Energy Fuels 2010, 24, 190–198 Published on Web 11/10/2009
: DOI:10.1021/ef900553j
Application of Chemical Looping Concept for Air Separation at High Temperatures† Behdad Moghtaderi* Priority Research Centre for Energy, School of Engineering (Chemical Engineering), Faculty of Engineering and Built Environment, The University of Newcastle, Callaghan, New South Wales 2308, Australia Received May 29, 2009. Revised Manuscript Received October 27, 2009
A novel approach, supported by preliminary experimental and theoretical data, is proposed here for separation of oxygen from air. The approach is based on the chemical looping concept and is inherently simple and costeffective, relying on the cyclic oxidation (i.e., combustion) and reduction of a metallic oxide as a means of separating oxygen from air. Thermodynamic calculations and preliminary experiments were carried out on oxides of Cu (CuO/Cu2O), Mn (Mn2O3/Mn3O4), and Co (Co3O4/CoO) to ascertain their feasibility for the proposed air-separation method. Results confirmed the feasibility of the approach and showed that Mn and Co oxide systems and their mixtures were, in particular, quite suited for air-separation applications.
temperature. However, cryogenic processes are generally expensive, owing to the energy intensity of the process.1-3 Conventional adsorption methods [e.g., pressure swing adsorption (PSA)] of producing oxygen rely on selective physical adsorption of O2 (or N2) on the internal pores of a high-surface-area adsorbent material. Adsorption plants operate in a cyclic manner, with the basic steps being adsorption (i.e., O2 or N2 removal from air) and regeneration (i.e., release of O2 or N2 form the saturated adsorbent material). The specific power consumptions of PSA and VPSA plants are not much lower than their cryogenic counterparts.1-3 Membranes rely on a barrier film to separate O2 from air. The film allows for selective permeation of O2 and can be made from a host of different materials.3 More advanced membrane systems, such as ITMs, allow for the rapid transfer of oxygen ions, achieving fluxes that are orders of magnitude higher than polymeric membranes. Perovskite membranes (e.g., La1-xAxCo1-yFeyO3-I) have also been employed in membrane reactors for in situ oxygen generation.8 In such systems, oxygen is stored in the oxide lattice during the sorption step and released during the desorption step. This approach is the basis of a high-temperature air-separation technology known as ceramic autothermal recovery (CAR) developed by the BOC group.9,10 Although membranes are generally modular and can be replicated to satisfy the throughput requirements, they generate a degree of complexity in terms of system integration and installation. While membranes have been in commercial use for several decades, their application to large volumetric gas flow rates has not yet been demonstrated.11 Membrane systems also suffer from high cost of manufacture.
1. Introduction Oxygen is the second largest volume chemical produced in the world, with a 30% share of the global industrial gas market.1-3 The global demand for oxygen in 2009 is forecast to be 850 billion cubic meters, with an annual growth rate of about 6%.2 It has major commercial applications in the metallurgical industry, chemical synthesis, glass manufacturing, pulp and paper industry, petroleum recovery/refining, and health services.1-3 Emerging markets for oxygen include advanced power generation systems, such as integrated gasification combined cycle (IGCC), oxyfuel combustion, and solid oxide fuel cells (SOFCs).4-6 Oxygen is commonly produced at industrial scales through air separation using cryogenic-distillation- and adsorption-based technologies.1-3 Advanced technologies, such as membrane separation [e.g., ion-transport membrane (ITM)] and in situ air separation are also being developed for small-volume point-of-use oxygen generation.6,7 Cryogenic systems are employed in large-scale production of high-purity oxygen, while adsorption systems are employed at the lower end of the production scale and for lower oxygen purities.1-3 In cryogenic separation, air is liquefied at very low temperatures and, hence, oxygen is selectively removed from the air by distillation. The process is very effective because it can be accurately controlled by adjusting the pressure and † Presented at the 2009 Sino-Australian Symposium on Advanced Coal and Biomass Utilisation Technologies. *To whom correspondence should be addressed. Telephone: þ61-(2)4985-4411. Fax: þ61-(2)-4921-6893. E-mail: behdad.moghtaderi@ newcastle.edu.au. (1) Kerry, F. G. Industrial Gas Handbook: Gas Separation and Purification; CRC Press: Boca Raton, FL, 2007; ISBN:0849390052. (2) H€ aring, H. W. Industrial Gases Processing; Wiley-VCH Verlag GmbH and Co. KGaA: Weinheim, Germany, 2008; ISBN: 978352731685-4. (3) Castle, W. F. Int. J. Refrig. 2002, 25, 158–172. (4) Andersson, K.; Johnsson, F. Energy Convers. Manage. 2006, 47, 3487–3498. (5) Qinghua, Y.; Kniep, J.; Lin, Y. S. Chem. Eng. Sci. 2008, 63, 2211– 2218. (6) Hutchings, K. N.; Bai, J.; Cutler, R. A.; Wilson, M. A.; Taylor, D. A. Solid State Ionics 2008, 179, 442–450. (7) Zeng, Y.; Tamhankar, S.; Ramprasad, M.; Fitch, F.; Acharya, D.; Wolf, R. Chem. Eng. Sci. 2003, 58, 577–582.
