Design and Fluid Dynamic Analysis of a Bench-Scale Combustion

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Ind. Eng. Chem. Res. 2005, 44, 546-556

PROCESS DESIGN AND CONTROL Design and Fluid Dynamic Analysis of a Bench-Scale Combustion System with CO2 Separation-Chemical-Looping Combustion Bernhard Kronberger,*,† Anders Lyngfelt,‡ Gerhard Lo1 ffler,† and Hermann Hofbauer† Institute of Chemical Engineering, Vienna University of Technology, A-1060 Vienna, Austria, and Department of Energy Conversion, Chalmers University of Technology, S-412 96 Gothenburg, Sweden

Chemical-looping combustion (CLC) is a novel combustion technology with inherent separation of the greenhouse gas CO2. The technique involves the combustion of gaseous fuel using a metal oxide for oxygen transfer. The system consists of two reactors, a fuel reactor and an air reactor, and an oxygen carrier in the form of a metal oxide that transports oxygen from the air to the fuel. Direct contact between fuel and combustion air is avoided, and the products from combustion are kept separated from the rest of the flue gases. A dual-fluidized bed reactor system representing a 10-kW CLC prototype was designed. A scaled flow model was built and investigated. Gas velocities and designs were varied, while solids circulation rate and gas leakage between the reactors as well as static pressure balance and residence time distribution of gas and particles were measured. Results show that the solids circulation rates were sufficient and the gas leakage could be decreased to very low values. Introduction

Me + 1/2O2 f MeO

Besides oxyfuel-technologies, precombustion and postcombustion methods, advanced options such as calcination/carbonization cycles1 and the AZEP process2 for CO2 separation are currently under investigation for CO2 capture. Another novel technology, which can be grouped together with fuel cells into indirect fuel oxidation technologies, is chemical-looping combustion. A chemical-looping combustor oxidizes a fossil fuel gas in two stages and was first mentioned by Knoche and Richter3 as a method to enhance the thermal efficiency of combustion. First, the fuel is used to reduce a metal oxide, thus producing an outlet stream of concentrated CO2 and steam from the fuel reactor, where the CO2 can be readily separated from water and is provided, for example, for storage (Figure 1). Second, the reduced metal oxide is reacted with air, producing an outlet stream from the air reactor consisting of N2 and any nonreacted O2 only. According to the scheme shown in Figure 1, the gaseous fuel introduced to the fuel reactor reacts with the oxygen carrier according to eq 1:

(2n + m)MeO + CnH2m f (2n + m)Me + mH2O + nCO2

∆Hox

(2)

Reaction 1 is, depending on the metal oxide type, often endothermic, and reaction 2 is always exothermic.4 The net chemical reaction over the two reactors, however, is the same as for normal combustion, and the total amount of heat evolved is equal to normal combustion of the same fuel (eq 3).

∆Hc ) ∆Hox + ∆Hred

(3)

The metal oxide is thus an oxygen and heat carrier simultaneously. Basic requirements for the oxygen carrier are high surface area for fast reaction and good physical properties such as crushing strength and attrition resistance. This can be reached by supporting the potential metal oxide types (Fe, Cu, Ni, Mn) by an inert such as alumina and titanium dioxide.5-13 The experimental results of these studies show that the rates

∆Hred (1)

The reaction in the oxidizer, the regeneration, follows eq 2: * To whom correspondence should be addressed. Tel.: +431-58801-15955. Fax: +43-1-58801-15999. E-mail: bkron@ mail.zserv.tuwien.ac.at. † Vienna University of Technology. ‡ Chalmers University of Technology.

Figure 1. Principle of chemical-looping combustion (CLC). Me and MeO denote oxidized and reduced oxygen carriers.

10.1021/ie049670u CCC: $30.25 © 2005 American Chemical Society Published on Web 01/08/2005

