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The use of petroleum coke as fuel in a 10 kWth chemical-looping combustor
international journal of greenhouse gas control 2 (2008) 169–179
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The use of petroleum coke as fuel in a 10 kWth chemical-looping combustor Nicolas Berguerand *, Anders Lyngfelt Department of Energy Conversion, Chalmers University of Technology, S-412 96 Go¨teborg, Sweden
article info
abstract
Article history:
Tests were made in a 10 kWth chemical-looping combustor with a petroleum coke as the
Received 28 August 2007
solid fuel and the oxygen carrier ilmenite, an iron titanium oxide. The fuel reactor is
Received in revised form
fluidized by steam and the oxygen carrier reacts with the volatiles released as well as
10 December 2007
the gasification intermediates CO and H2. A constant fuel flow corresponding to a thermal
Accepted 18 December 2007
power of 5.8 kW was introduced into the fuel reactor and a total of 11 h of operation was
Published on line 20 February 2008
reached. The effects of particle circulation and carbon stripper operation on solid fuel conversion, conversion of gas from the fuel reactor and CO2 capture were investigated. The
Keywords:
actual CO2 capture ranged between 60% and 75% while the solid fuel conversion was in the
Chemical-looping combustion
range of 66–78%. The low values of solid fuel conversion reflect loss of char due to low
Oxygen carrier
efficiency of the fuel reactor cyclone. The incomplete conversion of the gas from the fuel
Carbon dioxide capture
reactor is expressed as oxygen demand. The oxygen demand corresponds to the fraction of
Interconnected
oxygen lacking to achieve full gas conversion and was typically 25%, due to presence of CH4,
Fluidized beds
CO and H2 from the fuel reactor. Typical ratios of CH4, CO and H2 over the total gaseous
Solid fuel conversion
carbon from the fuel reactor are respectively 5, 10 and 25%. Low loss of non-combustible
Petroleum coke
fines from the system indicates very low attrition of the oxygen carrier. # 2007 Elsevier Ltd. All rights reserved.
1.
Introduction
Human activities are responsible for considerable atmospherical CO2 emissions, enhancing the natural greenhouse effect. The augmentation in the CO2 concentration due to fossil fuels combustion has been approximately 30% since the industrial revolution began in the 19th century (Karl and Trenberth, 2003). Numerous irregularities in local climate patterns have been observed during the past decades and those are much likely the consequence of anthropogenic greenhouse gas emissions. The accelerating frequency of warm summers in western and central Europe could be an illustration (Beniston and Diaz, 2004). Those emissions need to be lowered but since the fossil fuels are relatively abundant and the world production and consumption are so wide, it seems unlikely that the dependence on them can be reduced quickly enough. Separation and sequestration of the
produced CO2 appear to be a good option to reduce its release to the atmosphere. Chemical-looping combustion (CLC) has been introduced and developed as a new alternative to conventional combustion that prevents the CO2 from being mixed in the combustion gases. This is accomplished by preventing the air–N2 to be present in the part of the reactor system where the oxidation of the fuel takes place. A circulating metal oxide is submitted to numerous cycles in which it is oxidized in an air reactor and reduced in a fuel reactor, meanwhile providing the oxygen required for the combustion. The fuel and the metal oxide react according to the general equation: ð2n þ mÞMeO þ Cn H2m ! ð2n þ mÞMe þ mH2 O þ nCO2
(1)
where MeO denotes the oxygen carrier in the oxidized form and Me in the reduced form. The gases formed in the fuel
* Corresponding author. Tel.: +46 31 7725241; fax: +46 31 7723592. E-mail address: [email protected] (N. Berguerand). 1750-5836/$ – see front matter # 2007 Elsevier Ltd. All rights reserved. doi:10.1016/j.ijggc.2007.12.004
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particles and the gas surrounding them. Since this is generally obtained in the fluidized bed technology, two interconnected fluidized beds are chosen as reactor system, with the metal oxides circulating between the two beds. Most of the experimental work on chemical-looping combustion uses gaseous fuels, e.g. (Cho et al., 2001, 2004; Mattisson et al., 2004, 2006; Adanez et al., 2004, 2005; Johansson et al., 2004, 2006a; Johansson, 2005; Ishida et al., 1996), and the process has been demonstrated in pilots ranging from 300 W to 50 kW (Johansson et al., 2006a, 2006b; Abad et al., 2006; Lyngfelt et al., 2004; Lyngfelt and Thunman, 2005; Adanez et al., 2006; De Diego et al., 2007; Ryu et al., 2004). However, the major part of the CO2 emissions from large power plants comes from solid fuels, e.g. coal, making them an interesting option for the CLC. The reaction between the coal particles and the oxygen carrier proceeds via two steps, the fuel first needs to be devolatized/ gasified in order to produce gases that can react with the metal oxide. Investigation in laboratory reactors showed the reaction between solid fuel and oxygen carrier is feasible (Leion et al., in press, 2007). This was later confirmed by tests performed with a South African coal using a 10 kWth CLC solid fuel combustor built at Chalmers University of Technology (Berguerand and Lyngfelt, in press). For solid fuels, the volatiles released in the fuel reactor react according to reaction (1). However, the remaining char fraction needs to be gasified according to Eq. (3), after which the syngases CO and H2 can react with the metal oxide (Eq. (4)):
Nomenclature AR FR CI CLC f CO2 FC,FR FC,AR FO,AR FO,ARtot FN2 ;FR FAR,out FAR,in FC,FUEL m ˙ Fuel;in m ˙ in m ˙ loss r TAR, TFR (X)r hCC hOO hFuel t DPij FO,theor FO,act FO,meas VOD
air reactor fuel reactor circulation index (kPa Ln/min) chemical-looping combustion fraction of CO2 of total carbon leaving the FR flow of carbon containing species leaving the FR (Ln/min) flow of carbon leaving the AR (Ln/min) flow of oxygen used to oxidized the metal oxide in the AR (Ln/min) total flow of oxygen consumed in the AR (Ln/ min) flow of nitrogen fluid out of the FR (Ln/min) total gas flow out of the AR (Ln/min) total gas flow entering the AR (Ln/min) flow of gaseous carbon corresponding to the fuel flow added (Ln/min) flow of fuel entering the air reactor (g/h) flow of char entering the fuel reactor (g/h) loss of unreacted char to the air reactor (g/h) fractional conversion rate of the char (%/min) temperatures in the corresponding reactors (˚C) concentration of the species X coming from the reactor r (r can be AR or FR) carbon capture efficiency oxide oxygen fraction solid fuel conversion particle residence time in the fuel reactor (min) pressure difference measured between the pressure taps number i and j theoretical O2/C ratio for the fuel actual O2/C ratio needed under testing conditions measured O2/C ratio during the test oxygen demand
C þ H2 O=CO2 ! CO þ H2 =CO
(3)
CO=H2 þ MeO ! CO2 =H2 O þ Me
(4)
In CLC, the oxygen carriers are metal oxides that must have sufficient rates of reduction and oxidation, but also present enough strength to limit breakage and attrition. Moreover they have to be chosen so that sintering does not occur at the actual operating conditions of pressure and temperature. In previous work with gaseous fuels, oxides of the following metals have been tested: Fe, Ni, Co, Cu, and Mn. (Cho et al., 2001, 2004; Mattisson et al., 2004, 2006; Adanez et al., 2004, 2005; Johansson et al., 2004, 2006c; Johansson, 2005; Ishida et al.,
reactor are CO2 and H2O, and no air–N2 occurs in the exit stream of this reactor, which means that by condensing the H2O, it is possible to obtain almost pure CO2, while the stream coming from the air reactor is depleted in oxygen. The reduced metal oxide, Me, is then led to the air reactor where it is reoxidized by the O2 in the air:
Me þ 12O2 ! MeO
(2)
Fig. 1 shows a schematic picture of the CLC process. For the fuels and metal oxides used in chemical-looping combustion, reaction (1) is often endothermic and 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 air oxygen. As said, the advantage of this technology is to avoid the mixing of N2 from the combustion air with CO2, this mixing being the main source of difficulty, cost and energy penalty in any later separation process. The condition for the process is a good contact between the solid
Fig. 1 – Schematic picture of the CLC process. Two interconnected fluidized bed reactors, one air and one fuel reactor, with circulating oxygen carrier particles.
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1996). An important parameter which can be decisive in the choice of the oxygen carrier is its price, determinant for industrial application at large scale. Environmental and health aspects may also affect the choice. In order to increase the reactivity and durability of the particles, they are often doped with Al2O3, yttria-stabilized zirconium (YSZ), TiO2 or MgO. Their reactivity in the oxidation/reduction cycles depend on the type of oxygen carrier used, the particle size, the reduction gas used, and of course the operating temperature. In CLC with solid fuels, natural ores are considered because of their low price. A more detailed discussion on the feasibility of CLC for solid fuels is given by Berguerand and Lyngfelt (in press).
2.