r 2009 American Chemical Society
(8) Ito, W.; Nagai, T.; Sakon, T. Solid State Ionics 2007, 178, 809–816. (9) Acharya, D.; Krishnamurthy, K. R.; Leison, M.; Macadam, S.; Sethi, V.; Anheden, M.; Jordal, K.; Yan, J. Development of a high temperature oxygen generation process and its application to oxycombustion power plants with carbon dioxide capture. Proceedings of the 22nd Annual International Pittsburgh Coal Conference, Pittsburgh, PA, Sept 12-15, 2005. (10) Zhang, T.; Li, Z.; Cai, N. Korean J. Chem. Eng. 2009, 26 (3), 845– 849. (11) McKee, B. Solutions for the 21st century;Zero emissions technologies for fossil fuels. Technology Status Report, International Energy Agency (IEA), Paris, France, 2002.
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: DOI:10.1021/ef900553j
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Unconventional methods of air separation (i.e., noncryogenic and non-adsorption) have also been developed in the past.12 The earliest example is the thermal cycling of alkaline manganates for air separation, which was demonstrated for a short period in 1866 as a commercial operation.12 Processes based on absorption/desorption of barium oxide under strong vacuum have also been investigated by several researchers.12 A more recent air-separation method called “MOLTOX” was carried out by temperature swing absorption of oxygen from air using alkali metal nitrates and nitrites.13-15 The process did not lead to any commercial applications because of operational difficulties associated with handling molten salts. Electrolysis and thermo-chemical cycles for water splitting have also been studied for hydrogen and oxygen production.16,17 Over 250 thermo-chemical water-splitting cycles have been reported in the literature, although only a few have proven to be economically feasible.17 This is not surprising given that the water-splitting reaction is thermodynamically feasible at temperatures in excess of 1600 °C, requiring a complex and expensive reactor system driven by solar energy. Electrolysis of water is energy-intensive too. Integrated SOFC electrolyzers have recently been proposed to resolve this drawback.16 The throughput of such systems, although, is very low. The ever increasing demand for oxygen combined with the need to improve the energy efficiency and economic performances of air-separation methods have resulted in a continued and unabated search for alternative oxygen production methods. The objective of this study is to establish a novel approach that is inherently simple and cost-effective, relying on the cyclic oxidation (i.e., combustion) and reduction of a metallic oxide as a means of separating oxygen from air. The details are presented below.
Figure 1. Schematic of the CLAS process.
“CLAS”) incorporates the concept of oxygen decoupling into a two-step redox reaction mechanism oxidation Mex Oy -2 ðsÞ þ O2 ðgÞ f Mex Oy ðsÞ ð3Þ reduction Mex Oy ðsÞ f Mex Oy -2 ðsÞ þ O2 ðgÞ
As Figure 1 illustrates, the CLAS process works in a cyclic fashion by continuous recirculation of metal oxide particles between a set of two interconnected rectors, where oxidation (eq 3, O2 coupling) and reduction (eq 4, O2 decoupling) of carrier particles take place, respectively. The system consists of two fluidized bed reactors linked together through a loop seal to prevent gas leakage from one reactor to another. Air is fed into the oxidation reactor, so that the incoming reduced carrier particles can be regenerated to a higher oxidation state. The regenerated carrier particles, in turn, are transported back to the reduction reactor, where oxygen decoupling occurs in the presence of steam. The mixture of steam and oxygen exiting from the reduction reactor is passed through a condenser, so that steam can be fully separated from O2. The product oxygen can then be compressed for storage and delivery or directly fed to another process for on-site use (e.g., oxy-fuel combustion and IGCC). From an energy efficiency point of view, the CLAS process is quite efficient because of its low-energy demands. This is partly due to the fact that the theoretical net heat released over reactions 3 and 4 is zero. Therefore, in theory, the heat transported by the incoming carrier particles into the reduction reactor must be sufficient to support the endothermic reaction 4. Furthermore, under steady-state operation much of the heat required for production of steam and preheating of air is offset by the heat contents of the superheated steam stream, leaving the reduction reactor and the reduced air stream exiting from the oxidizer. As shown in Figure 1, this is achieved by exchanging (i) the sensible heat between various streams in a series of heat exchangers and (ii) the latent heat of phase change in a combined steam condenser/boiler unit. The additional thermal energy required to carryout the CLAS process can be provided by electrical power. Our preliminary mass and energy balance calculations suggest that the heat/power demand for the CLAS process will be much lower than that required in cryogenic systems (for details of mass and energy balance calculations, refer to section 3.2). The CLAS process, therefore, has the potential to create step-change improvements in the performance of air-separation systems.