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of reaction for both the oxidation and the reduction are fast enough for practical applications. The major advantage of this system is that CO2 and H2O are separated inherently from the rest of the flue gases without the need of extra energy for the separation. This is in contrast to the common techniques for separating carbon dioxide from flue gas, where large amounts of energy and expensive equipment are necessary and reduce the thermal efficiency. Another beneficial peculiarity of CLC is a reduction of the exergy destruction and was reported earlier.14 In conventional combustion processes, the largest irreversibilities occur directly in the combustion process. Options for improvements are to increase the combustion temperature, which is limited by material properties, or to oxidize the fuel in a controlled way as in a fuel cell or a chemical-looping combustor. So far, very limited work has been carried out in the design of a CLC reactor system. Batch experiments and general feasibility analysis of Lyngfelt et al.15 showed that a fluidized bed reactor concept is suitable for chemical-looping combustion. Johansson and co-workers16,17 have presented cold-flow modeling results of different layouts of small laboratory-scale reactor systems for chemical-looping combustion. Adanez et al.18 have presented a kinetic fuel reactor model that allows optimization of the reactor geometry. Ryu and colleagues19 have presented a preliminary design of a 50 kWth unit based on a pressurized fluidized bed technology. However, the development of the technology requires the demonstration of the process, preferably in a benchscale unit. The functionality of such a system and effects of continuous operation, which cannot be evaluated in batch operation, must be studied. The aim of this paper is to present a conceptual design of a 10-kW thermal power CLC prototype working at atmospheric pressure. The design criteria are discussed, and a final design is presented. A cold-flow model of the concept was manufactured, and an extensive test program to evaluate the hydrodynamic behavior was conducted. Results of these cold-flow modeling experiments are presented, along with conclusions on the suitability of the concept. Design Requirements of CLC Prototype Continuous testing in a chemical-looping prototype is not only needed to demonstrate the principle of this new combustion technology. It is also essential to verify the usefulness of the particles developed. Most laboratory tests of particle reactivity cover only a limited number of cycles. The number of hours of operation for the particles in a real system could be on the order of thousand, whereas the number of cycles for particles could be on the order of hundred thousand. Furthermore, the laboratory tests do not show if the particles would be degraded by attrition or fragmentation at the velocities in a full-scale unit. The purpose of the present project is to demonstrate this new combustion technology and to verify that the oxygen-carrier particles developed are able to survive the conditions of a real process. General Issues. An important requirement of the prototype testing unit is to create the possibility to demonstrate the new combustion technology in continuous operation. Further, performance tests of oxygencarrier materials having varying chemical and physical

properties, showing the long-term behavior at various temperature levels and operating conditions, are required. To this end, a very flexible unit is designed to fulfill the requirements but also to create a safety margin for uncertainties in the design. As for the scale of the test rig, it was decided to design a bench-scale unit. On the one hand, that scale will limit costs and ease handling of the reactor and, on the other hand, allow conclusions on and facilitate the scale-up of the process. Due the heterogeneous nature of the reaction and the required solids transport between the fuel reactor and the air reactor, a circulating fluidized bed (CFB) concept is the preferred reactor type. Dual fluidized beds are used in a multitude of processes such as biomass pyrolysis20 and gasification,21 where good contact between solids and gas is required. Additionally, such a system can guarantee the consideration of the following criteria, crucial for reliable operation of a CLC unit: (i) solids material circulation between the two reactors shall fulfill process requirements, (ii) particle inventory in each reactor shall be sufficient for conversion of gas, and (iii) gas leakage shall be minimized. The proposed dual fluidized-bed concept is built up by a transport reactor acting as a riser and a stationary fluidized bed. The riser gives the driving force for the solid material circulation, which has to fulfill two main objectives. First, it shall provide sufficient oxygencarrier capacity for complete conversion of the fuel gas in the stationary bed; second, the mass flow shall implicitly supply the energy transfer needed for balancing the temperatures between the oxidation and the reduction reactor. Design Procedure. The design criteria for the prototype were chosen for a thermal power of 5-10 kW. Although potential for higher cycle efficiency of a CLC power plant is attributed to pressurized CLC systems,22 this prototype unit is to be operated at atmospheric pressure. Clearly, the CFB concept allows pressurization of the process, but it is considered premature at this stage of development. The design procedure of the prototype unit (Figure 2) is based on the type of fuel gas. Natural gas is considered suitable for the process, although any other gaseous fuel like synthesis gas from coal gasification could be used as well.23 In particular, synthesis gas from solid fuels such as coal is gaining more attention, although the coupling of a gasification process with a novel combustor system seems premature. Most crucial design input data are the type of the oxygen carrier, the metal oxide. In the CLC prototype unit, mainly Ni and Fe type carriers are to be tested. The mean diameter of the carrier particles was chosen to be between 100 and 200 µm, having a mean density of 2300-5600 kg m-3. The mechanical properties must allow continuous cycling in the CLC unit at operating temperature without strong attrition or causing sintering of the material over several thousands of cycles. Experimental investigation of oxygen carrier material showed that 950 °C is a possible operating temperature for the fuel reactor. The air reactor temperature, however, is different because of diverse heat of reaction of oxidation and reduction reactions. The reactivity of the carrier material must allow sufficient reduction and oxidation rates. High gas conversion is important as it directly influences CO2

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Figure 3. Operating regime of the prototype unit according to Bi and Grace.34

Figure 2. Design procedure of a chemical-looping combustor.

