Chemical-Looping Combustion of Petroleum Coke ... - ACS Publications

Chemical-looping combustion with solid fuel was investigated in a 10 kWth chemical-looping combustor, using a petroleum coke as the fuel and ilmenite,...
0 downloads 0 Views 3MB Size
Download PDF
Energy Fuels 2009, 23, 5257–5268 Published on Web 09/23/2009

: DOI:10.1021/ef900464j

Chemical-Looping Combustion of Petroleum Coke Using Ilmenite in a 10 kWth Unit-High-Temperature Operation Nicolas Berguerand* and Anders Lyngfelt Department of Energy Conversion, Chalmers University of Technology, S-412 96 Goteborg, Sweden Received May 15, 2009. Revised Manuscript Received August 10, 2009

Chemical-looping combustion with solid fuel was investigated in a 10 kWth chemical-looping combustor, using a petroleum coke as the fuel and ilmenite, which is an iron titanium oxide, as an oxygen carrier. The fuel reactor (FR) was fluidized by steam to gasify the coke, and the oxygen carrier reacts with the gasification products CO and H2. The FR temperature was normally 950 °C. Testing involved the variation of operational parameters such as particle circulation, fluidizing velocity in the FR, and, most important, the FR temperature. In particular, successful testing was performed at a FR temperature of 1000 °C without any operation difficulty and the positive effect of temperature on carbon capture and solid fuel conversion was verified. The oxygen demand corresponds to the fraction of oxygen lacking to achieve full gas conversion and varied over a range of 27%-35%, because of the presence of CH4, CO, H2, and H2S at the FR outlet. During these tests, the CO2/CO ratios that were usually reached in the FR were in the range of 8-9 at stable operation. Most of the oxygen demand is associated with volatiles never in contact with oxidized particles, because the volatiles are released before the fuel particles reach the bed. Indeed, investigation based on gas concentration measurements during transitions phases, which correspond to the starting and stopping of fuel addition, indicate that as much as 80% of the total oxygen demand can be associated with the volatiles. The oxygen demand for the actual char oxidation is estimated to be 5%-9%, if sulfur is excluded.

more-controversial nuclear power. But, considering the abundance and comparatively low price of fossil fuels and screening the state of evolution of the other technologies, it seems unlikely that it would be possible to stop using fossil fuels in the near future. Thus, processes using CO2 capture and storage could give a most important contribution and provide a bridging technology for reducing those emissions.6 Chemical-looping combustion, commonly abbreviated as CLC, is an innovative combustion technology with inherent CO2 capture. In the CLC process, the CO2 produced in the combustion chamber is never mixed with the nitrogen in the air. Instead, a circulating metal oxide is used as the oxygen carrier, providing the oxygen that is needed for the combustion directly to the fuel, avoiding the presence of nitrogen in the exhaust flue gas (see Figure 1). The oxidation of the oxygen carrier occurs in the so-called air reactor (AR), while it is reduced in the fuel reactor (FR). MexOy/MexOy-1 are the symbols used in this paper for the oxidized/reduced form of the oxygen carrier. The general reaction between the fuel and the metal oxide reads ð2n þ mÞMex Oy þ Cn H2m f ð2n þ mÞMex Oy -1

Introduction The industrial revolution, which began in the 19th century, has been accompanied with growing drawbacks on the climate on Earth.1-3 The connection between the amounts of atmospheric CO2 and the global average temperature was first discovered by Arrhenius in the late 19th century.4 Since then, it has been claimed that the extensive use of fossil fuels for heat and power consumption or transportation has led to tremendous amounts of greenhouse gas (GHG) emissions into the atmosphere. In particular, the CO2 concentration in the atmosphere has now reached unprecedented levels, compared to that observed over the past 400 000 years. This is probably the most important cause of the increase in global temperature and the climate change currently witnessed, and this can have unpredictable consequences for life on Earth. The emissions must be reduced in one way or another if we do not want to face even more severe problems in the future.5 Different alternatives for energy production using nonfossil fuels have been proposed and are already available, such as wind power, hydropower, solar energy with photovoltaic panels, or the

þ mH2 O þ nCO2

*Author to whom correspondence should be addressed. Tel.: þ 4631-7725241. Fax: þ46-31-7723592. E-mail: [email protected]. (1) Weare, B. C. The Possible Link Between Net Surface Heating and El Ni~ no. Science 1983, 221, 947–949. (2) Arendt, A.; et al. Rapid Wastage of Alaska Glaciers and Their Contribution to Rising Sea Level. Science 2002, 297, 382–386. (3) Beniston, M.; Diaz, H. F. The 2003 heat wave as an example of summers in a greenhouse climate? Observations and Climate Modes Simulations for Basel, Switzerland. Global Planetary Change 2004, 44 (1-4), 73–81. (4) Arrhenius, S. On the Influence of Carbonic Acid in the Air upon Temperature of the Ground. Philos Mag. 1886, 41, 237–277. (5) Climate Change 2007. In Intergovernmental Panel on Climate Change; Cambridge University Press: Oxford, U.K., 2007. r 2009 American Chemical Society

while the oxygen carrier oxidation in the AR is 1 Mex Oy -1 þ O2 f Mex Oy 2

ð1Þ

ð2Þ

In the FR, CO2 and H2O are produced but these are never mixed with the nitrogen in the air, as is the case in normal (6) Carbon Capture and Storage. In Intergovernmental Panel on Climate Change; Cambridge University Press: Oxford, U.K., 2005.