Experimental
The Mexican petcoke used in the tests has the composition given in Table 1. The ash and volatile contents of petcoke are much smaller compared to the South African coal used in earlier testing on the same pilot (Berguerand and Lyngfelt, in press). Moreover, the sulphur content is almost ten times higher for the petcoke compared to the coal. This fuel is also less reactive (Leion et al., in press). The experimental setup was the same as in the tests with the South African coal and the operating conditions were unchanged except for the fuel flow which was increased to 655 g/h, which corresponds to a thermal power of approximately 5.8 kW. The temperature was 950 8C in the fuel reactor and 920–990 8C in the air reactor. Because of small size, the system is enclosed in an oven and not self-supporting in energy. A view of the fuel reactor’s different chambers and the particle circulation directions is presented in Figs. 2 and 3. Fig. 4 shows a general view over the whole reactor system: air and fuel reactors, particle filters, coal feeding, steam production unit and water seal. The reactor system is previously described by Berguerand and Lyngfelt (in press). Between the South African coal and the petcoke tests, the FR cyclone was redimensioned and rebuilt and therefore the solid fuel conversion was expected to be improved for the petcoke tests. This is detailed at the end of the Section 3.
2.1.
Figs. 2–3 – Top and front views of the fuel reactor with the particle circulation directions. No. (2) denotes the carbon stripper, (3) is the high velocity part and (4) is the low velocity part. Note: the drawings are not to scale.
Data evaluation
Below is an overview of the measured data and how these can be used to indicate the performance with respect to key issues. The concentration of CO2, noted (CO2), is given in percentage of the total volume flow in the fuel reactor’s
Table 1 – Petcoke analysis Petcoke C (wt%) H (wt%) N (wt%) S (wt%) O (wt%) H2O (wt%) Ash (wt%) Heating value (MJ/kg)
81.32 2.87 0.88 6.02 0.45 8.00 0.46 31.75
Fig. 4 – Whole pilot system: (a) air reactor; (b) riser; (c) air reactor cyclone; (d) fuel reactor.
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outlet. The value obtained is to a large extent dependent on the flow of nitrogen added, which dilutes the combustion products. The important point is that the nitrogen flow is known, giving a possibility to calculate the carbon flow. The fraction of CO2 of total carbon leaving the fuel reactor is defined by:
f CO2 ¼
ðCO2 Þ ðCO2 Þ þ ðCOÞ þ ðCH4 Þ
(5)
where (CO2), (CO) and (CH4) are measured concentrations. This gives a measure of the degree of completion of the combustion process in the fuel reactor. The ‘‘carbon flow’’ noted FC,FR is the flow of carbon containing species, i.e. CO, CO2 and CH4 leaving the fuel reactor given in Ln/min. It is calculated using the known nitrogen flow FN2 ;FR fluidizing the high velocity part, the carbon stripper, the fuel reactor particle lock, the sweep gas and half of the flows for the upper and lower particles locks. The carbon flow is given by:
FC;FR ¼ FN2 ;FR
ðCO2 Þ þ ðCOÞ þ ðCH4 Þ 1 ½ðCO2 Þ þ ðCOÞ þ ðCH4 Þ þ ðH2 Þ
(6)
This calculation assumes that the dry gases only contain N2, CO, CO2, CH4 and H2. The gas flow in the air reactor outlet is noted FAR,out and takes into account both char combustion, estimated by the (CO2)AR, and particle oxidation: 1 ðO2;exp ÞAR ¼ FAR;in 1 ðCO2 ÞAR ðO2;meas ÞAR
FAR;out
(7)
FC;FR FC;FR þ FC;AR
(9)
The oxygen demand VOD is the fraction of oxygen lacking to achieve a complete combustion of the carbon containing gases leaving the fuel reactor. It is calculated as:
VOD ¼
0:5ðCOÞ þ 0:5ðH2 Þ þ 2ðCH4 Þ ðCO2 Þ þ ðCOÞ þ ðCH4 Þ
TAR þ 273 273
(11)
Although the actual circulation is not known, this is a qualitative measure of the particle circulation in the reactor system. The ‘‘oxygen flow’’ noted FO,AR is the flow of oxygen used for oxidizing the metal oxide in the air reactor and is expressed in Ln/min. It is calculated taking into consideration the total air and nitrogen flow entering the air reactor, the expected oxygen concentration if no oxidation occurs, and the measured oxygen and CO2 concentrations in the gas leaving the air reactor. FO,ARtot is the total flow of oxygen consumed in the air reactor, accounting for particle oxidation and char combustion. The CO2 concentration gives the part of oxygen that is not used to oxidize the metal particles but instead to burn carbon. (O2,exp)AR takes into account the minor dilution of the combustion air with the particle locks nitrogen.
¼ FO;ARtot FC;AR
(12)
The ‘‘oxide oxygen fraction’’ noted hOO is another way of evaluating the carbon capture efficiency. This number is defined by the amount of oxygen used for oxidizing the oxide FO,AR in the air reactor divided by the sum of that used for oxidizing both char FC,AR (Eq. (6)) and oxide FO,AR in the air reactor.
hOO ¼ ¼
FO;ARtot FC;AR FO;AR ¼ FO;ARtot FO;AR þ FC;AR ðO2;exp ÞAR ðO2;meas ÞAR ðCO2 ÞAR ðO2;exp ÞAR ðO2;meas ÞAR ðO2;exp ÞAR ðCO2 ÞAR
(13)
(8)
the carbon capture efficiency is then defined by:
hCC ¼
CI ¼ DP810 FAR;out
FO;AR ¼ FAR;in ðO2;exp ÞAR FAR;out ððO2;meas ÞAR þ ðCO2 ÞAR Þ
where FAR,in is the flow of fluidizing gas at the inlet of the air reactor, ðO2;exp ÞAR is the (O2) expected due to the dilution by nitrogen from the particle locks, i.e. if no particle oxidation and char burning occurs, ðO2;meas ÞAR is the (O2) measured at the outlet, and (CO2)AR is the (CO2) in the air reactor outlet due to coal burning. The carbon capture efficiency hCC is the ratio of the carbon containing gas flow leaving the fuel reactor FC,FR to the total carbon containing gas flow leaving the system. The remaining part comes from the air reactor in form of CO2 and is noted: FC,AR. FC;AR ¼ FAR;out ðCO2 ÞAR
carbon on the fuel. In fact, for the pet coke used, the O2/C ratio is approximately 1.1 and thus the oxygen demand is somewhat overestimated here. The circulation index CI is defined by the pressure drop measured between pressure taps located at the risers entrance and outlet multiplied by the actual gas flow in the air reactor’s outlet, i.e. corrected for oxygen consumed and measured temperature in the air reactor:
(10)
The definition used here is a simplification, since it assumes that the oxygen needed is 1 mole of O2 per mole of
Note that both these flows are derived from the gas concentrations measured in the air reactor outlet, and any error is thus only associated with the gas concentration measurements. This makes the hOO more reliable than the carbon capture efficiency hCC defined previously. It is also possible to derive the oxide oxygen fraction from the carbon capture efficiency. First the O2/C ratio, FO,theor, is defined as moles of oxygen needed to burn the fuel completely per mole of carbon in the fuel. Then the ‘‘actual O2/C ratio’’, FO,act, is defined as actual oxygen needed under the testing conditions, i.e. corrected for the oxygen demand in the gas from the fuel reactor. This gives:
FO;act ¼ FO;theor ð1 hCC VOD Þ
(14)
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For the fuel used FO,theor is estimated to 1.1. Depending on the values for VOD and hCC, different values for VO,act can be obtained. This can be compared to the measured O2/C ratio for the tests:
FO;meas ¼
FO;ARtot FC;AR þ FC;FR
(15)
The closer to one the ratio between FO,meas and FO,act, the better the fulfilment of the mass balance in a test. Typical values for VOD and hCC are 25% and 65%, which means that FO,act is around 0.92. Then FO,meas should have the same value if the mass balance is fulfilled. Combining Eqs. (9), (13) and (15), it can be shown that the values for hCC and hOO should be equal for FO,meas = 1. The fractional loss of particle material noted Lf is the ratio between the loss of fines per hour, i.e. particles with a diameter below 45 mm, m ˙ f , and the total solid inventory in the reactor system, minv.:
Lf ¼
2.2.
m ˙f minv:
(16)
Mass balances
Mass balances of oxygen and carbon are included in the evaluation and are evaluated as follows: The mass balance of oxygen involves the total amount of oxygen consumed in the air reactor, i.e. FO,ARtot, given in Eq. (12), and the total amount of oxygen used for oxidizing the fuel. The latter can be derived from the total flow of gaseous carbon leaving the air and fuel reactors, i.e. FC,AR + FC,FR, given by Eqs. (6) and (8). The ratio showing the expected number of moles of oxygen needed per mole of carbon, taking into account the unconverted carbon containing gases at the outlet of the fuel reactor, is FO,act, and is given by Eq. (14). If the mass balance is fulfilled, the ratio between FO,ARtot, and FC,AR + FC,FR, i.e. FO,meas measured during the test (Eq. (15)), should then be
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equal to FO,act. Thus, comparing FO,meas and FO,act shows whether the mass balance of oxygen is fulfilled. The mass balance of carbon includes on the one hand carbon added with fuel, on the other the total gaseous carbon leaving the system, i.e. FC,AR + FC,FR, as well as the carbon leaving the process in the form of unconverted char. The mass balance is made in two steps. First, the ratio between FC,AR + FC,FR and FC,FUEL is calculated. Here, FC,FUEL is the flow of gaseous carbon corresponding to the fuel flow added. The ratio between the two expresses the solid fuel conversion, hFuel, given by:
hFuel ¼
FC;AR þ FC;FR FC;FUEL
(17)
This can then be compared to the solid fuel conversion derived from the amount of char collected in the water seal in relation to the total amount of fuel added during a given test period.