capture efficiency; otherwise, unreacted fuel must be recirculated to the fuel reactor. Simultaneously, the carrier must have an oxygen transport capacity that leads to full fuel oxidation. The desired power output of the prototype determines the fuel and air mass flows, thus the air-to-fuel ratio. The required solids circulation is determined from heat and mass balances, where for most oxygen carrier types the heat balance is the determining factor. Calculations by Kronberger et al.4 showed that a solids flow of 5 kg m-2 s-1 MW-1 is a feasible value. The choice for the riser acting as air reactor and the fuel reactor as stationary bubbling fluidized bed is based on proven reactivities. Most oxygen carrier types demand a higher particle residence time7-9 for the fuel oxidation. In turn, the bed mass in each reactor in combination with the solids circulation flow results in particle residence times that are dependent on the reactivity of the oxides. The hydrodynamics of the CFB system are predominantly determined by the operation conditions of the riser. Velocities similar to typical CFB risers are desired, and 4-10 times terminal velocity was assumed as superficial gas velocity (Figure 3). This is thought to cause for all particles sufficient solids entrainment for the required solids circulation rates. The major novelty of the transport reactor is that the bottom section has a wider cross section. The extended volume of the widened section, referred to as air reactor, causes a higher mean particle residence time in the oxidation zone. Velocities between 1.2 and 3 times terminal velocity are assumed to be appropriate in this zone. The stationary bed of the fuel reactor is operating in the bubbling fluidized bed regime. Low gas velocities below 0.7ut and above 2umf, depending on the particle properties and thermal power, are design parameters. If methane is used as fuel, the ratio of fuel molecules to

reactant molecules is three, which gives an upper limit of 0.75ut/3 ) 0.25ut based on the methane flow. No particle separator is planned for simplicity reasons for the fuel reactor, although very low solids entrainment is desired. To accommodate for the gas volume expansion in the fuel reactor, the cross section area increases with height by means of two conical sections. Gas leakage is difficult to avoid completely in a CFB system. However, if there is a gas leakage from the fuel reactor into the air reactor, carbon dioxide will be released to the atmosphere, and the CO2 capture efficiency decreases. If there is a leakage in the opposite direction, the carbon dioxide stream will be diluted with nitrogen that increases the cost of compression. Extra expenditures for separation could also be incurred. It is thus important that the system is designed to minimize leakage. The pressure drop of both fluidized beds is governed by the bed height, which itself follows from required particle residence time. The bed height of the fuel reactor was chosen on the basis of reactivities obtained in previous studies with iron-based oxygen carriers. Because reactivity data are different for the various metal oxides and also a safety margin shall be included, the bed height can be controlled by an adjustable overflow height. To do this, however, the outlet pipe has to be removed. For the freeboard, twice the height of the dense bed height (about 0.1 m) is assumed, which results in a total height of about 0.35 m. Based on the estimated reactivity of most metal oxide particles, the needed height of the dense bed in the air reactor is small. However, the bed mass in the air reactor and accordingly the bed height are related to the total solids inventory in the unit. This gives the need for a certain bed mass in the air reactor for stable operation. Also, for cooling purposes, an adequate surface area must be considered. The dimensions of the downcomers and the loop seals are based on considerations of required/expected solids flows and allowable particle velocities. The bed mass in each section determines the pressure situation in the system. The understanding of the pressure loop is important as, for instance, too large of a pressure difference between two reactors can cause malfunction of the loop seals. Design Concept of the CLC Prototype. The design concept of the CLC prototype (Figure 4a and b) shows the transport reactor (riser) and the fuel reactor. As two different particle separator designs were tested, the two graphics show the different concepts. In the case


Figure 4. (a) Configuration A; (b) configuration B. Design layout of a dual fluidized bed system (10-kW CLC prototype) with (A) air reactor, (B) riser, (C) particle separator, (D) downcomer, (E) loop seal, and (G) fuel reactor with standpipe overflow.

of the standard configuration A (Figure 4a), a cyclone is used for solid separation of the air reactor exit flow. Fuel reactor and air reactors are separated by L-type loop seals. The fuel reactor gas exit is designed as having a large cross section as high gas velocities would cause high particle loss due to entrainment. The solids flow leaving the fuel reactor by an overflow standpipe is returned into the air reactor at a height corresponding to an appropriate pressure level in the transport reactor. The advanced design (Figure 4b) is identical to the aforementioned configuration but includes an alternative particle separator design. The “hat” separator is based on the settling chamber principle and is intended to decrease exit effects on the solids flow. The particle back-flow in the riser is lower for such a concept, which increases the solids flow and leads to smaller pressure drops for a given solids flow. The particle/wall friction is minimized because of lower velocities, which compared to the cyclone has a positive effect on undesirable attrition of the oxygen carrier particles. Further, because of a smaller pressure drop of the separator, it leads to reduced energy consumption of the gas fans. A negative effect is expected on the particle separation efficiency. However, the synthetically produced metal oxide particles can be produced in a narrow particle size range and with very low fines content. Therefore, this disadvantage is considered acceptable. In addition to tests with the standard design of the prototype reactor geometry, tests with different riser