2. Chemical Looping Air Separation Concept The air-separation method proposed here relies on a chemical principle similar to that used in the chemical looping combustion (CLC), which is commonly carried out in a two-step redox reaction (eqs 1 and 2) by circulating metal oxide particles between two connected reactors16 oxidation Mex Oy -1 ðsÞ þ 0:5O2 ðgÞ f Mex Oy ðsÞ ð1Þ reduction
ð2n þ mÞMex Oy ðsÞ þ Cn H2m ðgÞ
f ð2n þ mÞMex Oy -1 ðsÞ þ nCO2 ðgÞ þ mH2 OðgÞ
ð4Þ
ð2Þ
The conventional CLC process is suitable for gaseous fuels only. More recently, however, a three-step CLC process has been proposed for solid fuels,19,20 where through an in situ decoupling step oxygen is released into the gas phase, subsequently reacting with the particles of the solid fuel. Inspired by this new development and driven by the need to resolve the shortcomings of conventional air-separation methods, the method proposed here (chemical looping air separation, (12) Gardner, J. B. History of cryogenics in BOC. In History and Origins of Cryogenics; Scurlock, R, Ed.; Clarendon Press, Oxford, U.K., 1992. (13) Erickson, D. C. U.S. Patent 4,132,766, 1979. (14) Erickson, D. C. U.S. Patent 4,340,578, 1982. (15) Erickson, D. C. U.S. Patent 4,746,502, 1988. (16) Iora, P.; Chiesa, P. J. Power Sources 2009, 190 (2), 408–416. (17) Fishtik, I.; Datta, R. Comput. Chem. Eng. 2008, 32, 1625–1634. (18) Hossain, M. M.; de Lasa, H. I. Chem. Eng. Sci. 2008, 63, 4433– 4451. (19) Mattisson, T.; Leion, H.; Lyngfelt, A. Fuel 2009, 88, 683–690. (20) Mattisson, T.; Lyngfelt, A.; Leion, H. Int. J. Greenhouse Gas Control 2009, 3, 11–19.
3. Result and Discussion 3.1. Redox Characteristics. The approach outline in the previous section faces a number of challenges and technical risks. Among these, perhaps the most important challenge is 191
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Table 1. Parameters Employed in the Chemical-Equilibrium Calculations metal oxide system Cu Mn Co Mn/Co
T (°C)
P (kPa)
elements
650-1100 650-1000 650-1000 650-950
101.3 101.3 101.3 101.3
Cu, O, N Mn, O, N Co, O, N Mn, Co, O, N
gas-phase species
pure bulk species
O2, N2 O2, N2 O2, N2 O2, N2
CuO, Cu2O Mn2O3, Mn3O4 Co3O4, CoO
the development of effective and robust metal oxide oxygen carriers. Very much like CLC-type oxygen carriers, the carriers suitable for the CLAS process should have a number of desirable properties including18 (i) excellent redox properties in terms of reactivity, degree of conversion, selectivity, and oxygen-transfer capacity, (ii) good chemical stability to avoid degradation of redox properties during repeated redox cycles, (iii) high mechanical strength, so that particle fragmentation and sintering can be avoided, (iv) low cost, and (v) sound environmental characteristics. Additionally, the oxygen carriers for the CLAS process should be able to react reversibly with gaseous oxygen at high temperatures. This additional thermodynamic constraint is a means of differentiating oxygen carriers feasible for the CLAS process from those only suitable for common redox applications. Numerous studies have been carried out on oxides of transitional metals, such as Fe, Cu, Co, Mn, and Ni, as potential candidates for redox applications.18 Thermodynamically, oxides of Cu, Mn, and Co are more promising for oxygen decoupling (i.e., CLAS process) because of their ability to reversibly react with oxygen.19 For this reason, thermodynamic chemical-equilibrium calculations were carried out on oxides of Cu (CuO/Cu2O), Mn (Mn2O3/Mn3O4), and Co (i.e., Co3O4/CoO) as well as a 1:1 physically mixed Mn/Co oxide system to ascertain its feasibility for the CLAS process. The specific objective was to determine the equilibrium partial pressure (EPP) of oxygen for each metal oxide system. As shown in the latter parts of this paper, the knowledge of EPP is vital in setting up the operational envelop of the CLAS process for a given metal oxide system. Equilibrium calculations were performed using the software package COSILAB (Combustion Simulation Laboratory), version 2.3, developed by SoftPredict (Rotexo Beteiligungs-GmbH, Germany). There are several different approaches for solving chemical-equilibrium problems. The approach adopted here was to minimize the Gibbs free energy of the heterogeneous system under consideration. For a given set of pressure and temperature, this was achieved by adjusting species mole numbers, so that the Gibbs function (G) defined by eq 5 is minimized " # I X P 0 ð5Þ ni gi ðTÞ þ RT ln Xi þ RT ln 0 G ¼ P i ¼1
bulk mixture species
Mn2O3, Mn3O4, Co3O4, CoO
Figure 2. Results of the chemical-equilibrium calculations.