3.
Results and discussion
Two tests using the pet coke were performed for a total of 11 h stable conditions. The fuel feeding was not an issue in contrast to previous tests with South African coal (Berguerand and Lyngfelt, in press).
3.1.
First test period
The first test period conducted gave 4.5 h stable operation time. Fig. 5 represents the gas concentrations at the outlet of the air and the fuel reactors. Note that the CO2 concentration (fuel reactor) and the other gas concentrations have different axis scales. The curve noted ‘‘20.5-CO2(AR)’’ indicates how much oxygen is left after oxidation of char in the air reactor, while the curve ‘‘O2(AR)’’ shows how much oxygen is left after both char combustion and ilmenite oxidation, i.e. the actually
Fig. 5 – Gas concentrations for the first petcoke run in (%) and on a dry basis. (Note: the CO2 concentration has a different scale.)
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Fig. 6 – CO2/CO ratio, carbon capture and oxygen demand.
measured oxygen concentration. The bigger the gap between those curves and the closer to 20.5% the upper curve is the higher are the carbon separation and the amount of oxide actually oxidized. The value of 20.5% is (O2,exp)AR, see above. The CO2 concentration from the fuel reactor reaches a stable value of approximately 16–17% within 15 min from the fuel feed start and decreases slowly to 15% the next 2.5 h. The methane curve reaches and keeps a constant value of 1% during the test period which indicates that the fuel feed is stable since the methane measured is mostly associated with fresh petcoke entering the system. The CO concentration stabilizes after approximately 20 min at a level of 2%, double the CH4 concentration. After 175 min, the nitrogen bottles were accidentally emptied and the carbon stripper, the high velocity part and the particle locks were fluidized with air. This is very clear in the curves from Fig. 5 above: the CO2 concentration is above 25% while the methane concentration decreases to 0. Note that meanwhile the low velocity parts of the system were still fluidized with steam. All concentration levels approach earlier values after the nitrogen bottles were replaced. During this test, the hydrogen concentration from the fuel reactor was not measured but estimated to approximately 4.5%. This value was taken from the results obtained for the second petcoke test where the ratio between the hydrogen and the carbon monoxide was stable at 2.5, see below. Fig. 6 shows respectively the CO2/CO ratio, the CO2 capture and two oxygen demand curves taking, or not taking, into account the estimated H2 concentration. As expected, during the incident where parts of the fuel reactor and the particle locks were fluidized with air, the CO2/CO ratio increased rapidly, the oxygen demand decreased and the CO2 capture increased. Within a few minutes, it goes from approximately 75% to over 85%. During stable conditions, the oxygen demand, with H2 assumed, averages at 27% while the CO2 capture averages at 75%. The petcoke mass analysis gives a value for FO,Theor of 1.1. Using Eq. (14): FO;act ¼ FO;theor ð1 hCC VOD Þ ¼ 1:1ð1 0:75 0:25Þ ¼ 0:88
Eq. (15) then indicates that at steady state conditions and if the mass balance is fulfilled, the total amount of oxygen consumed in the air reactor should be in a ratio 0.88 of the total amount of gaseous carbon leaving the reactor system. Fig. 7 compares the two corresponding curves. From average flow values during stable conditions, here from one hour after the fuel start to two hours after the start, the measured O2/C ratio for the test can be obtained: FO;meas ¼
FO;ARtot 10:6 ¼ 0:66 ¼ FC;AR þ FC;FR 16:06
This is to be compared with 0.88. Thus, the mass balance was not fulfilled. With the mass balance of the fuel reactor, it is clear that the average residence time cannot exceed about 10 min (Berguerand and Lyngfelt, in press). Thus, it cannot be any doubt that the system has been working under steadystate conditions for approximately 2 h before the incident with loss of nitrogen occurred. The most likely explanation to the error in the mass balance is that the nitrogen flow used to determine the carbon flow from the fuel reactor is incorrect. Indeed it is believed that nitrogen added to the fuel chute can leak through the fuel bin. Note the lower carbon capture efficiency of 60% compared to the previous 90–95% for the South African coal tests, which was expected since dealing with a less reactive fuel.
3.2.
Second test period
The second petcoke run gave 6.5 h stable operation. The experimental conditions were similar to the first test but this time parameter variations and gas chromatography measurements were made. This gave H2 concentration values and verified that the CO2, CO and CH4 concentrations matched the ones measured by the on-line gas analysers. Fig. 8 represents those concentrations. The crosses stand for the corresponding gas chromatography measurements. A total of five different tests were realized during this run and the effects of the carbon stripper and the particle circulation on the carbon capture were investigated. In test
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Fig. 7 – Flow of oxygen consumed in the air reactor (lower curve) and flow of carbon leaving the whole system (upper curve) (in Ln/min).
number 1, the fluidizing flow in the air reactor is 110 Ln/min and the carbon stripper flow is 4 Ln/min, which correspond to 0.2 m/s. Note that in test number 1, the circulation was clearly insufficient, leading to gradually reduced gas conversions. In test 2, the circulation is increased by augmenting the air flow to 120 Ln/min, the carbon stripper flow being unchanged. For test 3, the air flow is increased to 130 Ln/min while the carbon stripper is set to 10 Ln/min, i.e. 0.5 m/s. For test 4, the carbon stripper is unchanged but the circulation is again increased by augmenting the air flow to 145 Ln/min. Finally for test 5, the circulation is unchanged but the carbon stripper flow is increased to 14 Ln/min (0.7 m/s). Those tests are delimited by the small triangles above the xaxis in Fig. 8. The fuel feed was very stable, as can be seen from the constant methane concentration, and the hydrogen concentration decreased with improved operating conditions to values below 4%.
Fig. 9 represents the oxygen concentration from the air reactor, lower curve. The upper curve shows what the oxygen concentration would have been if only the combustion of char in the air reactor were considered. The difference between this curve and the oxygen actually measured indicates the amount of oxygen used to regenerate the oxygen carrier. As described earlier, the quality of the carbon capture can be estimated by considering the gap between the upper curve and a horizontal curve valuing 20.5%, i.e. (O2,exp)AR. A small gap indicates a good capture since the part of oxygen used to burn char in the air reactor is then small. The lower the lower curve and the higher the upper one, the more oxygen is used to oxidize ilmenite compared to oxidizing char, i.e. the higher is the CO2 capture. Fig. 10 shows the CO2/CO ratio in the fuel reactor, the CO2 capture and the oxygen demand. The oxygen demand is represented by the cross symbols and is calculated using the
Fig. 8 – Gas concentrations from the fuel reactor in (%) and on a dry basis (petcoke test 2); crosses show results from GCmeasurements, triangles show changes in operation. GC measurements of H2 are shown as + symbols.
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Fig. 9 – Oxygen concentration from the air reactor for different tests condition. Upper curve indicates the reduction in oxygen concentration due to combustion of char.
Fig. 10 – CO2/CO ratio in the fuel reactor, CO2 capture and oxygen demand (from GC measurements including H2, cross symbols). The CO2 capture and the oxygen demand have the same scale on the right of the figure whereas the CO2/CO ratio has the scale on the left here as the CO2/CO ratio has the scale on the left.
GC measurements of the gas concentrations shown in Fig. 8. The five different tests are delimited by the small triangles above the x-axis. During test 1, the CO2/CO ratio decreases from 8.5 to around 2, showing the decrease in gas conversion. The circulation is insufficient and not enough oxidized ilmenite enters the fuel reactor. Increasing the circulation, test 2, increases this ratio to 7. Further increase in test 3 and 4 lead to increased ratios stabilizing at approximately 9. Increased circulation also involves decrease of the oxygen demand: from around 70% for test 1 with insufficient circulation, to a stable 25% for the latest tests with the highest circulation. The carbon capture is discussed below. Table 2 summarizes the results obtained for cases 2–5. CI stands for the circulation index and is a measure of the particle circulation. The oxygen demand takes into account the H2 measurements represented in Fig. 8.
It is clear from Table 2 and Figs. 9 and 10 that the changes in circulation and carbon stripper flow have an effect on the carbon capture. As expected, a higher fluidizing velocity in the carbon stripper increases the capture, which is clearly seen if
Table 2 – summary of the results for the second petcoke test Case CI Carbon stripper (m/s) VOD (%) hCC (%) Solid fuel conv. (%) (FO,meas/FO,act)
2
3
4
50 0.2 32.6 70.4 75 0.79
68 0.5 25.5 74.6 70 0.82
123 0.5 25.6 68.7 78 0.99
5 116 0.7 23.6 70.1 66 0.90
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Table 3 – Comparison of the fuel particles PSD before and after the fuel reactor cyclone modifications (collected in the water seal) PSD (mm and wt%) 1 > 250 180 < 1 < 250 125 < 1 < 180 90 < 1 < 125 45 < 1 < 90 1 < 45
Test 6a
Petcoke 1
1.2 4.4 34 31.8 13.2 15.4
0.3 1.9 4.1 7.8 66.2 19.7
a Test 6 is a test with South African coal (Berguerand and Lyngfelt, in press).
cases 2 and 3 are compared. Note that the simultaneous increase in circulation should reduce the CO2 capture efficiency. This is actually the case, as seen if cases 3 and 4 are compared, Figs. 9 and 10. The ratio FO,meas/FO,act gives an indication about the fulfilment of the mass balance for the tests: the closer to 1 this ratio is the better. It appears that test number 4 with a rather high circulation and high carbon stripper velocity was the best, with a ratio of 0.99. The solid fuel conversion averaged 72% for the different cases. This reflects the loss of unreacted char and the low value is explained by the poor separation efficiency of the fuel reactor cyclone. Though, the conversion was higher than the average value obtained for the earlier coal tests. This is despite the lower reactivity of petcoke, which would be expected to lead to significantly lower fuel conversion. This can be explained by the modifications made on the FR cyclone. Table 3 compares the particle size distributions (PSD) of the fuel particles collected in the water seal for the sixth test with South African coal, see (Berguerand and Lyngfelt, in press), and for test 1 with the petroleum coke. Between those tests, the cyclone design was modified. Before the rebuilding, only 30% of the collected particles had a diameter below 90 mm, indicating very insufficient separation in the cyclone. After the rebuilding, 86% had a diameter below 90 mm. The separation was clearly better even if it is still very far from the efficiency of large, commercial cyclones.