heights and the alternative type of particle separator system were made. The riser height variations are aimed at giving more detailed information on the effect of riser height scaling for transferring the results to larger scale CLC units. The design configurations used in these tests are given in Table 1. For larger particle sizes, the thermal power of the unit is raised to 10 kW, and with the increased gas velocities the fluidization regimes of the riser and the stationary bed are still in a range that is considered appropriate for operation of the combustor. The air ratio in that case increases up to about 2.5. Although the heat balance has to be fulfilled for all thermodynamic particle properties, it is not required to keep the air ratio low as the unit is not designed to represent a large scale boiler. A summary of the design values chosen for the CLC laboratory unit is given in Table 2. The entire CLC prototype unit installation scheme is presented in Figure 5. The setup is equipped with measurement instrumentation, monitoring, and data collection with pressure transducers, thermocouples, and mass flow controllers. The following gases are supplied to the unit: inert gases for the loop seals (2 and 4), methane as fuel gas (3), and air (5) for the oxidation reactor where the latter is preheated by an electrical heater (6). The gas exit flows are passing a free convection cooling section (8). Additionally, the air reactor temperature can be automatically controlled by cooling air, which is fed to a jacket surrounding the air reactor. The fuel reactor temperature can be adjusted with the heating coils (7). Two large filters (9) are installed to achieve a low-pressure drop of the air reactor exit train to the chimney (10). Manual switching during operation enables the measurement of elutriation over a chosen time period. A water trap (1) is used to control the pressure balance of the reactors (air reactor and cyclone vs fuel reactor). The water seal has an overflow exit and also acts as a condensate trap for the humid gas from the fuel reactor, as well as a particle trap. Design Aspects of the Cold-Flow Model. The scaling of the hydrodynamics of the prototype dimensions into the scaled cold-flow model was pursued by applying the scaling rules of Glicksman24 (eq 4). Application of the scaling laws shows that for the material properties selected (cf., Table 2) it is difficult to develop reasonable scaling factors using ambient air. Therefore, a gas mixture of, for example, helium/ nitrogen is required, and, as bed material, glass beads are suitable. From this, a scaling factor of 0.55 results, and the relationships between the cold-flow model and the prototype reactor are given in Table 3. In general, the design of the cold-flow models is similar to the hot prototype reactor. The cold-flow model is built from acrylic glass and the gas distributor plates of the two reactors are of perforated plate type, whereas for the particle locks porous glass plates were used. The cyclone is designed according to design formulas by Barth-Muschelknautz25 and on data from Hugi.26 Glass beads were selected for bed material, and small amounts of an antistatic power (Larostat) were added to reduce electrostatic charges. The gas mixture for the cold-flow model is provided by a gas recirculation system where the exit gas flows, after passing a filter, are collected in a large gas bag that ensures atmospheric working pressure in the system. A gas pump leads the gas from the bag and raises the gas to operating

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Table 1. Configurations Used configuration A (standard configuration) configuration B configuration AS, BS

cyclone with standard riser alternative particle separator standard riser (Figure 4a) configuration A or B with short riser (Figure 4b))

Table 2. Design Values of the CLC Laboratory Unit Grace lab-scale prototype

operating parameter and design values thermal power fuel type air-to-fuel ratio operating pressure reactor temperatures particle density mean particle diameter gas fluidization velocity in the riser ) u/ut gas fluidization velocity in the air reactor ) u/ut gas fluidization velocity in the fuel reactor ) u/umf loop-seal gas fluidization velocity ) u/umf riser diameter riser height air reactor diameter air reactor height fuel reactor (maximum) diameter fuel reactor bed height fuel reactor total height total reactor system height

kW Pa °C kg m-3 m

5/10 methane 1.2-2.6 1 × 105 950 2500-5400 (100-200) × 10-6 4-10 1.2-3 5-15 1.2-4

m m m m m m m m

0.072 1.85 0.14 0.53 0.25 0.13 0.34 2.2

pressure of the mass flow controllers. A gas cooler ensures ambient temperature, and safety valves protect the Perspex model from any overpressure. Experiments The cold-flow model was operated with a total bed material inventory between 1.1 and 2.2 kg. Further, gas velocities in the reactors and also in the loop seals were varied. Their effects on the CFB pressure balance and solids circulation rate, mean gas and particle residence time, as well as residence time distribution (RTD) and

Figure 5. Drawing of prototype reactor test setup: (1) water trap, (2) nitrogen, (3) fuel gas, (4) argon, (5) air, (6) preheater, (7) heating coils, (8) finned tubes for cooling of gas streams, (9) filters, and (10) connection to chimney.

