5. 3.5.2. Key Data Obtained from the Cold Experimental System Useful for the Dimensional Design of the Hot Experimental System. Because the geometrical magnitude of the hot system, especially the most important CFB riser part,

air reactor

description

value

theoretical steam mass flow needed (kg/h)

4.3

excess steam coefficient (unitless) actual steam mass flow (kg/h) volatile (assumed as CH4) mass flow (kg/h) CO2 produced from char gasification (kg/h) estimated total gas volume flow in the fuel reactor Qf,hot (m3/h) theoretical air mass flow needed (kg/h) excess air coefficient (unitless) actual air mass flow (kg/h) total air volume flow in the air reactor Qa,hot (m3/h)

0.2 5.2 1.7 15.3 24.0 50.5 0.05 53.0 45.8

was projected to be close to that of the cold system, it is adequate to derive the necessary data for the dimensional design of the hot system by means of keeping the fluidization numbers under the hot condition the same as those acquired under the cold condition. As shown in Table 6, the required Table 6. Key Data Obtained from the Cold Experiments Useful for the Design of the Hot System description

value

minimum fluidizing number in the fuel reactor Nf,min maximum fluidizing number in the air reactor Na,max

10.9 0.132

parameters determined by the cold model experimental system are the minimum fluidizing number in the fuel reactor (Nf,min) and the maximum fluidizing number in the air reactor (Na,max). The detailed definitions of the fluidizing numbers Nf,min and Na,max could be seen in Section 2.2.4. By combining the calculated minimum fluidization gas velocity under hot conditions, we can estimate the minimum value of the hot superficial gas velocity in the fuel reactor and the maximum value of the hot gas velocity in the air reactor. Thus, by calculating the ratios between the gas volume flows and gas velocities, we can finally get the upper limit of the fuel reactor cross-sectional area and the lower limit of the air reactor cross-sectional area. Table 7 summarizes the dimensional 14215



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parameters of the hot experimental system that is to be tested in the near future.

bypassing, thereby ensuring high CO2 concentration and capture efficiency. (5) It is convenient to adjust the important flow parameters (e.g., gas−solids residence time, solids holdup, and solids inventory), which can help a real iG-CLC system to achieve high CO2 concentration and high conversions of fuel and oxygen carrier with low solids inventory.

Table 7. Dimensional Parameters of the Hot Experimental System reactor type fuel reactor

air reactor

description minimum fluidization gas velocity under hot conditions Umin,hot (m/s) minimum superficial gas velocity for particle circulating Uf,hot,min (m/s): Uf,hot,min = Nf,minUmin,hot maximum cross-sectional area of the fuel reactor Af,max (m2): Af,max = Qf,hot/Uf,hot,min selected inner diameter of fuel reactor (mm) maximum hot superficial gas velocity Ua,hot,max (m/s): Ua,hot, max = Na,maxUmin,hot minimum cross-sectional area of the air reactor Aa,min (m2): Aa,min = Qa,hot/Ua,hot,max selected height and depth of the air reactor (m)

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value 0.20 2.18

AUTHOR INFORMATION

Corresponding Authors

*B. Jin. Phone: +86-25-83794744. Fax: +86-25-83795508. Email: [email protected]. *H. Liu. Phone: +44-115-8467674. Fax: +44-115-9513159. Email: [email protected].

0.0030 60 0.027

Notes 0.47

The authors declare no competing financial interest.

1.2 and 0.5

ACKNOWLEDGMENTS The financial supports to the reported research by the National Natural Science Foundation of China (51076029), the Ministry of Science and Technology of China (China−EU International Collaboration Project 2010DFA61960), the UK Engineering and Physical Sciences Research Council (EPSRC/China Project EP/G063176/1), the Scientific Research Foundation of Graduate School of Southeast University (YBJJ1119), and China Academic Award for Doctoral Candidates are gratefully acknowledged. The authors would also like to acknowledge the provision of a scholarship to X.W. by the China Scholarship Council (CSC), which enables him to be able to carry out part of the reported work at the University of Nottingham.

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It can be noted, however, the cross-flow moving bed adopted in this study has a relative low value for the maximum fluidizing number Na,max of the air reactor, which indicates the requirement of a large cross-sectional area. Thus, this kind of novel iG-CLC system may be more suitable for applications in small boilers (e.g., for a 50 MWthermal utility boiler, it requires the air reactor has a height of about 30 m and depth of about 15 m). Further research to lower the requirement for the crosssectional area of the air reactor will be carried out in the future.