this method commonly known as the element-potential method21 was used in the present chemical-equilibrium analysis. The calculations involved solving a total of I þ J equations for the I unknown species mole numbers (ni) and J unknown element-potential multipliers (λj). These equations can be expressed using the following general equations, in which I and J are the total number of species and elements present in the heterogeneous mixture, respectively " # J X g0i ðTÞ P þ ln Xi þ ln 0 , i ¼ 1, :::, I ð6Þ aji λj ¼ P RT j ¼1 I X
aji ni ¼ n~j ,
j ¼ 1, :::, J
ð7Þ
i ¼1
where aji is the number of j elements in species i and n~j is the element mole number of j element in the system (i.e., 6.023 1023 multiplied by the number of atoms of type j present in the system). The above approach allows for different types of species to be considered, including gas-phase species, surface species, pure bulk species, and bulk mixtures. The following elements and species in Table 1 were employed in this study. The results of chemical-equilibrium calculations for the metal oxide systems listed in Table 1 are illustrated in Figure 2, where the EPP of oxygen has been plotted against temperature. As shown in Figure 2, all three oxide systems exhibit similar and promising trends. The Mn and Co systems, in particular, appear to be more favorable because of their lower equilibrium temperatures, which are advantageous from the energy-efficiency point of view and avoiding sintering. The Co system has the added advantage of featuring a narrow range of equilibrium temperatures between 750 and 900 °C. Note that a narrow range of equilibrium temperatures implies that the full range of partial pressures can be covered by a limited amount of heating and/or cooling. The minimum equilibrium temperature for the Co system, however, is relatively high, and as such, it cannot be employed at temperatures below 750 °C.
In eq 5, g0i is the molar-based Gibbs free energy of species i at the reference state, I is the total number of species in the heterogeneous mixture of metal oxide and gaseous species, ni is the mole number of species i, P is the pressure, P0 is the pressure at the reference state, R is the universal gas constant, T is the temperature of the mixture, and Xi is the mole fraction of species i. Minimization of Gibbs function is generally achieved by employing the method of the Lagrangian multiplier to ensure that the elemental conservation is met. A variant of (21) Reynolds, W. C. The element-potential method for chemicalequilibrium analysis: Implementation in the interactive program STANJAN. Technical Report, Stanford University, Stanford, CA, 1986.
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The Mn system, on the other hand, can be used at temperatures as low as 650 °C, but its range of equilibrium temperatures is quite wide (650-1000 °C), demanding a sizable heating and/or cooling load to cover the full range of partial pressures of interest. The separate set of calculations carried out for the 1:1 physically mixed mixture of the Mn/Co metal oxide system showed that the drawbacks of the Mn and Co systems can be notably minimized when pure oxides are mixed. As can be seen from Figure 2, while unlike the Co system, the Mn/Co mixture can be used at temperatures around 650 °C, it does not possess a temperature range as wide as that of the Mn system, which is its main advantage. As noted earlier, the successful execution of the CLAS process also largely depends upon our ability to exploit the reversible nature of the reactions occurring in both reactors. This can be best accomplished by manipulating the balance between the EPP and actual partial pressure (APP) of oxygen over the metal oxide. According to Le Chatelier’s principle, if a chemical system in equilibrium is disturbed, it will adjust itself to restore equilibrium. In a system with the equilibrium constant Kp and reaction quotient Qp (this quantity is calculated by the same expression as Kp but using actual conditions), the reaction will shift from products to
reactants if Qp > Kp. Conversely, if Qp < Kp, the reaction will proceed from reactants to products. This simple principle is actively employed in the operation of the CLAS process. For example, as shown in Figure 3, for the CuO/Cu2O system, the EPP is 10% at 1000 °C. Therefore, the APP in the oxidation reactor outlet can be maintained at levels of about 12% to ensure that newly regenerated metal oxide carrier particles do not prematurely release their oxygen content before being transported into the reduction reactor. Note that, in this case, (Qp = APP-1 = 12-1) < (Kp = EPP-1 = 10-1). Likewise, the spontaneous release of oxygen from carrier particles in the reduction reactor can be ensured by keeping the APP at around 5%, which is quite sufficient to satisfy the condition: (Qp = APP = 5) < (Kp = EPP = 10). The APP in the reduction reactor can be easily adjusted by controlling the flow rates of the incoming steam and outgoing steam/O2 mixture. The steam, therefore, is not merely a carrier gas but, most importantly, a means of controlling the partial pressure of O2 in the reduction reactor. Furthermore, unlike other carrier gases, such as helium and carbon dioxide, steam can be easily separated from the steam/ oxygen mixture simply by condensation. 3.2. Mass and Energy Balance Calculations. Mass and energy balance calculations were performed for Cu-, Mn-, and Co-based metal oxide systems using the process simulation package ASPEN-HYSYS, version 7.1. The HYSYS flow-sheet representation of the CLAS process is shown in Figure 4, where solid arrows represent material streams, whereas open arrows show energy streams. Separate subflow sheets have been included into the main flow sheet for the oxidation and reduction reactors, which are assumed to be Gibbs-type reactors. Oxygen carrier particles were defined as hypothetical solids, and their properties were imported into the simulation platform. HYSYS calculations for the Cu (CuO/Cu2O) system were performed at atmospheric pressure over a range of temperatures between 850 and 1000 °C, assuming a molar flow rate of 1 kmol/h for air (i.e., air intake material stream in Figure 4). The mole numbers of steam and oxygen at each reaction temperature (see Table 2) were determined on the basis of the method outlined in Figure 3. Panels a-d of Figure 5 also demonstrate typical temperature profiles in the air preheater and heat exchangers EX-01 and EX-02 (this particular set corresponds to a reaction temperature of 1000 °C).
Figure 3. Determination of the APP in the oxidation and reduction reactor outlets using the EPP.