Table 3. Relationships between the Laboratory Unit and the Cold-Flow Model Grace prototype

parameter temperature pressure fluidization gas type (riser) solid material type mean particle diameter particle size distribution solid density particle sphericity mass length area velocity volume flow solids circulation rate

°C Pa

950 1 × 105 air

CFM prototype 25 1 × 105 He/N2

m

oxygen carrier glass beads 120 × 10-6 67.5 × 10-6

m

120 × 10-6

kg m-3 2550 ∼1 1 1 1 1 1 1

(40-80) × 10-6 2550 1 0.17 0.55 0.552 0.74 0.22 0.23

gas leakage were tested. An additional pressure relief valve was installed at the fuel reactor exit, and variations were carried out to study the effect of the fuel reactor pressure on the pressure balance of the system, in particular, the gas leakage. Experimental Procedure. All in- and outgoing gas flows of the CFM were measured using mass flow controllers (type MKS flows 5-200lN/min (basis nitrogen) and commercial diaphragm gas meters (type Elster BK). For very low gas flows and for calibration purposes, also a gas bubble meter was utilized. A total number of 20 pressure transducers, type Honeywell, Micro Switch, were used. The solids circulation rate was determined by a short interruption of the fluidization of the lower particle lock. Repeated measurements of the time necessary to fill a dedicated volume in the downcomer were used to determine the solids flux. Total measurement error was found to be below (6% with deviations attributed to variations of the bulk behavior due to electrostatic charges caused by changing air humidity. The test method, instead, proofed a very high accuracy. For the residence time distribution test, a tracer measurement technique of Rhodes et al.27 was adapted. A pulse function of sodium chloride was injected into the solids flow, and bed material samples were taken at the downcomer of the fuel reactor particle overflow at given time intervals. The concentration of the solid sample is determined by a conductivity method, and the RTD distribution function is derived by standard methods. For gas leakage measurements, a tracer gas method was used. Propane was added to the inlet air of all four fluidization gas flows, one at a time, and the concentration of propane in the incoming, as well as in the outgoing, gas flows from the cyclone and the fuel reactor was measured with a flame ionization detector. Solving the mass balances of this overdetermined set of equations leads to the leakage gas flows of each particle lock. For the operation of the prototype, it is important to keep the loss of solids very low. The performance of the particle separators was determined by fractional separation efficiency.


Figure 6. Pressure loop of the cold-flow model (configuration A, total solid inventory (TSI) ) 1.9 kg, uAR ) 0.45 m s-1; letters refer to Figure 4).

Figure 7. Specific solids circulation rate design chart of configuration A. Parameters: total solid inventory (TSI) and bed mass distribution, riser velocity, air/fuel ratio. Gs,min gives the required solid flow from mass and energy balance.

Results and Discussion From start-up of the unit, it could already be seen that the system was working without any difficulties. Nonetheless, a few design variations were indicated and carried out before the main test program was carried out: The fuel reactor geometry was modified as observations suggested that particle mixing was not satisfactory. The cyclone inlet duct was shortened to reduce the particle accumulation. Due to the high operating temperatures, particle softening cannot be excluded totally and that might cause agglomeration and total blocking of the system. The static pressure loops in the unit (Figure 6) for configuration A (Figure 4a) show that the loop seals can balance the pressure differences between the reactors. The absolute pressure in the fuel reactor should be between the air reactor and the cyclone-downcomer pressure. Preferably, the latter is lower than the fuel reactor pressure because then the fuel reactor exit flow, that is, CO2, is not diluted. For similar reasons, the pressure in the fuel reactor should be lower than that of the air reactor to avoid leakage of CO2 into the air reactor. Solids Circulation Rate. In the CLC system, the solids circulation has to fulfill two main tasks: on one hand, providing enough oxygen transport capacity for complete oxidation of the fuel in the fuel reactor and, on the other hand, providing sufficient energy transfer between the two reactors to keep the temperatures at desired values. During the test runs, it could be proved that neither the fluidization velocity in the fuel reactor nor the fluidization velocities of the particle locks are directly influencing the solids circulation rates. This is mainly a consequence of the “overflow type” particle return system in the fuel reactor. All numbers for the solid flow in this section represent averaged values from a minimum of 10 measurements and values and refer to the “hot” prototype unit. Variation of Air Reactor Fluidization Velocity and Total Solids Inventory. As can be seen from Figure 7, the total solids inventory and the velocity in the riser are strongly influencing the specific solids circulation rate. Also, it can be observed that for the

Figure 8. Comparison of specifics solids circulation rate versus riser velocity for riser height variation of configuration A (standard conf.) and AS (A with short riser), TSI/mFR of 2.2 and 3.3.

standard configuration (configuration A) a wide range of values for solids circulation can be achieved in the system. However, it was found that for Gs values higher than about 90 kg m-2 s-1 the system becomes unstable and the fuel reactor starts being filled up higher than the overflow level, which was attributed to the flow limitation of the downcomers. Due to the resulting higher pressure on the particle column in the downcomer, however, it is possible to further increase the solids circulation, but this is considered not to be a desired operation point for the current CFB system. Variation of TSI for Standard Configuration and Reduced Riser Length. Variation of the riser height was carried out throughout the experimental program, and it was found that a riser length reduction of 0.1 m results in an increase of the solids circulation rate up to 20% (Figure 8). The graph shows a comparison for two different solids inventories of the system. The increasing solids flow is a consequence of the fact that the riser height in all design configurations is well below the transport disengagement height (TDH), which is defined according to Geldart28 as the height above which the elutriation rate remains constant. In our case, this was calculated according to Fournol et al.,29 being at least an order of magnitude higher than the actual riser height. Comparison of System Design A and B. Experimental test results obtained with the alternative particle separator, but for equal riser length, are given in Figure 9. It appears that a significant increase of the