Table 4 – PSD and particles ‘‘lifetimes’’ of the ilmenite particles for the two petcoke tests Diameters (mm)
1 > 250 180 < 1 < 250 125 < 1 < 180 90 < 1 < 125 45 < 1 < 90 1 < 45 Total loss (g) Fuel operation (h) Lf, loss of finesa (%/h) Estimated lifetime (h), 1/Lf
Petcoke test 1st (%)
2nd (%)
0.8 14.2 64.7 13.8 3.8 2.7 710 4.5 0.03 3000
0 21.4 65.9 8.2 3.7 0.8 1130 6.5 0.01 9500
a Lf is calculated with a solid fuel inventory in the reactor system estimated to 13 kg.
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The oxygen demand not counting H2 is similar to the one for the previous coal tests: 17–18%. The CO2 capture ranged from 68 to 75%. This is definitely lower compared to the previous tests with coal, which is a consequence of the lower reactivity of petcoke.
3.3.
Elutriated particles
For both petcoke tests, PSD of the ilmenite collected in the air reactor filter were obtained. The results are shown in Table 4. The fines are particles with a diameter under 45 mm. The total mass of particles in the reactor system minv is estimated to 13 kg. Using Table 4, it is possible to calculate m ˙ f for each test. Then Eq. (16) gives the estimated fractional loss and particle ‘‘lifetimes’’. Fractional losses of 0.03 and 0.01%/h were obtained for respectively the first and the second tests, corresponding to lifetimes of 3000–9500 h, respectively.
3.4.
Simple model
Below two simplified reactor model concepts are considered for the fuel reactor, i.e.: - the continuous stirred-tank reactor (CSTR); - the plug flow reactor (PFR), i.e. a reactor where the fuel and the oxygen carrier particles are in co-current flow. Furthermore, it is assumed that the solid fuel reacts at a rate which is proportional to the mass, which is also a simplification. It is also assumed that there is no internal char separation in the fuel reactor and no char separation of the particle flow leaving for the air reactor. The reactor calculations here are used to calculate the loss from the fuel reactor. During initial devolatilization the solid fuel is converted into char. The loss is here defined as the char lost to the air reactor over char added to fuel reactor (in the form of fuel), on a carbon basis. Thus, if for instance the fuel has no volatiles and a loss of 10% to the air reactor, this means that 10% of the carbon added will be oxidized to CO2 in the air reactor. This means that the efficiency of CO2 capture will be 90%. If the loss is 10%, but the fraction of volatiles is 50% (on a carbon basis), the CO2 capture will be 95%. The loss from the fuel reactor can be expressed as a function of rt, where r is the fractional conversion rate of the solids and t is the residence time. The loss is obtained as
1 m ˙ loss ¼ rt þ 1 m ˙ in
m ˙ loss ¼ ert m ˙ in
for the CSTR
for the PFR
(18)
(19)
For the particles tested the fractional conversion rate in laboratory varies in the range 10–45%/minute with 100% steam (Leion et al., unpublished results). Here, 20%/minute is assumed. Assuming 5 kg oxide particles in the fuel reactor
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low velocity part, see (Berguerand and Lyngfelt, in press), a maximum conversion of 3% on mass basis and a fuel power of 5.8 kW, the oxide would be completely reduced within 6 min. Taking into account incomplete fuel and gas conversion as well as incomplete CO2 capture, the reduction time would not exceed 10 min. If the residence time in this part of the fuel reactor was longer than 10 min, a gradual decrease in the fuel oxidation efficiency would be observed. This is not the case. The residence time is then less than 10 min. With a residence time estimated as 5–10 min, the rt could be in the range 1–2, which would give mass losses in the range 33–50% for the CSTR case (Eq. (18)) and 13.5–36.8% for the PFR (Eq. (19)). The actual mass loss is in the range 25–31%, Table 2. However, the fuel conversion is 70% and the volatiles content of the fuel is 10%. The actual mass loss of char converted is then 29–36%. Since the actual fuel reactor most likely behaves like a mixture of the two cases, the loss agrees reasonably well.
3.5.
Discussion on the key issues for solid fuel CLC
The following points are discussed in detail in (Berguerand and Lyngfelt, in press) presenting the results obtained for the same unit using South African coal as the fuel. Fuel conversion: the conversion of the fuel in the reactor system is mostly dependent on the residence time of the fuel particles in the fuel reactor and the separation efficiency of the fuel reactor cyclone which recirculates unburnt particles. The rather low conversion obtained, typically 70%, can in part be attributed to the low reactivity of the fuel but is mainly due to poor efficiency of the cyclone. A large-scale efficient cyclone can of course be expected to give much better performance. Carbon capture: the carbon capture is dependent on the loss of unconverted carbon to the air reactor. It gives an indication about the efficiency of the separation in the carbon stripper. Test 2 clearly indicated the role of the carbon stripper in the carbon capture. The values obtained are typically 60–75% for the tests with this fuel. This is rather low but can be explained by the low reactivity of the petroleum coke making the residence time in the fuel reactor too short to achieve high conversion of the fuel. Thus, higher solids inventory is needed for this fuel. Redimensioning of the carbon stripper is also very likely to improve the carbon capture. Gas conversion in the fuel reactor: the degree of gas conversion shows if the circulation between the air reactor and the fuel reactor is sufficient to maintain the reactions in the fuel reactor. However, even with sufficient circulation, unconverted gases are present in the stream from the fuel reactor. Indeed CH4, CO and H2 are responsible for an oxygen demand of about 25%. The reason is inadequate contact between reducing gases released from particles and oxygen carrier. The reactor design could be changed to improve the contact between the oxygen carrier and the fuel particles. Other choice of metal oxide can also be considered. A possibility to address the reducing gases in the exhaust from the fuel reactor is to add an ‘‘oxygen polishing’’ step downstream: after the fuel reactor cyclone, pure oxygen is
injected to the gas flow to fully oxidize combustibles to CO2 and H2O.
4.
Conclusion
A chemical-looping combustor was run for a total of 11 h under stable tests conditions using pet coke as the fuel. The CO2 capture ranged from 60 to 75% which is lower than for the previous tests with a South African coal (Lyngfelt et al., 2004) due to a less reactive fuel. Thus, the residence time in the fuel reactor was not sufficient to reach high CO2 capture for this fuel. The carbon stripper effect on the CO2 capture was verified. The oxygen demand, with H2 included, was around 25%. The oxygen demand can be addressed for instance by a downstream ‘‘oxygen polishing’’ step, or a better fuel reactor design providing a better contact between the gas released from the fuel and the oxygen carrier particles. The solid fuel conversion reached higher figures of 66 to 78% thanks to an improved FR cyclone but still this means a high loss of unconverted char because of poor efficiency of the small cyclone. Particle size distribution of ilmenite particles recovered from the air reactor gives estimated fractional losses of fines from 0.01 to 0.03%/h. The low tendency for attrition/fragmentation of this material and its low market price make it an interesting option for use in solid fuel CLC.
Acknowledgement This work has been made in the EU financed research project Enhanced Capture of CO2 (ENCAP), SES6-2004-502666.
references
Abad, A., et al., 2006. Chemical-looping combustion in a 300 W continuously operating reactor system using a manganesebased oxygen carrier. Fuel 85 (9), 1174–1185. Adanez, J., et al., 2004. Selection of oxygen carriers for chemical-looping combustion. Energy Fuel. 18 (2), 371–377. Adanez, J., Garcia-Labiano, F., de Diego, L.F., Gayan, P., Abad, A., Celaya, J., 2005. Development of oxygen carriers for chemical-looping combustion. Carbon Dioxide Capture for Storage in Deep Geological Formations – Results from the CO2 Capture Project, vol. 1, pp. 587, 604. Adanez, J., Gayan, P., Celaya, J., de Diego, L.F., Garcia-Lubiano, F., Abad, A., 2006. Chemical-looping combustion in a 10 kWth prototype using a CuO/Al2O3 oxygen carrier: effect of operating conditions on methane combustion. Indus. Eng. Chem. Res. 45, 6075–6080. Beniston, M., Diaz, H.F., 2004. The 2003 heat wave as an example of summers in a greenhouse climate? Observations and Climate Modes Simulations for Basel Switzerland,. Global Planet. Change 44 (1–4), 73–81. Berguerand, N., Lyngfelt, A. Design and running of a 10 kWth Chemical-Looping Combustor for Solid Fuels. Fuel, in press.