5257

pubs.acs.org/EF

Energy Fuels 2009, 23, 5257–5268

: DOI:10.1021/ef900464j

Berguerand and Lyngfelt

Figure 1. Schematic picture of the CLC process. Two interconnected fluidized bed reactors, one an air reactor (AR) and one a fuel reactor (FR), with circulating oxygen carrier particles.

Figure 2. Design proposed by Lyngfelt for a chemical-looping combustor for natural gas. Legend: (1) air reactor (AR) and riser, (2) cyclone, and (3) fuel reactor (FR).

combustion with air. This avoids further cost and energy penalty in gas separation processes. By condensing the steam, almost pure CO2 can be obtained. At the AR outlet, the flue gas contains nitrogen and some unused oxygen. Depending on the fuel and the oxygen carrier, reaction 1 is often endothermic while reaction 2 is exothermic. The total amount of heat resulting from reactions 1 and 2 is the same as for a normal combustion where the fuel is in direct contact with the oxygen in the air. CLC was first introduced in a 1954 patent as a way to produce pure CO2 from fossil fuels 7 and later tested for gaseous fuel combustion in different laboratory scale prototypes and larger units with two interconnected fluidized beds. Such a reactor design shown in Figure 2 was proposed by Lyngfelt et al.8,9 and provides a good contact between the gases and the oxygen carrier, which is crucial for the process. More recently, attention has been focusing on solid fuels such as, for instance, coal or petroleum coke (see refs 10-15). Indeed, they represent the main source of CO2 emissions; considering the cost and abundance of, e.g., coal in the world, solid fuel CLC seems to be interesting. For solid fuels, the volatile compounds can still react according to reaction 1. However, because no solid-solid

reaction between the oxygen carrier and the remaining char fraction is expected to occur, the char must be gasified in the FR using steam or CO2, according to eq 3: ð3Þ C þ H2 O=CO2 f CO þ H2 =CO The synthesis gases H2 and CO can then react according to reaction 4: ð4Þ CO=H2 þ Mex Oy f CO2 =H2 O þ Mex Oy -1 Solid fuel CLC has been studied in laboratory units using different types of fuels and oxygen carriers. The fuels used were bituminous coals and petroleum cokes; however, testing with lignite has also been performed.16,17 In these tests, the oxygen carriers used were iron oxides, an iron titanium oxide called ilmenite and copper oxides. Tests were also conducted in a 10 kWth unit at Chalmers University, using ilmenite as the oxygen carrier and bituminous coals and a petroleum coke as the fuels.18,19 More recently, chemical looping with oxygen uncoupling (CLOU) has been studied as a novel method to burn solid fuels in gas-phase oxygen (see refs 20-22). In this process, the oxygen carrier directly releases its oxygen in the FR, which then reacts with the char. This gas-solid reaction is fast and avoids the time-consuming char gasification step of eq 3. The metal oxide behavior is decisive for success in CLC processes, and several properties determine a suitable oxygen carrier. It must have (i) a high reactivity toward the fuel and high reduction/oxidation rates, (ii) the ability to fully convert the fuel to CO2 and H2O, (iii) a low tendency for agglomeration