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Figure 9. Comparison of specific solids circulation rate versus riser velocity for variation of the particle separator type A (standard conf.) and B (alternative particle separator) and TSI (4.8/ 8.3 kg).

circulation rate is achieved with the “hat” separator (B) for identical mass distribution in the system. The reason for this is the reduced particle reflux at the exit of the riser, the bend into the cyclone inlet duct. For the alternative particle separator, the riser top exit is entirely open and thus presents less resistance to the suspension flow. For all design configurations, the particle separation efficiency of the gas solids separators was determined. In long-term measurements, the particle separation efficiency for both types was above 99.93% and up to 99.97% per cycle. Certainly, this is sufficient for the CLC prototype unit and allows operation with very low loss of precious oxygen carrier in the laboratory. Dimensionless Solids Circulation versus Solids Holdup. The pressures measured along the riser height can be converted into solids hold-up values. At first, this gives the solids distribution along the riser. Although not presented in this paper, the examination confirms the effectiveness of the bottom widening for increasing the mean particle residence time in the riser. Another important result of this analysis is a comparison of the solids holdup detected in the upper portion of the riser with the dimensionless specific circulation rate according to the estimation of eq 4. A successful comparison gives a basis for a mathematical riser model, but, more importantly, the verification allows the determination of the solids circulation rate in the prototype reactor from pressure measurement along the riser. This is of relevance because the test method presented in this study cannot be applied in the hot reactor and other methods are often complicated or somewhat inaccurate.

1-)

Gs (u - ut)Fp

(4)

Although the pressure drop at the top of the riser is very low and thus causes higher measurement inaccuracy and also some uncertainty exists on the actual value of the terminal velocity at the top of the riser, the results are satisfactory. The solids holdup used for this chart is not determined at the very top of the riser but about 0.1 m lower to eliminate the effects of slightly higher solids density due to particle reflux from the riser exit bending. The results of this study are presented in Figure 10. It can be seen that the estimation according to eq 4, which ideally gives the dashed line, agrees well with

Figure 10. Solids holdup versus dimensionless specific solids circulation rate for standard configuration (A). Dashed line gives ideal relationship according to eq 4.

Figure 11. Effect of fluidization velocity (and solids circulation rate) on cumulative RTD function (F(θ) vs dimensionless time θ). CSTR gives comparison with an ideal continuous stirred tank reactor (CSTR).

most measurements data. It is therefore concluded that this analysis can be applied in the laboratory prototype. Particle Residence Time Distribution in the Fuel Reactor. The particle residence time distribution (RTD) can provide vital information for system designers and operators. It is of importance in particular when the gas-solids reaction is the limiting factor and at the low gas velocities that are advantageous for good gassolid contacting and therefore considered for the prototype reactor. The particle age distribution in the fuel reactor influences the kinetics of the fuel oxidation and, thus, the conversion, which is crucial for high thermal efficiency and also for environmental concerns; otherwise unreacted fuel must be recirculated. The residence time distribution in the fuel reactor of the prototype was determined by a solid tracer method as described earlier. The effect of variations of the solids mass flow and the gas fluidization velocity on the cumulative residence time distribution of the reactor versus dimensionless time θ ()t/τ) is shown in Figure 11. The cumulative RTD function F(θ) gives the fraction of molecules exiting the reactor that have spent a time t or θ or less in the reactor and is the integral of the age distribution function E(θ). Curves A and B differ only in the solid mass flow but show an identical behavior for F(θ), which confirms the expectations for dimensionless time scale. Curves B-D


Figure 12. Comparison for RTD between measurement and model data for fuel reactor.

Figure 13. Mass fractions of reactor zones versus reactor fluidization for a solids circulation rate of 15.3 kg m-2 s-1.