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4. CONCLUSIONS A novel reactor system for in situ gasification chemical looping combustion (iG-CLC) has been proposed, and the experimental results of the cold experimental system have been presented in this paper. The reactor system is mainly comprised of a high-flux circulating fluidized bed riser as the fuel reactor and a cross-flow moving bed as the air reactor. The results of the present study show that the reactor system has the following main advantages compared to the previously proposed concepts: (1) The direct hydraulic link between the two reactors brings inherent simplicity and stabilization of the whole system. (2) The CFB riser as the fuel reactor promises favorable gas−solids contacts over the whole reactor height. The realization of the high solids flux condition increases the solids holdups in the riser, and further enhances the gas− solids contacts and reactions. (3) The cross-flow moving bed as the air reactor has the advantages of low pressure drop, continuous solids flow, and large gas−solids contact area. Especially, the addition of the cross-flow moving bed on the basis of a CFB system hardly impacts the original solids circulating structure. However, using the cross-flow moving bed as the air reactor has a disadvantage in terms of its relative large cross-sectional area requirement, and this may cause the proposed novel iG-CLC system to be more suitable for applications in small boilers (e.g., 50 MWthermal). Further research to lower the requirement for the cross-sectional area of the air reactor will be conducted in the future. (4) Pressure adjustment between the two reactors could control the gas flow directions and restrain the gas 14216

NOMENCLATURE Aa,min = minimum cross-sectional area of the air reactor of the hot iG-CLC system, m2 Af,max = maximum cross-sectional area of the fuel reactor of the hot iG-CLC system, m2 dp = particle diameter, mm f I = fraction of the fuel reactor flue gas passing into the air reactor fΠ = fraction of gas from the air reactor inlet passing into the separator Gp = solids mass flux, kg m−2 s−1 na = molar flow rates of the total gas flow at the air reactor inlet, mol s−1 na,1 = part of the total gas flow from the air reactor inlet exiting from the air reactor outlet, mol s−1 na,2 = part of the total gas flow from the air reactor inlet passing into the separator, mol s−1 nf = molar flow rates of the total fuel reactor flue gas, mol s−1 nf,1 = part of the fuel reactor flue gas exiting from the separator outlet, mol s−1 nf,2 = part of the fuel reactor flue gas passing into the air reactor, mol s−1 Na = fluidizing number in the air reactor Na,max = maximum fluidizing number in the air reactor Nf = fluidizing number in the fuel reactor Nf,min = minimum fluidizing number in the fuel reactor P = pressure, kPa ΔP1 = pressure drop of riser distributor, kPa ΔP2 = differential pressure between the dense-phase fluidized region and gas chamber, kPa ΔPb = pressure drop of cross-flow moving bed, kPa dx.doi.org/10.1021/ie3023884 | Ind. Eng. Chem. Res. 2013, 52, 14208−14218


Article

ΔPZ = local pressure drop at two adjacent riser elevations, kPa Q1,sta = inlet air flow of the riser, m3 h−1 Q2,sta = fluidizing air flow of the J-valve, m3 h−1 Q3,sta = aeration air flow of the J-valve, m3 h−1 Q4,sta = aeration air flow of the makeup material tank, m3 h−1 Q5,sta = inlet air flow of the moving bed, m3 h−1 Qa,sta = outlet flow of the inertial separator, m3 h−1 Qb,sta = outlet air flow of the moving bed, m3 h−1 Qg,sta = sum of Q1,sta, Q2,sta, Q3,sta ,and Q4,sta, m3 h−1 Qin,sta = total inlet flow rate of the system, m3 h−1 Qout,sta = total outlet flow rate of the system, m3 h−1 Qa,hot = total air volume flow in the air reactor under the hot conditions, m3 h−1 Qf,hot = total gas volume flow in the fuel reactor under the hot conditions, m3 h−1 Ua = superficial gas velocity of the air reactor under the operating conditions of the air reactor, m3 s−1 Uf = the superficial gas velocity of the fuel reactor under the operating conditions of the fuel reactor, m s−1 Umin = the minimum fluidization gas velocity under the operating conditions of the fuel reactor, m s−1 Ua,hot,max = maximum superficial gas velocity of the air reactor under the hot conditions, m s−1 Uf,hot,min = minimum superficial gas velocity of the fuel reactor under the hot conditions, m s−1 Umin,hot = minimum fluidization gas velocity under the hot conditions, m s−1 xa,CO = concentration of tracer gas I at the outlet of inertial separator, ppm xb,CO = concentration of tracer gas I at the outlet of moving bed, ppm x′a,CO = concentration of tracer gas II at the outlet of inertial separator, ppm xb,CO ′ = concentration of tracer gas II at the outlet of moving bed, ppm Z = riser elevation, m ΔZ = height difference in riser, m

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Greek Letters

εp = apparent cross-sectional average solids holdup ρg = air density, kg m−3 ρp = density of iron ore particles, kg m−3

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Experimental Investigation on Flow Behaviors in a Novel In Situ

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