(7) Lewis, W. K., Gilliland, E. R. (S.O.D. Company). Production of Pure Carbon Dioxide, U.S. Patent 2,665,971, 1954. (8) Lyngfelt, A; Leckner, B; Mattisson, T A fluidized-bed combustion process with inherent CO2 separation-application of chemical-looping combustion. Chem. Eng. Sci. 2001, 56, 3101–3313. (9) Lyngfelt, A., Thunman, H. Construction and 100 h of Operational Experience of a 10 kW Chemical-Looping Combustor. Chapter 36 in Carbon Dioxide Capture for Storage in Deep Geologic Formations-Results from the CO2 Capture Project, Vol. 1;Capture and Separation of Carbon Dioxide from Combustion Sources; Thomas, D., Ed.; Elsevier Science: London, 2005; Chapter 36, pp 625-646. (10) Cao, Y.; et al. Reduction of Solid Oxygen Carrier (CuO) by Solid Fuel (Coal) in Chemical-Looping Combustion. Prepr. Pap.-Am. Chem. Soc., Div. Fuel Chem. 2005, 50 (1), 99–102. (11) Lyon, R. K.; Cole, J. A. Unmixed Combustion: An Alternative to Fire. Combust. Flame 2000, 121 (1/2), 249–261. (12) Pan, W.-P. et al. Application of a Circulating Fluidized-Bed Process for the Chemical-Looping Combustion of Solid Fuels. In Abstracts of Papers, 228th ACS National Meeting, Philadelphia, PA, August 22-26, 2004; American Chemical Society: Washington, DC, 2004; Paper No. 155. (13) Cao, Y.; Liu, K.; Riley, J. T.; Pan, W.-P. Application of a circulating fluidized-bed process for the chemical-looping combustion of solid fuels. Prepr. Pap.-Am. Chem. Soc., Div. Fuel Chem. 2004, 49 (2), 815–816. (14) Dennis, J. S.; Scott, S. A.; Hayhurst, A. N. In situ gasification of coal using steam with chemical looping: A technique for isolating CO2 from burning a solid fuel. J. Energy Inst. 2006, 79, 187–190. (15) Gao, Z.; Shen, L.; Xiao, J.; Qing, C.; Song, Q. Use of Coal as Fuel for Chemical-Looping Combustion with Ni-based Oxygen Carrier. Ind. Eng. Chem. Res. 2008, 47 (23), 9279–9287.

(16) Leion, H.; Mattisson, T.; Lyngfelt, A. Solid Fuels in ChemicalLooping Combustion. Int. J. Greenhouse Gas Control 2008, 2, 180–193. (17) Leion, H.; Mattisson, T.; Lyngfelt, A. The Use of Petroleum Coke as Fuel in Chemical Looping Combustion. Fuel 2007, 86, 1947– 1958. (18) Berguerand, N.; Lyngfelt, A. Design and Operation of a 10 kWth Chemical-Looping Combustor for Solid Fuels;Testing with South African Coal. Fuel 2008, 87, 2713–2726. (19) Berguerand, N.; Lyngfelt, A. The Use of a Petroleum Coke as Fuel in a 10 kWth Chemical-Looping Combustor. Int. J. Greenhouse Gas Control 2008, 2 (2), 169–179. (20) Mattisson, T.; Lyngfelt, A.; Leion, H. Chemical-Looping with Oxygen Uncoupling for Combustion of Solid Fuels. Int. J. Greenhouse Gas Control 2009, 3 (1), 11–19. (21) Leion, H., Mattisson, T. Lyngfelt, A., Combustion of a German lignite using chemical-looping with oxygen uncoupling (CLOU). The Clearwater Coal Conference ; The 33rd International Technical Conference on Coal Utilization & Fuel Systems, Clearwater, FL, June 1-5, 2008. (22) Mattisson, T.; Leion, H.; Lyngfelt, A. Chemical-Looping with Oxygen Uncoupling using CuO/ZrO2 with Petroleum Coke. Fuel 2009, 88 (4), 683–690.

5258

Energy Fuels 2009, 23, 5257–5268

: DOI:10.1021/ef900464j

Berguerand and Lyngfelt

and low attrition/fragmentation, and (iv) low production cost and low environmental impact In CLC with solid fuels, because fuel ash is likely to reduce the lifetime of the metal oxide, low-cost natural ores are considered in priority. Key Performance Criteria for CLC with Solid Fuels. CLC for solid and gaseous fuels differs in several ways. With solid fuels, two types of particles are present in the FR and the oxygen carrier must not only convert the volatiles but also the syngases produced during gasification of the remaining char. This involves important issues to consider when dealing with solid fuel CLC and evaluating the overall process efficiency. Those key performance criteria are as follows: • The solid fuel conversion in the FR (i.e., the conversion of char). This is strongly dependent on the fuel reactivity, together with the separation efficiency in any device (e.g. cyclone, recovering particles from the gas stream leaving the FR). • The carbon capture of the system. Here, the loss is defined as the amount of unburnt carbon that enters into and burns in the AR divided by the total gaseous carbon exiting the reactor system. Thus, high carbon capture is synonymous with a small loss of carbon in the solids flow from the FR to the AR. • The gas conversion in the FR (i.e., the oxidation of volatiles and syngases in contact with the oxygen carrier in the FR). • The integrity of the oxygen carrier, in terms of attrition/ fragmentation and remaining activity in relation with its cost. This point is not studied in the present publication (see ref 18 for more details). Purpose. In this study, CLC with solid fuel is investigated in a 10 kWth pilot that was built at Chalmers University, with consideration of the key performance criteria previously mentioned. The purpose of this work is (i) to study hightemperature operation (earlier tests at high FR temperature are presented in ref 23; during these tests, however, a broken porous plate led to an inadequate fluidization of the FR and the results were not fully conclusive), (ii) investigate the respective importance of char and volatiles on the oxygen demand, and (iii) verify that the particle circulation and the fluidization of the FR chamber applied correspond to proper operating conditions.