are experimental results at different reactor fluidization velocities. All experimental data were taken only to about twice the space time, but it is clear that the curves must reach the value 1 for F(θ) at infinite time. A comparison of the general trend of the RTD curves A-D shows strong deviation from the common assumption of ideal mixing of an ideal continuous stirred tank reactor as proposed, for example, by Geldart.30 Further, a tendency toward a fully mixed bed appears with increasing gas fluidization. A detailed analysis of the RTD function E(θ) results at further parameter variations confirmed the visual observation of nonideal mixing of the bed at low fluidization velocities. It was therefore concluded that in the bed dead or less active regions exist at the low fluidization velocities considered for the prototype. These are most likely located close to the distributor plate between the gas discharge holes but primarily in the annulus of the conical sections. The latter is consistent with visual observations and calculations of the mass in the conical section. As reported by Riviere et al.,31 increasing the fluidization velocity and decreasing the particle size improve particle mixing and reduce the size of the stagnant zones. A mathematical model was developed to represent the phase flows and reactor zones in the bubbling fluidized bed reactor (Figure 12). The model is based on the physical mass transfer mechanisms involved and comprises a spilt of the reactor in a dense zone, consisting of a bubbling and a stagnant zone, and a freeboard section above. Additionally, mass transfer between the different zones is included. The resulting number of parameters allows the determination of meaningful values modeling the behavior of the reactor. Good agreement between model predictions and test results could be found for the RTD function (Figure 12). Thereby, an exponential function was used to add data for higher θ to the experimental values that are only available up to θ of about 2. Mass fractions and solids flux in each reaction zone were determined, and sensitivities on operating conditions were established. The most interesting outcome of this analysis is the significant mass fraction that is apparently not involved in the mixing process of the reactor. Figure 13 shows model results for variation of the reactor fluidization and a clear tendency toward decreasing stagnant zone with increasing gas flow rate. Also, the bed mass in the freeboard zone was identified as significant, and on average only about half of the bed mass is ideally mixed in the bubbling bed for the considered operation range.

Whereas the mass transfer between the freeboard and the bubbling bed was very good in the range of several times the circulation rate, the mass transfer between the bubbling bed and the stagnant zones was found to be small. It was determined as a few percent of the solids circulation mass flow. It should be noted that the results of the RTD analysis must be transferred to the prototype unit with special caution. Although appropriate scaling criteria for fluid dynamic similarity were used for the design of the model, it is clear that the volume increase occurring during methane combustion could not be represented. Also, the gas properties change across height in the prototype fuel reactor, which hinders the direct transfer of the flow model results to the prototype unit. Kinetic modeling results of Adanez et al.18 and Kronberger et al.32 show that about half of the gas conversion in the bubbling bed occurs close above the gas distributor plate because of the highest rates of reaction and high gas-solid mass transfer in this region. Therefore, it is suggested that as an approximation the ratio u/umf is twice for equivalent conditions in the fuel reactor of the hot unit, which gives a stagnant zone fraction of maximum 20% of the total reactor mass. Gas Leakage. The gas leakage rates have been studied because they are important for the CO2 capture performance of the CLC process. Total solids inventory, reactor fluidization velocities, loop seal fluidization, and the pressure balance between fuel and air reactor were varied in these tests. The leakage was determined by propane addition into the inlet gas flows of the different fluidized beds. The exit concentrations were measured by a flame-ionization-detector. In Figure 14, a correlation of the measured gas leakage flow to the gas volume (flow) in the void between the particles of the solids flux is presented. The latter is the net gas volume that is carried in the interparticle void fraction of the moving solids flux. Thereby, the bed voidage in the solids bulk was calculated assuming minimum fluidization regime. The effective gas velocity results in a difference from the particle velocity and the loop seal fluidization gas velocity and was kept constant for the presented data. The results in Figure 14 show gas leakage flows into the air and the fuel reactor, for measurements at different solids circulation rates as well as different pressure differences between the fuel reactor exit and the cyclone downcomer. The basic correlation that can be seen in this figure is the proportionality between the two parameters. As the gas flow in the interparticle void

554


Figure 14. Correlation of gas flow in the void between particles and measured gas leakage for variation of solids circulation rate and pressure difference between FR and downcomer ∆pL. Positive values give the overpressure of the FR in comparison to the pressure in the downcomer of the cyclone.

is also proportional to the solids circulation rate, also this proportionality is noticed. Furthermore, it could be observed that there is no dependency of gas leakage on the pressure drop across the loop seals and on the total solids inventory. As both axes represent measurement data, deviations are attributed to inaccuracies. The understanding of the gas leakage mechanism, that is, the proportionality of the gas leakage and the solids circulation rate, makes possible countermeasure obvious. An increase of the siphon fluidization reduces the gas leakage by a stripping effect of the loop seal fluidization gas. An analysis of this can be seen in Figure 14. In this case, the (specific) gas leakage is defined as follows: absolute gas leakage flow into the reactor divided by the (inlet) fluidization gas flow of the concerned reactor. In the context of CLC, however, a gas stream of special interest is the fuel reactor outlet flow, that is, the CO2/H2O mixture, because a too large dilution of the CO2 produced could give technical problems, and therefore this was used as basis for the representation in this figure. The leakage rates in Figure 14 represent the dilution of the CO2 stream. As can be seen, this would be low, 0.1-0.6%. A leakage into the air reactor, on the other hand, would not give any technical difficulties, but it would reduce the CO2 capture efficiency and hence the “zero-emission” attribute of chemical-looping combustion. It can be seen in Figure 15 that increasing the velocity in the loop seals causes a significant decrease of the gas leakage, which theoretically gives the possibility to totally prevent gas leakage by completely stripping of the solids flow. Another possibility for this is the injection of fluidization gas into the downcomer. From the measurements, also the flow of the particle lock fluidization gas (steam or inert gas in the case of the hot CLC process) could be tracked precisely, and it was found that for siphon velocities up to about 3umf almost the entire gas flow (>97%) is following the solids flow. This result is valuable as it allows also the calculation of the dilution of the gas streams with loop seal fluidization agent. From these observations, it can be concluded that increasing the siphon fluidization velocity is an appropriate measure to decrease the gas leakage into both