Figure 4. HYSYS process flow-sheet representation of the CLAS process.
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Table 2. Parameters Used in the HYSYS Calculations for the Cu (CuO/Cu2O) Metal Oxide System T (°C) of the oxidation reactor
oxygen EPP from Figure 2 (%)
oxygen partial pressure in the oxidizer (%)
oxygen partial pressure in the reducer (%)
oxygen molar flow (kmol/h)
850 900 950 1000
0.5 1.5 4.5 10.0
1.0 2.0 10.0 12.0
0.2 1.0 3.0 5.0
0.20 0.19 0.11 0.09
stack reduced air molar steam molar flow temperature flow (kmol/h) (kmol/h) (°C) 0.80 0.81 0.89 0.91
99.80 18.81 3.56 1.71
28 28 28 28
Figure 5. Plots of axial temperature profile for the (a) air preheater, (b) heat exchanger EX-01, (c) heat exchanger EX-02, and (d) enlarged view of section A of heat exchanger EX-02.
constrains (see section 3.1) impose lower oxygen molar flows through the system as the reaction temperature is increased (see also column 5 in Table 2). Figure 6 illustrates the plot of specific power versus reaction temperature for the CuO/Cu2O metal oxide system. The specific power (SP) has been calculated from eq 8
Table 3. Summary of HYSYS Results for the Cu System for 1 kmol/h of Air T (°C) of the oxidation reactor
Qnet [kW/ (kmol/h)air]
oxygen production [m3/(kmol/h)air]a
power demand (kW/m3n)
850 900 950 1000
0.2724 0.2202 0.1098 0.0900
4.89 4.65 2.69 2.20
0.053 0.047 0.041 0.041
a
SP ¼ Qnet =½VO2 st
At normal conditions (25 °C and 101.3 kPa).
ð8Þ
where Qnet is the net input power to the process and VO2 is the volume of oxygen produced per kmol/h of air. The subscript “st” refers to the fact that VO2 is expressed under standard conditions (25 °C and 101.3 kPa). As can be seen from Figure 6, the specific power over the range of temperatures studied here can be approximated by a second-order polynomial, which plateaus at temperatures around 950 °C. The decrease in the specific power for temperatures greater than 850 °C is primarily due to the drop in Qnet as a result of lower steam demands at higher reaction temperatures (see section 3.1 and Table 2).
Table 3 summarizes the results of the heat and mass balance analysis for the CuO/Cu2O system. Because much of the heat required for the CLAS process is provided through waste heat recovery from various parts of the flow sheet, the net input power (Qnet) is relatively small. It is evident from Table 3 that Qnet consistently decreases as the reaction temperature is increased. The volume of product oxygen also diminishes as the reaction temperature is increased. This can be assigned to the fact that equilibrium 194
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Figure 6. Specific power versus reaction temperature.
Figure 7. Plots of the steam molar flow and specific power versus VO2.
It is interesting to note that, for the present case, the specific power varies between 0.041 and 0.053 kW h/m3n, with an average value of 0.045 kW h/m3n. This is about 11% of the specific power of conventional cryogenic systems,3 which typically require 0.4 kW h per cubic meters of O2 produced (i.e., 0.4 kW h/m3n). More advanced cryogenic systems3 due to enter the market by 2012, however, are expected to reach specific powers in the vicinity of 0.3 kW h/m3n. Such a specific power is still 6.7 times greater than the average specific power for the CLAS process. Even if heat losses to the ambient are incorporated into the energy balance calculations, the average specific power of the CLAS process will rise to about 0.08 kW h/m3n or just 26% of the specific power of an advanced cryogenic air-separation system. Mass and energy balance calculations for Mn and Co metal oxide systems (not presented here) led to very similar conclusions to those outlined above for the Cu system, confirming the energy-efficient nature of the CLAS process. Figure 7 shows plots of the steam molar flow rate and specific power against the volume of oxygen produced per kmol/h of air entering the process (i.e., VO2). The quantity of steam used in the process is an important design consideration, which underpins not only the physical dimensions of the reduction reactor and, hence, its capital cost but also, more importantly, the heating and cooling demands. A low steam molar flow rate is generally preferred because it corresponds to a relatively small specific power (Figure 7). However, as Figure 7 indicates on a logarithmic scale, the molar flow rate of steam rises linearly with VO2, implying that in real terms very large quantities of steam may be required to maximize VO2. For example, a steam molar flow rate of 99.8 kmol/h is required at 850 °C for the production of 4.89 m3/h of oxygen per kmol/h of air, whereas only 1.71 kmol/h of steam is needed at 1000 °C to produce 2.2 m3/h of oxygen per kmol/h of air (see also Tables 2 and 3). A compromise must be made to balance the need for minimization of the steam flow rate and specific power against the need for maximization of VO2. Given that the steam molar flow curve does not show any minima, the above compromise can be best made at the point for which the slope of the specific power curve is zero. According to Figure 6, at this optimum point for the Cu system, we have VO2 = 3.0 m3/h, steam flow rate = 5 kmol/h, and SP = 0.03 kW h/m3n. The corresponding reaction temperature for the