Figure 3. Top view of the fuel reactor (FR) with the particle circulation directions. Legend: 2, carbon stripper; 3, high-velocity part; and 4, low-velocity part. Note: the drawings are not to scale.

Figure 4. Front view of the fuel reactor (FR) with the particle circulation directions. Legend: 2, carbon stripper; 3, high-velocity part; and 4, low-velocity part. Note: the drawings are not to scale.

unburnt char;are elutriated; the corresponding amounts depending on the operation/fluidizing velocity. Indeed, because a certain fraction of the char is likely to remain unconverted downstream of LOVEL, the reactor system has an internal recirculation loop via a smaller cyclone called the FR cyclone, to increase the overall residence time of char particles. (3) The carbon stripper (CS), where char and oxygen carrier are separated; the unreacted char is recirculated to the low-velocity section, whereas the oxygen carrier enters the AR for a new oxidation. These notations will be used below. The operating conditions were as follows: a constant fuel flow of 655 g/h, corresponding to a thermal power of ∼6 kW. The temperature in the FR was initially set to 950 °C but was increased up to 1000 °C during some of the test periods. The temperature in the AR ranged between 800 and 950 °C for the experiments. Changes in the fluidization velocity of LOVEL and the particle circulation were made; the latter being achieved by changes in air reactor fluidizing flow or temperature changes, affecting the velocity. Because of small size, the system is enclosed in an oven and is not self-supporting, with regard to energy. During operation, the AR is fluidized with air while the loop seals, and the HIVEL and CS are fluidized with nitrogen. To prevent steam from moving up into the fuel bin and condensing, a downward flow of 20 Ln/min nitrogen is also added. The combustion chamber LOVEL is fluidized with steam that has been delivered by a steam generator. The oxygen carrier used is ilmenite, which contains iron titanium oxide (FeTiO3) with a purity of 94.3%. It has the particle size distribution (PSD) given in Table 1 and a bulk density of 2100 kg/m3. The fuel is a Mexican petroleum coke with a cut diameter (D50) of 100 μm; its composition is given in Table 2. Its sulfur content of 6.6% is rather high, but comparable to the content in the petroleum coke used in ref 19.

Experimental Section The unit and the reactor system are described in two previous publications presenting tests with ilmenite and two fuels: a bituminous coal and a pet coke.18,19 A schematic view of the different chambers in the FR and the directions of particle circulation are presented in Figures 3 and 4. Figure 5 shows a general view over the entire reactor system: ARs and FRs, particle filters, fuel feeding system, steam generator, and water seal. In particular, the FR is divided in three main chambers (see Figures 3 and 4): (1) The low-velocity section (LOVEL), which is subdivided into two chambers, where the fuel is expected to devolatilize, the char is gasified, and the volatiles and syngases react with the oxygen carrier. (2) The high-velocity section (HIVEL), which is located below the riser and where both types of particles;but especially the (23) Berguerand, N.; Lyngfelt, A. Operation in a 10 kWth ChemicalLooping Combustor for Solid Fuel;Testing with a Mexican Petroleum Coke. Energy Procedia 2009, 1, 407–414.

5259

Energy Fuels 2009, 23, 5257–5268

: DOI:10.1021/ef900464j

Berguerand and Lyngfelt

Figure 5. Schematic of the entire pilot system: (a) AR, (b) riser, (c) AR cyclone, and (d) FR.

the system and the fluidizing gas flows for the different chambers. In addition, online gas analysis allows the oxygen, carbon dioxide and carbon monoxide concentrations from both reactors to be measured. From the FR, the methane and sulfur dioxide concentrations are also measured and gas chromatography (GC) samples give the hydrogen concentration. From the fuel composition and the total concentration of carbon-containing gases at the FR outlet, the total concentration of the sulfur-containing gases (SO2 and H2S) is calculated, assuming that the S/C ratio is the same in the fresh fuel as that in the gas leaving the FR. The H2S content can be deduced, knowing the measured SO2 content. All these measures and calculated data are used to assess the operation. The “carbon flow”, where FC,FR is the flow of carboncontaining species (i.e., CO, CO2, and CH4 leaving the FR (rates given in terms of Ln/min). It is calculated using the known nitrogen flow (FN2,FR) fluidizing the high-velocity part, the carbon stripper, the FR loop seal, the sweep gas, and half of the flows for the upper and lower loops seals. The carbon flow reads FC;FR ¼ FN2 ;FR