Figure 15. Specific gas leakage versus loop seal fluidization velocity for zero pressure difference between air and fuel reactor and standard operating conditions.

reactors. However, it is obvious that this also increases the dilution of both exit flows by the siphon fluidization agent, which, on the other hand, would reduce the system efficiency and thus an optimum for the overall process shall be determined. The transfer of the specific gas leakage results can be pursued directly, but attention must be given to gas mixing between fluidized bed reactors caused by porous particles. Because of the required reactivity, the oxygen carrier particles comprise a porous structure. Franke et al.33 reported an effect of the capture of heavy oils in pores on fluidized bed incineration quality, and, similarly, also in this case an effect is expected in the prototype. An exact analysis can be carried out easily once test data of the prototype and particle porosity and mass flow are available, but additional specific gas leakage up to 1% must be taken into consideration. Conclusions A bench-scale dual fluidized bed combustor for continuous atmospheric chemical-looping combustion was designed and evaluated. A wide range of stable and suitable operating conditions was identified, and the reactor geometry was optimized with respect to solids circulation rate, gas leakage, and bed mass in the reactors. In more detail, the following conclusions can be drawn: The experiments on solids circulation rates showed that high solids flow rates are reached. The circulation rate can be varied in a wide range and depends primarily on the riser fluidization and the total solids inventory of the system. This will allow adjusting of the oxygen transport capacity in the CLC prototype as well as control of the temperature difference between the air and the fuel reactor. Determination of residence time distribution reveals that a considerable fraction of the bed material may not be included in the mixing process of the bubbling bed. The cold-flow model results, however, are conservative, and a significant improvement is expected in the prototype unit because of gas volume increase inside the bed caused by the fuel oxidation reaction. Gas leakage was found to be very low, and also the leakage mechanisms were understood. An effective countermeasure is to increase the flow of inert gas for the particle lock fluidization. Another option is the


injection of inert gas, for example, steam, into the downcomers to create a gas barrier for the leakage gases coming from the reactors. Attention must be paid to possible increase of gas leakage in case of extremely porous bed material particles, and it is necessary to consider this aspect for the operation of the prototype unit. Throughout the work, a number of design alternatives and improvements, such as an advanced option of particle recirculation, were examined, and the findings significantly support the work with the chemical-looping prototype. The results of this study present very useful information on the fluid dynamic behavior of the dual fluidized bed system. Although the reactor concept was specifically designed for chemical-looping combustion, it is clear that such a system may also be used for a number of processes that are based on cyclic reaction and regeneration.35-37 Acknowledgment This work was performed in the framework of the EU founded research project GRangemouth Advanced CO2 CapturE Project (GRACE), ENK5-CT-2001-00571. Nomenclature dp* ) dimensionless particle diameter ≡ Ar1/3 E(θ) ) age distribution function F(θ) ) cumulative RTD function Gs ) specific solids circulation rate (kg m-2 s-1) Gs,min ) minimum required specific solids circulation rate (kg m-2 s-1) ∆Hc ) heat of combustion (J mol-1) ∆Hox ) heat of oxidation (J mol-1) ∆Hred ) heat of reduction (J mol-1) mAR ) bed mass in the air reactor (kg) mFR ) bed mass in the fuel reactor (kg) TSI ) total solid inventory (kg) uc ) superficial gas velocity corresponding to the maximum pressure fluctuation amplitude (m s-1) uFR ) superficial gas velocity in the fuel reactor (m s-1) umf ) minimum fluidization velocity (m s-1) uRIS ) superficial gas velocity in the riser (m s-1) use ) onset velocity of significant solids entrainment (m s-1) ut ) terminal settling velocity (m s-1) U* ) dimensionless velocity ∆pL ) pressure difference between fuel reactor and downcomer (Pa) ) voidage Fp ) density of the bed material (kg m-3) τ ) space time of bed particles (s) θ ) dimensionless residence time of bed particles

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Received for review April 23, 2004 Revised manuscript received October 15, 2004 Accepted November 9, 2004 IE049670U

Design and Fluid Dynamic Analysis of a Bench-Scale Combustion

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