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Cho, P., Mattisson, T., Lyngfelt, A., 2001. Reactivity of iron oxide with methane in a laboratory fluidized bed – application of chemical-looping combustion. In: 7th Int. Conf. Circulating Fluidized Beds, Niagara Falls, Ontario, May 5–7, 2002, pp. 599–606. Cho, P., Mattisson, T., Lyngfelt, A., 2004. Comparison of iron-, nickel-, copper- and manganese-based oxygen carriers for chemical-looping combustion. Fuel 83 (9), 1215–1225. De Diego, L.F., Garcia-Lubiano, F., Gayan, P., Celaya, J., Palacios, J.M., Adanez, J., 2007. Operation of a 10 kWth chemicallooping combustor during 200 h with a CuO/Al2O3 oxygen carrier. Fuel 86, 1036–1045. Ishida, M., Jin, H., Okamoto, T., 1996. A fundamental study of a new kind of medium material for chemical-looping combustion. Energy & Fuels 10 (4), 958–963. Johansson, M., 2005. Selection of oxygen carriers for chemicallooping combustion using methane as fuel. In: Environmental Inorganic Chemistry, Chalmers University of Technology, Go¨teborg. Johansson, M., Mattisson, T., Lyngfelt, A., 2004. Investigation of Fe2O3 with MgAl2O4 for chemical-looping combustion. Indus. Eng. Chem. Res. 43 (22), 6978–6987. Johansson, M., Mattisson, T., Ryde´n, M., Lyngfelt, A., 2006a. Carbon Capture via Chemical-Looping Combustion and Reforming, in International Seminar on Carbon Sequestration and Climate Change. Rio de Janeiro, Brazil. Johansson, E., Mattisson, T., Lyngfelt, A., Thunman, H., 2006b. Combustion of syngas and natural gas in a 300 W chemicallooping combustor. Chem. Eng. Res. Des. 84 (9), 819–827. Johansson, E., et al., 2006c. A 300 W laboratory reactor system for chemical-looping combustion with particle circulation. Fuel 85 (10–11), 1428–1438.
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Karl, T.R., Trenberth, K.E., 2003. Modern global climate change. Science 302, 1719–1723. Leion, H., Mattisson, T., Lyngfelt, A., Solid fuels in chemicallooping combustion. Int. J. Greenhouse Gas Control, in press Leion, H., Mattisson, T., Lyngfelt, A., 2007. The use of petroleum coke as fuel in chemical looping combustion. Fuel 86, 1947– 1958. Lyngfelt, A., Thunman, H., 2005. Construction and 100 h of operational experience of a 10 kW chemical-looping combustor. In: Thomas, D. (Ed.), 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. Elsevier Science, London, Chapter 36, pp. 625–646. Lyngfelt, A., et al., 2004. The grace project the development of oxygen carrier particles for chemical-looping combustion. Design and operation of a 10 kW chemical-looping combustor. In: Proceedings of the 7th Int. Conf. Greenhouse Gas Contr. Technol., Vancouver, Canada. Mattisson, T., Johansson, M., Lyngfelt, A., 2004. Multicycle reduction and oxidation of different types of iron oxide particles – application to chemical-looping combustion. Energy & Fuels 18 (3), 628–637. Mattisson, T., Johansson, M., Lyngfelt, A., 2006. The use of NiO as an oxygen carrier in chemical-looping combustion. Fuel 85, 736–747. Ryu, H.-J., et al., 2004. Continuous operation of a 50 kWth chemical-looping combustor: long-term operation with Niand Co-based oxygen carrier particle. In: The fifth ChinaKorea Joint Workshop on Clean Energy Technology, October 25–28, Qingdao University, China, pp. 221–230.
available at www.sciencedirect.com
journal homepage: www.elsevier.com/locate/ijggc
The use of petroleum coke as fuel in a 10 kWth chemical-looping combustor Nicolas Berguerand *, Anders Lyngfelt Department of Energy Conversion, Chalmers University of Technology, S-412 96 Go¨teborg, Sweden
article info
abstract
Article history:
Tests were made in a 10 kWth chemical-looping combustor with a petroleum coke as the
Received 28 August 2007
solid fuel and the oxygen carrier ilmenite, an iron titanium oxide. The fuel reactor is
Received in revised form
fluidized by steam and the oxygen carrier reacts with the volatiles released as well as
10 December 2007
the gasification intermediates CO and H2. A constant fuel flow corresponding to a thermal
Accepted 18 December 2007
power of 5.8 kW was introduced into the fuel reactor and a total of 11 h of operation was
Published on line 20 February 2008
reached. The effects of particle circulation and carbon stripper operation on solid fuel conversion, conversion of gas from the fuel reactor and CO2 capture were investigated. The
Keywords:
actual CO2 capture ranged between 60% and 75% while the solid fuel conversion was in the
Chemical-looping combustion
range of 66–78%. The low values of solid fuel conversion reflect loss of char due to low
Oxygen carrier
efficiency of the fuel reactor cyclone. The incomplete conversion of the gas from the fuel
Carbon dioxide capture
reactor is expressed as oxygen demand. The oxygen demand corresponds to the fraction of
Interconnected
oxygen lacking to achieve full gas conversion and was typically 25%, due to presence of CH4,
Fluidized beds
CO and H2 from the fuel reactor. Typical ratios of CH4, CO and H2 over the total gaseous
Solid fuel conversion
carbon from the fuel reactor are respectively 5, 10 and 25%. Low loss of non-combustible
Petroleum coke
fines from the system indicates very low attrition of the oxygen carrier. # 2007 Elsevier Ltd. All rights reserved.
1.
Introduction
Human activities are responsible for considerable atmospherical CO2 emissions, enhancing the natural greenhouse effect. The augmentation in the CO2 concentration due to fossil fuels combustion has been approximately 30% since the industrial revolution began in the 19th century (Karl and Trenberth, 2003). Numerous irregularities in local climate patterns have been observed during the past decades and those are much likely the consequence of anthropogenic greenhouse gas emissions. The accelerating frequency of warm summers in western and central Europe could be an illustration (Beniston and Diaz, 2004). Those emissions need to be lowered but since the fossil fuels are relatively abundant and the world production and consumption are so wide, it seems unlikely that the dependence on them can be reduced quickly enough. Separation and sequestration of the
produced CO2 appear to be a good option to reduce its release to the atmosphere. Chemical-looping combustion (CLC) has been introduced and developed as a new alternative to conventional combustion that prevents the CO2 from being mixed in the combustion gases. This is accomplished by preventing the air–N2 to be present in the part of the reactor system where the oxidation of the fuel takes place. A circulating metal oxide is submitted to numerous cycles in which it is oxidized in an air reactor and reduced in a fuel reactor, meanwhile providing the oxygen required for the combustion. The fuel and the metal oxide react according to the general equation: ð2n þ mÞMeO þ Cn H2m ! ð2n þ mÞMe þ mH2 O þ nCO2
(1)
where MeO denotes the oxygen carrier in the oxidized form and Me in the reduced form. The gases formed in the fuel
* Corresponding author. Tel.: +46 31 7725241; fax: +46 31 7723592. E-mail address: [email protected] (N. Berguerand). 1750-5836/$ – see front matter # 2007 Elsevier Ltd. All rights reserved. doi:10.1016/j.ijggc.2007.12.004
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international journal of greenhouse gas control 2 (2008) 169–179
particles and the gas surrounding them. Since this is generally obtained in the fluidized bed technology, two interconnected fluidized beds are chosen as reactor system, with the metal oxides circulating between the two beds. Most of the experimental work on chemical-looping combustion uses gaseous fuels, e.g. (Cho et al., 2001, 2004; Mattisson et al., 2004, 2006; Adanez et al., 2004, 2005; Johansson et al., 2004, 2006a; Johansson, 2005; Ishida et al., 1996), and the process has been demonstrated in pilots ranging from 300 W to 50 kW (Johansson et al., 2006a, 2006b; Abad et al., 2006; Lyngfelt et al., 2004; Lyngfelt and Thunman, 2005; Adanez et al., 2006; De Diego et al., 2007; Ryu et al., 2004). However, the major part of the CO2 emissions from large power plants comes from solid fuels, e.g. coal, making them an interesting option for the CLC. The reaction between the coal particles and the oxygen carrier proceeds via two steps, the fuel first needs to be devolatized/ gasified in order to produce gases that can react with the metal oxide. Investigation in laboratory reactors showed the reaction between solid fuel and oxygen carrier is feasible (Leion et al., in press, 2007). This was later confirmed by tests performed with a South African coal using a 10 kWth CLC solid fuel combustor built at Chalmers University of Technology (Berguerand and Lyngfelt, in press). For solid fuels, the volatiles released in the fuel reactor react according to reaction (1). However, the remaining char fraction needs to be gasified according to Eq. (3), after which the syngases CO and H2 can react with the metal oxide (Eq. (4)):
Nomenclature AR FR CI CLC f CO2 FC,FR FC,AR FO,AR FO,ARtot FN2 ;FR FAR,out FAR,in FC,FUEL m ˙ Fuel;in m ˙ in m ˙ loss r TAR, TFR (X)r hCC hOO hFuel t DPij FO,theor FO,act FO,meas VOD
air reactor fuel reactor circulation index (kPa Ln/min) chemical-looping combustion fraction of CO2 of total carbon leaving the FR flow of carbon containing species leaving the FR (Ln/min) flow of carbon leaving the AR (Ln/min) flow of oxygen used to oxidized the metal oxide in the AR (Ln/min) total flow of oxygen consumed in the AR (Ln/ min) flow of nitrogen fluid out of the FR (Ln/min) total gas flow out of the AR (Ln/min) total gas flow entering the AR (Ln/min) flow of gaseous carbon corresponding to the fuel flow added (Ln/min) flow of fuel entering the air reactor (g/h) flow of char entering the fuel reactor (g/h) loss of unreacted char to the air reactor (g/h) fractional conversion rate of the char (%/min) temperatures in the corresponding reactors (˚C) concentration of the species X coming from the reactor r (r can be AR or FR) carbon capture efficiency oxide oxygen fraction solid fuel conversion particle residence time in the fuel reactor (min) pressure difference measured between the pressure taps number i and j theoretical O2/C ratio for the fuel actual O2/C ratio needed under testing conditions measured O2/C ratio during the test oxygen demand
C þ H2 O=CO2 ! CO þ H2 =CO
(3)
CO=H2 þ MeO ! CO2 =H2 O þ Me
(4)
In CLC, the oxygen carriers are metal oxides that must have sufficient rates of reduction and oxidation, but also present enough strength to limit breakage and attrition. Moreover they have to be chosen so that sintering does not occur at the actual operating conditions of pressure and temperature. In previous work with gaseous fuels, oxides of the following metals have been tested: Fe, Ni, Co, Cu, and Mn. (Cho et al., 2001, 2004; Mattisson et al., 2004, 2006; Adanez et al., 2004, 2005; Johansson et al., 2004, 2006c; Johansson, 2005; Ishida et al.,
reactor are CO2 and H2O, and no air–N2 occurs in the exit stream of this reactor, which means that by condensing the H2O, it is possible to obtain almost pure CO2, while the stream coming from the air reactor is depleted in oxygen. The reduced metal oxide, Me, is then led to the air reactor where it is reoxidized by the O2 in the air:
Me þ 12O2 ! MeO
(2)
Fig. 1 shows a schematic picture of the CLC process. For the fuels and metal oxides used in chemical-looping combustion, reaction (1) is often endothermic and 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 air oxygen. As said, the advantage of this technology is to avoid the mixing of N2 from the combustion air with CO2, this mixing being the main source of difficulty, cost and energy penalty in any later separation process. The condition for the process is a good contact between the solid
Fig. 1 – Schematic picture of the CLC process. Two interconnected fluidized bed reactors, one air and one fuel reactor, with circulating oxygen carrier particles.