Table 1. Particle Size Distribution of the Ilmenite diameter, Ø (μm)

wt %

Ø > 250 180 < Ø < 250 125 < Ø < 180 90 < Ø < 125 90 < Ø

7.41 44.97 38.10 8.73 0.79

Table 2. Pet Coke Analysis parameter

value Proximate Analysis

moisture ash volatiles

1.09 wt % 1.83 wt % 9.91 wt % Ultimate Analysis

composition C H S N O heating value

84.93 wt % 3.41 wt % 6.59 wt % 1.66 wt % 0.49 wt % 32.96 MJ/kg

Data Evaluation



The purpose of the data evaluation is to indicate the performance, with respect to the previously mentioned key performance criteria. During operation, the temperature in the ARs and FRs are monitored, as well as the important pressure drops in

½CO2  þ ½CO þ ½CH4  1 -ð½CO2  þ ½CO þ ½CH4  þ ½H2  þ ½H2 Sþ½SO2 Þ ð5Þ

This calculation assumes that the dry gases only contain N2, CO, CO2, CH4, H2, H2S, and SO2. Note that higher 5260

Energy Fuels 2009, 23, 5257–5268

: DOI:10.1021/ef900464j

Berguerand and Lyngfelt

It is calculated, taking into consideration the total air and nitrogen flow entering the air reactor, the expected oxygen concentration if no oxidation occurred, and the measured oxygen and CO2 concentrations at the AR outlet. FO,ARtot is the total flow of oxygen consumed in the AR for particle oxidation and char combustion (this also is expressed in terms of Ln/min). The CO2 concentration gives the part of oxygen that is not used to oxidize the metal particles but instead is sued to burn char in the AR. The expected oxygen concentration ([O2,exp]AR) takes into account the minor dilution of N2 with the loop seals. The “oxide oxygen fraction”, noted as ηOO, is another measure that is used to evaluatie the carbon capture efficiency. This number is defined by the amount of oxygen used to oxidize the oxide in the AR (FO,AR) divided by the total amount of oxygen consumed in the air reactor (i.e., FO,ARtot). Derivation using eqs 7 and 11 yields FO;ARtot - FC;AR FO;AR ¼ ηOO ¼ FO;ARtot FO;AR þ FC;AR

hydrocarbons or nitrogen compounds were not measured in this study. The gas flow in the AR outlet is noted as FAR,out and expressed as 1 - ½O2, exp AR ð6Þ FAR;out ¼ FAR;in  1 - ½CO2 AR - ½O2, meas AR where FAR,in is the total air and nitrogen flow at the AR inlet, [O2,exp]AR is the O2 concentration that is expected if no particle oxidation and char burning occurred but taking into consideration the minor nitrogen dilution from the loop seals, [O2,meas]AR is the oxygen concentration measured at the outlet, and [CO2]AR is the measured carbon dioxide concentration in the AR outlet due to char burning. The carbon flow leaving the AR is in the form of CO2 and is given by ð7Þ FC;AR ¼ FAR;out  ½CO2 AR The CO2 capture of the reactor system given by the “carbon capture efficiency”, denoted as ηCC and defined as the ratio of carbon-containing gas flow leaving the FR (FC,FR) to the total carbon-containing gas flow leaving the system (i.e., from the ARs and FRs). It reads FC;FR ð8Þ ηCC ¼ FC;FR þ FC;AR

¼

ð12Þ

Note that ηOO is dependent only on the gas concentrations in the AR outlet, which eliminates any uncertainty that is due to flows. Thus, any possible error would be associated with the gas concentration measurements. In this sense, ηOO is more reliable than the carbon capture efficiency (ηCC) in eq 8. The “actual O2/C ratio” (ΦO,act) is defined as actual oxygen needed under the testing conditions (i.e., corrected for the oxygen demand in the gas from the FR). This gives ð13Þ ΦO;act ¼ ΦO;theor ð1 - ηCC ΩOD Þ

The oxygen demand (ΩOD) is the fraction of oxygen that is lacking to achieve complete combustion of the gases leaving the FR. It is calculated as 0:5½CO þ 2½CH4  þ 0:5½H2  þ 1:5½H2 S ð9Þ ΩOD ¼ ΦO;theor ð½CO2  þ ½CO þ ½ CH4 Þ where ΦO,theor is defined as the ratio of moles of oxygen needed to burn the fuel completely per mole of carbon in the fuel. For the pet coke studied here, its analysis gives ΦO,theor = 1.15. In eq 9, the use of ΦO,theor assumes that all carbon in the fuel is converted to gas, which is an ideal case. In practice, ΦO is greater, because unconverted char, with ΦO = 1.0, leaves the FR. To include a correction for the loss of char in eq 9 would introduce uncertainties in the mass balance, because it is difficult to determine this loss versus time accurately. With the definition in eq 9, the oxygen demand is a few percent higher; however, on the other hand, this is likely compensated by the fact that higher hydrocarbons are not included. The circulation index (CI) is defined by the pressure drop measured between pressure taps located at the riser entrance and outlet multiplied by the actual gas volume flow in the AR outlet (i.e., corrected for oxygen consumed and measured temperature in the AR): TAR þ 273 Þ ð10Þ CI ¼ jΔPRISER jFAR;out ð 273