optimum point is 942 °C, which has been calculated by interpolation of relevant data in Table 3. 3.3. Experimental Measurements. The chemical-equilibrium calculations of section 3.1 highlighted that the 1:1 mixture of Mn/Co metal oxide systems has more favorable redox characteristics than its parent materials for the CLAS process. This particular mixture was selected for experimental studies discussed below. The 1:1 mixture of Mn and Co oxide systems used in experimental work was prepared by physical mixing of Mn2O3 and Co3O4 particles in a 1:1 weight ratio. Particles of each species had been previously prepared according to the method outlined below using three different binders (i.e., support), namely, yttriastabilized zirconia (YSZ; 92% ZrO2 and 8% Y2O3), alumina (Al2O3), and TiO2. The preparation of mono-species metal oxide samples typically initiated by the direct mixing of commercial CuO, Mn, and CO powders (SigmaAldrich) with the binder at a 3:2 weight ratio (i.e., particle/support). Distilled water was then added to the particle/support mix to form a paste. The paste was dried in an oven at 105 °C for 36 h to free up the capillary water. The dry paste was then calcined for 5 h under N2 in a high-temperature furnace. The calcined sample was pulverized in a ball mill and sieved to a particle size range of 90-106 μm. Experiments were performed in a bench-scale unit comprising a small electrically heated quartz tube reactor, a gas analysis train [comprising gas manifolds, a gas chromatography/mass spectrometry (GC/MS) gas analyzer, and gas bottles] for online monitoring of the oxygen concentration, and a data acquisition/control system. The quartz reactor had an inner diameter of 15 mm and a length of 500 mm. The reactor was fitted with a porous plate sample holder placed 200 mm away from the reactor inlet. Experiments were systematically carried out on different metal oxide systems at temperatures between 650 and 850 °C at atmospheric pressure. A typical experiment consisted of several reduction and oxidation cycles to allow for the sample to stabilize. The data set corresponding to the final redox cycle was used for determination of reaction properties because the first cycle was usually slower than the others. The experimental procedure involved the following steps: At first, a charge of 5 g of metal oxide was loaded into the reactor and the heating system was turned on. The reaction 195
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temperature was set to a predetermined level, and the reactor was then purged with nitrogen to prevent any premature reaction. Upon reaching the desired temperature, the nitrogen flow was turned off and, instead, steam was introduced. The metal oxide sample was then allowed to react with steam until the oxygen concentration stabilized (verified through a micro-GC reading). At this point, the flow of steam was stopped and a purge of N2 was introduced for about 1 min to avoid the mixing of air and steam. Air was then introduced to commence the oxidation process and, thereby, regenerate the reduced metal oxide particles. The oxidation process was seen to full completion by monitoring the oxygen content of air at the reactor outlet. During the reduction phase of the cycle, the steam flow rate was adjusted, so that even at 100% conversion, the condition of Qp < Kp could be met. For example, at 800 °C, the APP was maintained at 1.5% just under the EPP of 2%, corresponding to the 1:1 mixture of Mn/Co metal oxide systems (see Figure 3). The instantaneous measurements of the oxygen partial pressure during the reduction of all three 1:1 mixtures have been plotted in Figure 7 at an actual to equilibrium partial pressure ratio (APP/EPP) of 75%. Clearly, all three mixtures are capable of releasing oxygen into the gas phase. However, only the mixture with YSZ support approaches the APP, indicating the mixed oxide system has reached almost full conversion. Given that the operating conditions (i.e., pressure, temperature, and heating rate) were identical for all cases studied, the observed discrepancies in Figure 4 can be attributed to the impact of particle transformation phenomena (i.e., structural, morphological, and compositional changes during reaction) on the reactivity of carrier particles. It is hypothesized here that the migration of Meyþ ions provides paths for oxygen vacancies to migrate from the surface to the bulk (during oxygen decoupling) and the bulk to the surface (during oxygen coupling), leading to a degree of oxygen vacancy diffusion. It is this mechanism that is influenced by the particle transformation phenomena (that is the combined effects of structural, morphological, and compositional changes). Similar reduction experiments to that outlined above were carried out for several other temperatures between 650 and 850 °C. The APP/EPP ratio was kept at 75% to maintain the consistency with the results shown in Figure 8 for the set-point temperature of 800 °C. Additional experiments were also carried out at a 50% actual to equilibrium partial pressure ratio (APP/EPP = 50%) to investigate the effect of the oxygen concentration during the reduction stage. The measured reaction times for 95% conversion of the oxide system have been plotted in Figure 9A for a range of different set-point temperatures at both APP/EPP ratios of 75 and 50%. The average reaction rate for 95% conversion (ravg expressed in percentage per second or %/s) has also been plotted against the reaction temperature in Figure 9B for APP/EPP ratios of 75 and 50%. Clearly, the conversion rate and temperature have a very strong correlation. At low temperatures (e.g., ≈650 °C), approximately 3 min (≈170 s) is needed to reach 95% conversion, while the corresponding conversion time at the highest set-point temperature of 850 °C is about 20 s. The average rate of reaction increases with the temperature in a parabolic fashion (i.e., for APP/ EPP of 75%, ravg = 4.00 10-5T2 - 3.68 10-2T þ 8.38, with R2 = 0.9984 and for APP/EPP of 50%, ravg = 5.00 10-5T2 - 4.41 10-2T þ 10.06, with R2 = 0.9984).