international journal of greenhouse gas control 2 (2008) 169–179
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1996). An important parameter which can be decisive in the choice of the oxygen carrier is its price, determinant for industrial application at large scale. Environmental and health aspects may also affect the choice. In order to increase the reactivity and durability of the particles, they are often doped with Al2O3, yttria-stabilized zirconium (YSZ), TiO2 or MgO. Their reactivity in the oxidation/reduction cycles depend on the type of oxygen carrier used, the particle size, the reduction gas used, and of course the operating temperature. In CLC with solid fuels, natural ores are considered because of their low price. A more detailed discussion on the feasibility of CLC for solid fuels is given by Berguerand and Lyngfelt (in press).
2.
Experimental
The Mexican petcoke used in the tests has the composition given in Table 1. The ash and volatile contents of petcoke are much smaller compared to the South African coal used in earlier testing on the same pilot (Berguerand and Lyngfelt, in press). Moreover, the sulphur content is almost ten times higher for the petcoke compared to the coal. This fuel is also less reactive (Leion et al., in press). The experimental setup was the same as in the tests with the South African coal and the operating conditions were unchanged except for the fuel flow which was increased to 655 g/h, which corresponds to a thermal power of approximately 5.8 kW. The temperature was 950 8C in the fuel reactor and 920–990 8C in the air reactor. Because of small size, the system is enclosed in an oven and not self-supporting in energy. A view of the fuel reactor’s different chambers and the particle circulation directions is presented in Figs. 2 and 3. Fig. 4 shows a general view over the whole reactor system: air and fuel reactors, particle filters, coal feeding, steam production unit and water seal. The reactor system is previously described by Berguerand and Lyngfelt (in press). Between the South African coal and the petcoke tests, the FR cyclone was redimensioned and rebuilt and therefore the solid fuel conversion was expected to be improved for the petcoke tests. This is detailed at the end of the Section 3.
2.1.
Figs. 2–3 – Top and front views of the fuel reactor with the particle circulation directions. No. (2) denotes the carbon stripper, (3) is the high velocity part and (4) is the low velocity part. Note: the drawings are not to scale.
Data evaluation
Below is an overview of the measured data and how these can be used to indicate the performance with respect to key issues. The concentration of CO2, noted (CO2), is given in percentage of the total volume flow in the fuel reactor’s
Table 1 – Petcoke analysis Petcoke C (wt%) H (wt%) N (wt%) S (wt%) O (wt%) H2O (wt%) Ash (wt%) Heating value (MJ/kg)
81.32 2.87 0.88 6.02 0.45 8.00 0.46 31.75
Fig. 4 – Whole pilot system: (a) air reactor; (b) riser; (c) air reactor cyclone; (d) fuel reactor.
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outlet. The value obtained is to a large extent dependent on the flow of nitrogen added, which dilutes the combustion products. The important point is that the nitrogen flow is known, giving a possibility to calculate the carbon flow. The fraction of CO2 of total carbon leaving the fuel reactor is defined by:
f CO2 ¼
ðCO2 Þ ðCO2 Þ þ ðCOÞ þ ðCH4 Þ
(5)
where (CO2), (CO) and (CH4) are measured concentrations. This gives a measure of the degree of completion of the combustion process in the fuel reactor. The ‘‘carbon flow’’ noted FC,FR is the flow of carbon containing species, i.e. CO, CO2 and CH4 leaving the fuel reactor given in Ln/min. It is calculated using the known nitrogen flow FN2 ;FR fluidizing the high velocity part, the carbon stripper, the fuel reactor particle lock, the sweep gas and half of the flows for the upper and lower particles locks. The carbon flow is given by:
FC;FR ¼ FN2 ;FR
ðCO2 Þ þ ðCOÞ þ ðCH4 Þ 1 ½ðCO2 Þ þ ðCOÞ þ ðCH4 Þ þ ðH2 Þ
(6)
This calculation assumes that the dry gases only contain N2, CO, CO2, CH4 and H2. The gas flow in the air reactor outlet is noted FAR,out and takes into account both char combustion, estimated by the (CO2)AR, and particle oxidation: 1 ðO2;exp ÞAR ¼ FAR;in 1 ðCO2 ÞAR ðO2;meas ÞAR
FAR;out
(7)
FC;FR FC;FR þ FC;AR
(9)
The oxygen demand VOD is the fraction of oxygen lacking to achieve a complete combustion of the carbon containing gases leaving the fuel reactor. It is calculated as:
VOD ¼
0:5ðCOÞ þ 0:5ðH2 Þ þ 2ðCH4 Þ ðCO2 Þ þ ðCOÞ þ ðCH4 Þ
TAR þ 273 273
(11)
Although the actual circulation is not known, this is a qualitative measure of the particle circulation in the reactor system. The ‘‘oxygen flow’’ noted FO,AR is the flow of oxygen used for oxidizing the metal oxide in the air reactor and is expressed in Ln/min. It is calculated taking into consideration the total air and nitrogen flow entering the air reactor, the expected oxygen concentration if no oxidation occurs, and the measured oxygen and CO2 concentrations in the gas leaving the air reactor. FO,ARtot is the total flow of oxygen consumed in the air reactor, accounting for particle oxidation and char combustion. The CO2 concentration gives the part of oxygen that is not used to oxidize the metal particles but instead to burn carbon. (O2,exp)AR takes into account the minor dilution of the combustion air with the particle locks nitrogen.
¼ FO;ARtot FC;AR
(12)
The ‘‘oxide oxygen fraction’’ noted hOO is another way of evaluating the carbon capture efficiency. This number is defined by the amount of oxygen used for oxidizing the oxide FO,AR in the air reactor divided by the sum of that used for oxidizing both char FC,AR (Eq. (6)) and oxide FO,AR in the air reactor.