Depending on the test results for ΩOD and ηCC, different values for ΦO,act are obtained. These can be compared to the measured O2/C ratio for the tests (ΦO,meas): FO;ARtot ð14Þ ΦO;meas ¼ FC;AR þ FC;FR Using eqs 8 and 12, it is also possible to derive the relationship between ηCC, ηOO, and ΦO,act: ð15Þ ηCC ¼ 1 -ð1 - ηOO ÞΦO;act Mass Balances. Mass balances of oxygen and carbon are useful to understand the results and evaluate their quality. They involve the definitions presented above: • mass balance of oxygen: it is based on the comparison between ΦO,meas and ΦO,act during a test (i.e., between the total oxygen consumed per gaseous carbon leaving the system and the expected ratio of oxygen and carbon for the test conditions). The closer the values, the better the fulfilment of the oxygen mass balance. • the carbon mass balance: this compares the solid fuel conversion ηSF,rec. derived from the amount of elutriated char collected in the water seal to the total amount of fuel added during a given test and the calculated solid fuel conversion, ηSF, which is given by FC;FR þ FC;AR ηSF ¼ ð16Þ FC;FUEL

CI does not give the actual particle circulation (in units of kg/ min); instead, it provides a qualitative measure of the particle circulation between the reactors so that the different tests can be compared. More details on how eq 10 is derived are presented in ref 18. The “oxygen flow”, which is noted as FO,AR, is the flow of oxygen used to oxidize the metal oxide in the AR and is expressed in terms of Ln/min by eq 11:  FO;AR ¼ FAR;in ½O2;exp AR -FAR;out ½O2;meas AR þ½CO2 AR ¼ FO;ARtot -FC;AR

½O2;exp AR - ½O2;meas AR - ½CO2 AR ½O2;exp AR -½O2;meas AR - ½O2;exp AR ½CO2 AR

i.e., the ratio between the total flow of gaseous carbon leaving the FR and the total flow of carbon added to the system with the fuel feed (noted as FC, FUEL).

ð11Þ 5261

Energy Fuels 2009, 23, 5257–5268

: DOI:10.1021/ef900464j

Berguerand and Lyngfelt

If the mass balance of carbon is fulfilled, this fuel conversion and the one based on char recovery should be equal.

Table 3. First Test: Different Operating Conditions for the Five Test Periods

Results and Discussion

flow at air flow at lower-velocity fuel reactor reactor, AR section, LOVEL test period (Ln/min) temperature, TFR (°C) (Ln/min)

Four tests were performed, totalizing ∼18 h of stable operation. The results for three of them are presented below. The test not presented involved experiments at higher temperature and the results were rather similar to those in the first test below, but it was repeated because of malfunction of the steam generator. For all the figures in this section, time zero corresponds to the start of fuel feed and the small triangles above the x-axis delimit test periods corresponding to changes in the operating conditions. The temperature normally used in the FR was 950 °C, while the steam flow to LOVEL was 30 Ln/min. The fuel addition corresponded to a thermal power of 6 kW. Note that the fuel feed slowly decreases with time during every test and is directly related to the remaining amount of pet coke in the fuel feeder bin. Because methane is associated with volatiles release from fresh fuel entering the FR, proper correlation between the initial calibrated flow and methane concentrations allows the exact fuel feed to be determined throughout an experiment. This was always taken into account in the data evaluations. Finally, hydrogen concentrations measured via GC were compared to values derived from the water-gas shift equilibrium, using the equilibrium constant. For the tests presented below, results indicated that calculated values were lower, compared to the measurements and the deviation was in the range of 50%-70%, meaning that the shift equilibrium was not reached. Important parameters to assess the quality of results include the following: • The CO2/CO and SO2/Ctot ratios in the flue gas of the FR outlet. They should be as high as possible, meaning high conversion degrees of CO to CO2 and H2S to SO2. • The oxygen demand, ΩOD in eq 9, which indicates the oxygen that is lacking to convert the gases to higher degrees in the FR; this value should be as low as possible. • The CO2 capture, ηOO in eq 12, which indicates the loss of carbon to the AR (i.e., showing whether the fuel has sufficient residence time in the FR). • The solid fuel conversion (ηSF) in eq 16. First Test. The first test involved 3 h, 40 min of operation, and the purpose of the test was to investigate the effect of increased FR temperature on the system performance. During operation, changes were made in the particle circulation, the fluidization of the LOVEL section, and the FR temperature. The operating conditions for the corresponding test periods are summarized in Table 3. Figure 6 shows the temperatures in the ARs and FRs, as well as the averaged circulation index (CI), given by eq 10. Figure 7 shows the concentrations of CO, CO2, CH4, and H2 at the FR outlet, and Figure 8 represents the CO2/CO and SO2/Ctot ratios, as well as the oxygen demand in the FR, the CO2 capture (ηOO), and the solid fuel conversion (ηSF) in the system. In test periods 2 and 3, the fluidizing air flow to the air reactor was increased to ensure sufficient particle circulation (see Figure 6). In test period 4, the LOVEL fluidization velocity was increased, which led to some increase in the CO2/CO ratio, but, otherwise, no significant effect was seen (see Figure 8). The higher CO2/CO ratio could be an effect of increasing steam concentration pushing the water-gas shift equilibrium toward less CO and more H2 (see eq 17):