Figure 8. Instantaneous oxygen partial pressure measurements at 800 °C.
Figure 9. (A) Conversion time and (B) average conversion rate during reduction of a 1:1 mixture of Mn and Co metal oxide systems.
The comparison of the plots corresponding to APP/EPP of 75 and 50% in panels A and B of Figure 9 shows that the increased difference between the actual and equilibrium concentrations of oxygen (i.e., greater difference between the actual and equilibrium partial pressures) promotes the 196
Energy Fuels 2010, 24, 190–198
: DOI:10.1021/ef900553j
Moghtaderi
characteristics. The solid recirculation rate is defined as the mass flow rate of solids within a chemical looping system, while the solid inventory refers to the total hold up of solid particles within the system. Generally speaking, in chemical looping applications, low recirculation rates and small inventories are preferred because they lower the demand for metal oxide particles, in turn, improving the cost effectiveness of the process. The solid recirculation rate and solid inventory for the 1:1 mixture of Mn/Co metal oxide systems were calculated using the conversion times and rates (both reduction and oxidation) reported in section 3.4 as well as equilibrium data presented in Figure 2. Given that the reduction rates for the Mn/Co mixture were lower than their oxidation counterparts (see section 3.3 and Figures 9 and 10 for details), m_ s was calculated on the basis of reduction data. Typically, the mass contents of Mn2O3, Co3O4, and binder in the 5 g samples were 1.5, 1.5, and 2.0 g, respectively. These correspond to 9.5 10-6 kmol of Mn2O3 and 6.23 10-6 kmol of Co3O4 in the sample. The reactions taking place during the reduction of the Mn/Co mixture are ð9Þ 6Mn2 O3 f 4Mn3 O4 þ O2
spontaneous release of oxygen as indicated by lower conversion times and higher conversion rates of plots associated with APP/EPP = 50%. However, the greater the difference between the actual and equilibrium O2 concentrations, the lower the overall cumulative yield of oxygen. Therefore, to strike a balance between the conversion rate and yield, the APP/EPP ratio has to be optimized. On the basis of a series of preliminary assessment, the optimum APP/EPP for the 1:1 mixture of Mn and Co oxide systems was found to be about 68%. The oxidation process for the Mn/Co oxide system investigated in this study was found to be generally faster than the reduction process discussed above. Figure 10 presents the plots of the conversion time and average conversion rate during the oxidation of the 1:1 mixture of Mn/Co metal oxide systems. The APP/EPP for this particular set of results is 1.5 (i.e., 150%) which corresponds to redox cycles with reduction APP/EPP ratios of 75% (see panels A and B of Figure 9). Comparisons of the conversion time curves in Figures 9A and 10 reveal that on average the oxidation time for the Mn/Co system is 27% shorter than its reduction time. Similarly, the comparisons of ravg curves in Figures 9B and 10 show that ravg for oxidation is between 30 and 80% (an average of about 50%) higher than corresponding reduction rates. The mismatch between reduction and oxidation rates may lead to operational problems (most notably in terms of the solid circulation rate and solid inventory), which have to be investigated further. The above data also imply that system characteristics, such as the rate of solid recirculation for the CLAS process, should be estimated on the basis of the slower reduction process (see section 3.4 for details). 3.4. Solid Recirculation Rate and Solid Inventory. The solid recirculation rate (m_ s) and solid inventory (Ms) of a given chemical looping system are among its main performance
2Co3 O4 f 6CoO þ O2
ð10Þ
Assuming 95% conversion, the amount of oxygen produced by Mn2O3 and Co3O4 content of the 5 g sample are 1.5 10-6 and 2.96 10-6 kmol, respectively, or a combined total of 8.93 10-7 kmol/g of the sample. For the Mn/Co mixture, the partial pressure of oxygen in the oxidation and reduction reactors can be estimated from the EPP (Figure 2) using the method outlined in section 3.1, in turn, allowing us to determine the molar and volumetric flow rates of the oxygen product as a function of the reaction temperature. The solid recirculation rate is then calculated from ~ O2 =ð8:93 10-7 Þ m_ s ¼ m
ð11Þ
~ O2 is the molar flow rate of the oxygen product where m and the constant 8.93 10-7 is the value calculated earlier as
Figure 10. Conversion time and average conversion rate during oxidation of a 1:1 mixture of Mn and Co metal oxide systems (APP/ EPP = 1.5).
Figure 11. Plot of solid inventory against the reaction temperature for the Mn/Co metal oxide mixture.