hOO ¼ ¼
FO;ARtot FC;AR FO;AR ¼ FO;ARtot FO;AR þ FC;AR ðO2;exp ÞAR ðO2;meas ÞAR ðCO2 ÞAR ðO2;exp ÞAR ðO2;meas ÞAR ðO2;exp ÞAR ðCO2 ÞAR
(13)
(8)
the carbon capture efficiency is then defined by:
hCC ¼
CI ¼ DP810 FAR;out
FO;AR ¼ FAR;in ðO2;exp ÞAR FAR;out ððO2;meas ÞAR þ ðCO2 ÞAR Þ
where FAR,in is the flow of fluidizing gas at the inlet of the air reactor, ðO2;exp ÞAR is the (O2) expected due to the dilution by nitrogen from the particle locks, i.e. if no particle oxidation and char burning occurs, ðO2;meas ÞAR is the (O2) measured at the outlet, and (CO2)AR is the (CO2) in the air reactor outlet due to coal burning. The carbon capture efficiency hCC is the ratio of the carbon containing gas flow leaving the fuel reactor FC,FR to the total carbon containing gas flow leaving the system. The remaining part comes from the air reactor in form of CO2 and is noted: FC,AR. FC;AR ¼ FAR;out ðCO2 ÞAR
carbon on the fuel. In fact, for the pet coke used, the O2/C ratio is approximately 1.1 and thus the oxygen demand is somewhat overestimated here. The circulation index CI is defined by the pressure drop measured between pressure taps located at the risers entrance and outlet multiplied by the actual gas flow in the air reactor’s outlet, i.e. corrected for oxygen consumed and measured temperature in the air reactor:
(10)
The definition used here is a simplification, since it assumes that the oxygen needed is 1 mole of O2 per mole of
Note that both these flows are derived from the gas concentrations measured in the air reactor outlet, and any error is thus only associated with the gas concentration measurements. This makes the hOO more reliable than the carbon capture efficiency hCC defined previously. It is also possible to derive the oxide oxygen fraction from the carbon capture efficiency. First the O2/C ratio, FO,theor, is defined as moles of oxygen needed to burn the fuel completely per mole of carbon in the fuel. Then the ‘‘actual O2/C ratio’’, FO,act, is defined as actual oxygen needed under the testing conditions, i.e. corrected for the oxygen demand in the gas from the fuel reactor. This gives:
FO;act ¼ FO;theor ð1 hCC VOD Þ
(14)
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For the fuel used FO,theor is estimated to 1.1. Depending on the values for VOD and hCC, different values for VO,act can be obtained. This can be compared to the measured O2/C ratio for the tests:
FO;meas ¼
FO;ARtot FC;AR þ FC;FR
(15)
The closer to one the ratio between FO,meas and FO,act, the better the fulfilment of the mass balance in a test. Typical values for VOD and hCC are 25% and 65%, which means that FO,act is around 0.92. Then FO,meas should have the same value if the mass balance is fulfilled. Combining Eqs. (9), (13) and (15), it can be shown that the values for hCC and hOO should be equal for FO,meas = 1. The fractional loss of particle material noted Lf is the ratio between the loss of fines per hour, i.e. particles with a diameter below 45 mm, m ˙ f , and the total solid inventory in the reactor system, minv.:
Lf ¼
2.2.
m ˙f minv:
(16)
Mass balances
Mass balances of oxygen and carbon are included in the evaluation and are evaluated as follows: The mass balance of oxygen involves the total amount of oxygen consumed in the air reactor, i.e. FO,ARtot, given in Eq. (12), and the total amount of oxygen used for oxidizing the fuel. The latter can be derived from the total flow of gaseous carbon leaving the air and fuel reactors, i.e. FC,AR + FC,FR, given by Eqs. (6) and (8). The ratio showing the expected number of moles of oxygen needed per mole of carbon, taking into account the unconverted carbon containing gases at the outlet of the fuel reactor, is FO,act, and is given by Eq. (14). If the mass balance is fulfilled, the ratio between FO,ARtot, and FC,AR + FC,FR, i.e. FO,meas measured during the test (Eq. (15)), should then be
173
equal to FO,act. Thus, comparing FO,meas and FO,act shows whether the mass balance of oxygen is fulfilled. The mass balance of carbon includes on the one hand carbon added with fuel, on the other the total gaseous carbon leaving the system, i.e. FC,AR + FC,FR, as well as the carbon leaving the process in the form of unconverted char. The mass balance is made in two steps. First, the ratio between FC,AR + FC,FR and FC,FUEL is calculated. Here, FC,FUEL is the flow of gaseous carbon corresponding to the fuel flow added. The ratio between the two expresses the solid fuel conversion, hFuel, given by:
hFuel ¼
FC;AR þ FC;FR FC;FUEL
(17)
This can then be compared to the solid fuel conversion derived from the amount of char collected in the water seal in relation to the total amount of fuel added during a given test period.
3.
Results and discussion
Two tests using the pet coke were performed for a total of 11 h stable conditions. The fuel feeding was not an issue in contrast to previous tests with South African coal (Berguerand and Lyngfelt, in press).
3.1.
First test period
The first test period conducted gave 4.5 h stable operation time. Fig. 5 represents the gas concentrations at the outlet of the air and the fuel reactors. Note that the CO2 concentration (fuel reactor) and the other gas concentrations have different axis scales. The curve noted ‘‘20.5-CO2(AR)’’ indicates how much oxygen is left after oxidation of char in the air reactor, while the curve ‘‘O2(AR)’’ shows how much oxygen is left after both char combustion and ilmenite oxidation, i.e. the actually
Fig. 5 – Gas concentrations for the first petcoke run in (%) and on a dry basis. (Note: the CO2 concentration has a different scale.)
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Fig. 6 – CO2/CO ratio, carbon capture and oxygen demand.
measured oxygen concentration. The bigger the gap between those curves and the closer to 20.5% the upper curve is the higher are the carbon separation and the amount of oxide actually oxidized. The value of 20.5% is (O2,exp)AR, see above. The CO2 concentration from the fuel reactor reaches a stable value of approximately 16–17% within 15 min from the fuel feed start and decreases slowly to 15% the next 2.5 h. The methane curve reaches and keeps a constant value of 1% during the test period which indicates that the fuel feed is stable since the methane measured is mostly associated with fresh petcoke entering the system. The CO concentration stabilizes after approximately 20 min at a level of 2%, double the CH4 concentration. After 175 min, the nitrogen bottles were accidentally emptied and the carbon stripper, the high velocity part and the particle locks were fluidized with air. This is very clear in the curves from Fig. 5 above: the CO2 concentration is above 25% while the methane concentration decreases to 0. Note that meanwhile the low velocity parts of the system were still fluidized with steam. All concentration levels approach earlier values after the nitrogen bottles were replaced. During this test, the hydrogen concentration from the fuel reactor was not measured but estimated to approximately 4.5%. This value was taken from the results obtained for the second petcoke test where the ratio between the hydrogen and the carbon monoxide was stable at 2.5, see below. Fig. 6 shows respectively the CO2/CO ratio, the CO2 capture and two oxygen demand curves taking, or not taking, into account the estimated H2 concentration. As expected, during the incident where parts of the fuel reactor and the particle locks were fluidized with air, the CO2/CO ratio increased rapidly, the oxygen demand decreased and the CO2 capture increased. Within a few minutes, it goes from approximately 75% to over 85%. During stable conditions, the oxygen demand, with H2 assumed, averages at 27% while the CO2 capture averages at 75%. The petcoke mass analysis gives a value for FO,Theor of 1.1. Using Eq. (14): FO;act ¼ FO;theor ð1 hCC VOD Þ ¼ 1:1ð1 0:75 0:25Þ ¼ 0:88
Eq. (15) then indicates that at steady state conditions and if the mass balance is fulfilled, the total amount of oxygen consumed in the air reactor should be in a ratio 0.88 of the total amount of gaseous carbon leaving the reactor system. Fig. 7 compares the two corresponding curves. From average flow values during stable conditions, here from one hour after the fuel start to two hours after the start, the measured O2/C ratio for the test can be obtained: FO;meas ¼
FO;ARtot 10:6 ¼ 0:66 ¼ FC;AR þ FC;FR 16:06
This is to be compared with 0.88. Thus, the mass balance was not fulfilled. With the mass balance of the fuel reactor, it is clear that the average residence time cannot exceed about 10 min (Berguerand and Lyngfelt, in press). Thus, it cannot be any doubt that the system has been working under steadystate conditions for approximately 2 h before the incident with loss of nitrogen occurred. The most likely explanation to the error in the mass balance is that the nitrogen flow used to determine the carbon flow from the fuel reactor is incorrect. Indeed it is believed that nitrogen added to the fuel chute can leak through the fuel bin. Note the lower carbon capture efficiency of 60% compared to the previous 90–95% for the South African coal tests, which was expected since dealing with a less reactive fuel.
3.2.
Second test period
The second petcoke run gave 6.5 h stable operation. The experimental conditions were similar to the first test but this time parameter variations and gas chromatography measurements were made. This gave H2 concentration values and verified that the CO2, CO and CH4 concentrations matched the ones measured by the on-line gas analysers. Fig. 8 represents those concentrations. The crosses stand for the corresponding gas chromatography measurements. A total of five different tests were realized during this run and the effects of the carbon stripper and the particle circulation on the carbon capture were investigated. In test
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Fig. 7 – Flow of oxygen consumed in the air reactor (lower curve) and flow of carbon leaving the whole system (upper curve) (in Ln/min).
number 1, the fluidizing flow in the air reactor is 110 Ln/min and the carbon stripper flow is 4 Ln/min, which correspond to 0.2 m/s. Note that in test number 1, the circulation was clearly insufficient, leading to gradually reduced gas conversions. In test 2, the circulation is increased by augmenting the air flow to 120 Ln/min, the carbon stripper flow being unchanged. For test 3, the air flow is increased to 130 Ln/min while the carbon stripper is set to 10 Ln/min, i.e. 0.5 m/s. For test 4, the carbon stripper is unchanged but the circulation is again increased by augmenting the air flow to 145 Ln/min. Finally for test 5, the circulation is unchanged but the carbon stripper flow is increased to 14 Ln/min (0.7 m/s). Those tests are delimited by the small triangles above the xaxis in Fig. 8. The fuel feed was very stable, as can be seen from the constant methane concentration, and the hydrogen concentration decreased with improved operating conditions to values below 4%.