1 2 3 4 5

130 140 155

30

950

38 1000

Figure 6. First test: Circulation index (CI) and temperatures in ARs and FRs.

Figure 7. First test: Gas concentrations at the FR outlet.

CO þ H2 O f CO2 þ H2

ð17Þ

The effect of LOVEL fluidization is further examined in the third test. In test period 5, the temperature in the FR was increased to 1000 °C. A direct effect was the decrease in the CO2/CO ratio (see Figure 8). Again, this could be associated with a change 5262

Energy Fuels 2009, 23, 5257–5268

: DOI:10.1021/ef900464j

Berguerand and Lyngfelt

Figure 8. First test: CO2/CO and SO2/Ctot ratios, solid fuel conversion (ηSF), oxygen demand (ΩOD), and CO2 capture (ηOO).

in the water-gas shift equilibrium, which is shifted toward less H2 and more CO at higher temperatures. More important is the fact that higher temperatures led to a significant increase in the CO2 capture (from 74% to 84%). Figure 8 also suggests that solid fuel conversion increased somewhat at higher temperatures. Moreover, the oxygen demand decreased slightly to ∼27%, which means the operation was moving toward better gas conversion. The main conclusion from this first test is that the temperature increase was clearly accompanied by an increase in the CO2 capture of the system. Second Test. The second test involved 5 h, 20 min of stable operation, and the goal was to verify the repeatability of higher-temperature results and also to check that performance was similar before and after a high-temperature period. The FR temperature was first switched from 950 °C to 1000 °C at the end of test period 1 and eventually set back to 950 °C at the end of test period 2. Figures 9-11 represent the test results, similar to Figures 6-8 in the first test. Increasing the temperature in the FR to 1000 °C led to increased concentrations of CO2, CO, and H2 (see Figure 10). This is responsible for a decrease in CH4 concentration due to larger gas dilution. As seen in the first test, higher temperature also led to a decrease in CO2/CO (see Figure 11). However, no increase but rather a decrease in the oxygen demand was observed, again suggesting that the decrease in the CO2/CO ratio could be associated with the shift in equilibrium. As in the first test, the CO2 capture was considerably improved by the temperature increase (see Figure 11). Also, the increase in solid fuel conversion could be clearly observed. Decreasing the temperature in test period 3 gave a transition back to the conditions observed in test period 1. In conclusion, the increase in FR temperature from 950 °C to 1000 °C had a significant positive effect on the CO2 capture but also a positive effect on the solid fuel conversion. Also, the oxygen demand decreased somewhat at higher temperature. Third Test. The purpose of the third test was to investigate the effect of steam flow in LOVEL and involved 3 h, 30 min of stable operation at a FR temperature of 950 °C. Table 4 below summarizes the flows and corresponding steam velocities in the LOVEL section. Figure 12 represents the performance data.

Figure 9. Second test: Circulation index (CI) and temperature profiles in ARs and FRs.

Figure 10. Second test: Gas concentrations at the FR outlet.

The increase in steam flow to LOVEL in test period 2 resulted in an increase in the SO2/Ctot ratio, to 7.2; otherwise, however, no significant effects were observed. However, decreasing the steam flow in test period 3 led to a drastic decrease in performance, as seen by the CO2/CO and SO2/Ctot ratios, the CO2 capture, and the oxygen demand. A further decrease in steam flow in test period 4 led to even more decreased performance. Clearly, the system limit had been reached and it was not possible to operate properly at such a low steam flow in LOVEL. Indeed, the corresponding steam velocity was only 5 cm/s, to be compared with the particles minimum fluidizing velocity (umf = 2 cm/s). The effect of increased steam flow in test period 5 was immediate: rapidly increasing CO2 capture, as well as CO2/ CO and SO2/Ctot ratios, and decreasing oxygen demand (see Figure 12). Unfortunately, the test period was discontinued, because of failure in the gas supply before the system had 5263