Table 4. Summary of Solids Recirculation and Inventory Results for the Mn/Co Mixture Assuming a Molar Flow Rate of 1 kmol/h for Air T (°C)
O2 EPP
O2 partial pressure (reduction)
O2 partial pressure (oxidation)
product O2 molar flow (kmol/h)
product O2 volumetric flow (m3/h)
700 750 800 850
0.00187 0.00633 0.02000 0.07030
0.0015 0.0050 0.0150 0.0600
0.003 0.010 0.030 0.100
0.207 0.200 0.180 0.110
5.1 4.9 4.5 2.7
197
solid circulation solid inventory rate (kg/s) tT (s) (kg) 0.064 0.062 0.056 0.034
121 79 53 38
7.8 4.9 3.0 1.3
Energy Fuels 2010, 24, 190–198
: DOI:10.1021/ef900553j
Moghtaderi
Table 5. Summary of Solid Inventory and Reactor Volume Calculations for Two Hypothetical Product Capacities of 138 000 and 277 000 m3n/h O2 production rate = 138 000 m3n/h
O2 production rate = 277 000 m3n/h
T (°C)
solid inventory (ton)
combined reactor volumes (m3)
solid inventory (ton)
combined reactor volumes (m3)
700 750 800 850 average
212 139 92 66 127
73 48 32 23 44
426 280 185 133 256
146 95 63 46 88
the number of moles of oxygen produced per gram of the metal oxide mixture. The solid inventory (Ms) can be approximated using the following expression: ð12Þ Ms ¼ m_ s tT
the bottom row of Table 5, the height of the reactors can be estimated. In the case of the 138 000 m3n/h production rate, the oxidation and reduction reactors will each be 7 m high. In the case of the 277 000 m3n/h production rate, the oxidation and reduction reactors will each be 14 m high. Alternatively, two 7 m high oxidation and two 7 m high reduction reactors can be employed. These dimensions are quite small when compared to the footprints of typical cryogenic systems.
in which tT denotes the sum of reduction and oxidation conversion times. The value of Ms determined from the above approximate approach is a conservative estimate because the use of eq 12 implies that the solid recirculation rates and solid densities during both reduction and oxidation are identical. However, as noted earlier, the oxidation rate is generally higher than the rate of reduction, meaning that m_ s for oxidation is generally lower than its reduction equivalent. Table 4 summarizes the results of the solid recirculation rate and solid inventory calculations carried out based on the above approach for the Mn/Co mixture. Calculations were performed assuming an inlet molar air flow of 1 kmol/h. As can be seen, the recirculation rate appears to decrease as the reaction temperature is increased. However, it should be noted that the volumetric flow rate of the product oxygen also decreases as the reaction temperature is increased. Therefore, if the solid recirculation rate is expressed per unit volume of oxygen produced, its value does not strongly depend upon the reaction temperature. This is shown in Figure 11, where the normalized solid circulation rate (i.e., solid recirculation per cubic meter of product oxygen) has been plotted against the reaction temperature. The same statement, however, cannot be made about the solid inventory, even if it is normalized by the volume of oxygen produced. The solid inventory shows a much stronger correlation with the reaction temperature. For example, an increase of 50 °C in the reaction temperature can lead to a drop as big as 60% in the solid inventory. The solid inventory is a particularly important parameter because it underpins the physical dimensions of the oxidation and reduction reactors. To put the above results into perspective, the data presented in Figure 10 were used to determine the solid inventory and the physical dimensions of two hypothetical CLAS plants with production capacities identical to two classes of conventional cryogenic-based airseparation systems,4 one with an O2 production of 138 000 m3n/h and the other with an O2 production of 277 000 m3n/h. Results have been summarized in Table 5, where the combined reactor volume has been assumed to be 2.0 times larger than the actual volume of solid particles. Reactor volume calculations were carried out using a bulk density of 5850 kg/ m3 for the Mn/Co metal oxide mixture. Assuming a diameter of 2 m for reactors and using the average figures reported in
Conclusions The conventional technologies for oxygen production while well-established suffer from high-energy demand and cost, which can no longer be tenable under the current economic, energy, and environmental constraints. The proposed CLAS process with its features, such as low energy demand, inexpensive manufacture, and simplicity, will offer an effective solution to the above problem and, thus, should be an attractive proposition to industry players who are becoming increasingly concerned with their untenable position. A technology option, such as the CLAS process, may also hold the key to successful commercial-scale deployment of emerging fossil fuel energy technologies, such as oxy-fuel combustion and IGCC, indirectly leading to a range of additional environmental benefits. Our preliminary results presented in this paper confirm the feasibility of the CLAS process for air-separation applications. Theoretical and experimental studies carried out on Cu (CuO/Cu2O), Mn (Mn2O3/Mn3O4), and Co (i.e., Co3O4/ CoO) oxide systems and their mixtures have ascertained their effectiveness for chemical looping air separation. Among these, however, the 1:1 mixture of Mn and Co oxide systems was found to be more effective than the other oxide systems studied. The actual and operating partial pressures of oxygen and the operation temperature were identified as the major controlling parameters. Conversion rates of up to 5%/s can be achieved by careful selection of the temperature and APP/ EPP ratio. Preliminary cost analysis showed that the specific power consumption of the CLAS process is about 0.045 kW h per cubic meter of air produced, which is approximately 11% of the specific power consumption of conventional cryogenic air-separation systems. Acknowledgment. The author acknowledges the financial support provided to this project by the Australian Research Council, the Newcastle Port Corporation, and the Priority Research Centre at the University of Newcastle (Australia). The author also thanks the staff at USyd for assisting him with some of the experimental measurements.
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