Energy Fuels 2009, 23, 5257–5268

: DOI:10.1021/ef900464j

Berguerand and Lyngfelt

The terminal velocity of the fuel particles varies over a range of 0.02-0.23 m/s, depending on their size. However, the gas velocity from the addition of 20 Ln/min sweep gas is as high as 0.76 m/s in the “hot” zone of the fuel chute, at 950 °C, and somewhat lower in the cooler zones higher up. Thus, it is the sweep gas that controls the maximum downward velocity of the fuel particles, which should be 0.23 þ 0.76 (i.e., 1 m/s). Because the “hot” zone of the fuel chute is 1 m long, this means that the fuel particles stay for at least 1 s in this zone. According to the devolatilization time above, this is long enough to fully devolatilize the fuel, meaning that only the remaining char actually reaches the bed of LOVEL. Thus, the volatiles escape the FR via the freeboard and the FR cyclone outlet and are virtually never in contact with the ilmenite. This suggests that an important part of the outgoing flow of unconverted gases from the FR could come from the volatiles. Figure 13 shows the contribution of the different gases on the oxygen demand (ΩOD) for the second test, using calculations derived from the definition of ΩOD in eq 9. Methane and hydrogen clearly are responsible for the largest contribution. A short test of 30 min was performed at TFR = 950 °C and for a fuel feed corresponding to a thermal power of 6 kW. The purpose was to analyze the transitions in the gas concentrations at the FR outlet during the first minute following the fuel feed start and stop, respectively, to quantify the volatiles in the total gas outflow. Indeed, immediately after the fuel feed start, only volatiles and nitrogen are present in the dry gases exiting the FR, because the slower char gasification is not yet significant. On the other hand, after shutting down the fuel feed, the volatiles rapidly disappear and the remaining char slowly reacts. The results are presented in Figures 14 and 15, respectively. Bag samples, later analyzed via GC, were performed to allow sufficient H2 measurements during the transitions. Those correspond to the cross symbols (þ) in Figures 14 and 15. The curves indicate a step change in the CO concentration of ∼0.7% units to be attributed to the volatiles. For H2, the corresponding step is ∼3%. Moreover, all the methane clearly comes from the volatiles (i.e. 1.1%). This suggests that most of the CO and H2 and all the CH4 in the exiting gas from the FR actually come from the volatiles, which, again, is consistent with expectations. Figure 15 also shows no variation in CO2 when the fuel feed is stopped, clearly indicating that no volatiles are oxidized to CO2. Table 5 refers to test period 1 in the second test (see Figures 10 and 11) and is used to illustrate the role of the volatiles in the oxygen demand. This test was operated at a FR temperature of 950 °C and with fluidizing conditions and fuel load similar to that in the short test previously described. Table 5 shows the following data in columns 2-7: column 2: the gas concentrations of the different species at the FR outlet given by Figure 10, as well as the total concentration of carbon-containing species column 3: the oxygen demand for each species, as well as the total oxygen demand (ΩOD) column 4: the concentrations for each species that should be related to volatiles based on Figures 14 and 15 column 5: the corresponding oxygen demands for the volatiles species (Ωvol OD), as well as the total oxygen demand associated with the volatiles

Figure 11. Second test: CO2/CO and SO2/Ctot ratios, solid fuel conversion (ηSF), oxygen demand (ΩOD), and CO2 capture (ηOO). Table 4. Third Test: Different Operating Conditions for the Five Test Periods

test period

flow at lowervelocity section, LOVEL (Ln/min)

steam velocity (cm/s)

1 2 3 4 5

30 38 20 10 30

15 19 10 5 15

stabilized fully. Nevertheless, the curve trends indicate results moving toward previously observed conditions when the steam flow was increased. In conclusion, steam flows that are too low result in poorer mixing, which leads to lower gas and solid fuel conversions. Moreover, the tests verify that the flow normally used for operation (i.e., 30 Ln/min) is sufficient and further increases in flow have no significant beneficial effect. Oxygen Demand and Fate of the Volatiles. During stable operation, the oxygen demand (ΩOD) given by eq 9 varied over a range of 27%-36%. This indicates that a rather high fraction of the reducing gases in the FR does not convert and eventually escapes the FR. These gases can either be unconverted volatiles or syngases, which are products from the char gasification. If devolatilization already occurs in the fuel feed chute, there is little chance that the volatiles will ever be in contact with oxidized ilmenite and convert. This would result in a significant part of the oxygen demand actually coming from the volatiles. In contrast, char gasification will most likely occur in the LOVEL section of the FR. In other words, to better understand the results, it is important to know the effect of volatiles and char on the oxygen demand. Estimations on the devolatilization time of different fuel particles with sizes of