Fig. 9 represents the oxygen concentration from the air reactor, lower curve. The upper curve shows what the oxygen concentration would have been if only the combustion of char in the air reactor were considered. The difference between this curve and the oxygen actually measured indicates the amount of oxygen used to regenerate the oxygen carrier. As described earlier, the quality of the carbon capture can be estimated by considering the gap between the upper curve and a horizontal curve valuing 20.5%, i.e. (O2,exp)AR. A small gap indicates a good capture since the part of oxygen used to burn char in the air reactor is then small. The lower the lower curve and the higher the upper one, the more oxygen is used to oxidize ilmenite compared to oxidizing char, i.e. the higher is the CO2 capture. Fig. 10 shows the CO2/CO ratio in the fuel reactor, the CO2 capture and the oxygen demand. The oxygen demand is represented by the cross symbols and is calculated using the
Fig. 8 – Gas concentrations from the fuel reactor in (%) and on a dry basis (petcoke test 2); crosses show results from GCmeasurements, triangles show changes in operation. GC measurements of H2 are shown as + symbols.
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Fig. 9 – Oxygen concentration from the air reactor for different tests condition. Upper curve indicates the reduction in oxygen concentration due to combustion of char.
Fig. 10 – CO2/CO ratio in the fuel reactor, CO2 capture and oxygen demand (from GC measurements including H2, cross symbols). The CO2 capture and the oxygen demand have the same scale on the right of the figure whereas the CO2/CO ratio has the scale on the left here as the CO2/CO ratio has the scale on the left.
GC measurements of the gas concentrations shown in Fig. 8. The five different tests are delimited by the small triangles above the x-axis. During test 1, the CO2/CO ratio decreases from 8.5 to around 2, showing the decrease in gas conversion. The circulation is insufficient and not enough oxidized ilmenite enters the fuel reactor. Increasing the circulation, test 2, increases this ratio to 7. Further increase in test 3 and 4 lead to increased ratios stabilizing at approximately 9. Increased circulation also involves decrease of the oxygen demand: from around 70% for test 1 with insufficient circulation, to a stable 25% for the latest tests with the highest circulation. The carbon capture is discussed below. Table 2 summarizes the results obtained for cases 2–5. CI stands for the circulation index and is a measure of the particle circulation. The oxygen demand takes into account the H2 measurements represented in Fig. 8.
It is clear from Table 2 and Figs. 9 and 10 that the changes in circulation and carbon stripper flow have an effect on the carbon capture. As expected, a higher fluidizing velocity in the carbon stripper increases the capture, which is clearly seen if
Table 2 – summary of the results for the second petcoke test Case CI Carbon stripper (m/s) VOD (%) hCC (%) Solid fuel conv. (%) (FO,meas/FO,act)
2
3
4
50 0.2 32.6 70.4 75 0.79
68 0.5 25.5 74.6 70 0.82
123 0.5 25.6 68.7 78 0.99
5 116 0.7 23.6 70.1 66 0.90
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Table 3 – Comparison of the fuel particles PSD before and after the fuel reactor cyclone modifications (collected in the water seal) PSD (mm and wt%) 1 > 250 180 < 1 < 250 125 < 1 < 180 90 < 1 < 125 45 < 1 < 90 1 < 45
Test 6a
Petcoke 1
1.2 4.4 34 31.8 13.2 15.4
0.3 1.9 4.1 7.8 66.2 19.7
a Test 6 is a test with South African coal (Berguerand and Lyngfelt, in press).
cases 2 and 3 are compared. Note that the simultaneous increase in circulation should reduce the CO2 capture efficiency. This is actually the case, as seen if cases 3 and 4 are compared, Figs. 9 and 10. The ratio FO,meas/FO,act gives an indication about the fulfilment of the mass balance for the tests: the closer to 1 this ratio is the better. It appears that test number 4 with a rather high circulation and high carbon stripper velocity was the best, with a ratio of 0.99. The solid fuel conversion averaged 72% for the different cases. This reflects the loss of unreacted char and the low value is explained by the poor separation efficiency of the fuel reactor cyclone. Though, the conversion was higher than the average value obtained for the earlier coal tests. This is despite the lower reactivity of petcoke, which would be expected to lead to significantly lower fuel conversion. This can be explained by the modifications made on the FR cyclone. Table 3 compares the particle size distributions (PSD) of the fuel particles collected in the water seal for the sixth test with South African coal, see (Berguerand and Lyngfelt, in press), and for test 1 with the petroleum coke. Between those tests, the cyclone design was modified. Before the rebuilding, only 30% of the collected particles had a diameter below 90 mm, indicating very insufficient separation in the cyclone. After the rebuilding, 86% had a diameter below 90 mm. The separation was clearly better even if it is still very far from the efficiency of large, commercial cyclones.
Table 4 – PSD and particles ‘‘lifetimes’’ of the ilmenite particles for the two petcoke tests Diameters (mm)
1 > 250 180 < 1 < 250 125 < 1 < 180 90 < 1 < 125 45 < 1 < 90 1 < 45 Total loss (g) Fuel operation (h) Lf, loss of finesa (%/h) Estimated lifetime (h), 1/Lf
Petcoke test 1st (%)
2nd (%)
0.8 14.2 64.7 13.8 3.8 2.7 710 4.5 0.03 3000
0 21.4 65.9 8.2 3.7 0.8 1130 6.5 0.01 9500
a Lf is calculated with a solid fuel inventory in the reactor system estimated to 13 kg.
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The oxygen demand not counting H2 is similar to the one for the previous coal tests: 17–18%. The CO2 capture ranged from 68 to 75%. This is definitely lower compared to the previous tests with coal, which is a consequence of the lower reactivity of petcoke.
3.3.
Elutriated particles
For both petcoke tests, PSD of the ilmenite collected in the air reactor filter were obtained. The results are shown in Table 4. The fines are particles with a diameter under 45 mm. The total mass of particles in the reactor system minv is estimated to 13 kg. Using Table 4, it is possible to calculate m ˙ f for each test. Then Eq. (16) gives the estimated fractional loss and particle ‘‘lifetimes’’. Fractional losses of 0.03 and 0.01%/h were obtained for respectively the first and the second tests, corresponding to lifetimes of 3000–9500 h, respectively.
3.4.
Simple model
Below two simplified reactor model concepts are considered for the fuel reactor, i.e.: - the continuous stirred-tank reactor (CSTR); - the plug flow reactor (PFR), i.e. a reactor where the fuel and the oxygen carrier particles are in co-current flow. Furthermore, it is assumed that the solid fuel reacts at a rate which is proportional to the mass, which is also a simplification. It is also assumed that there is no internal char separation in the fuel reactor and no char separation of the particle flow leaving for the air reactor. The reactor calculations here are used to calculate the loss from the fuel reactor. During initial devolatilization the solid fuel is converted into char. The loss is here defined as the char lost to the air reactor over char added to fuel reactor (in the form of fuel), on a carbon basis. Thus, if for instance the fuel has no volatiles and a loss of 10% to the air reactor, this means that 10% of the carbon added will be oxidized to CO2 in the air reactor. This means that the efficiency of CO2 capture will be 90%. If the loss is 10%, but the fraction of volatiles is 50% (on a carbon basis), the CO2 capture will be 95%. The loss from the fuel reactor can be expressed as a function of rt, where r is the fractional conversion rate of the solids and t is the residence time. The loss is obtained as
1 m ˙ loss ¼ rt þ 1 m ˙ in
m ˙ loss ¼ ert m ˙ in
for the CSTR
for the PFR
(18)
(19)
For the particles tested the fractional conversion rate in laboratory varies in the range 10–45%/minute with 100% steam (Leion et al., unpublished results). Here, 20%/minute is assumed. Assuming 5 kg oxide particles in the fuel reactor
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low velocity part, see (Berguerand and Lyngfelt, in press), a maximum conversion of 3% on mass basis and a fuel power of 5.8 kW, the oxide would be completely reduced within 6 min. Taking into account incomplete fuel and gas conversion as well as incomplete CO2 capture, the reduction time would not exceed 10 min. If the residence time in this part of the fuel reactor was longer than 10 min, a gradual decrease in the fuel oxidation efficiency would be observed. This is not the case. The residence time is then less than 10 min. With a residence time estimated as 5–10 min, the rt could be in the range 1–2, which would give mass losses in the range 33–50% for the CSTR case (Eq. (18)) and 13.5–36.8% for the PFR (Eq. (19)). The actual mass loss is in the range 25–31%, Table 2. However, the fuel conversion is 70% and the volatiles content of the fuel is 10%. The actual mass loss of char converted is then 29–36%. Since the actual fuel reactor most likely behaves like a mixture of the two cases, the loss agrees reasonably well.
3.5.
Discussion on the key issues for solid fuel CLC
The following points are discussed in detail in (Berguerand and Lyngfelt, in press) presenting the results obtained for the same unit using South African coal as the fuel. Fuel conversion: the conversion of the fuel in the reactor system is mostly dependent on the residence time of the fuel particles in the fuel reactor and the separation efficiency of the fuel reactor cyclone which recirculates unburnt particles. The rather low conversion obtained, typically 70%, can in part be attributed to the low reactivity of the fuel but is mainly due to poor efficiency of the cyclone. A large-scale efficient cyclone can of course be expected to give much better performance. Carbon capture: the carbon capture is dependent on the loss of unconverted carbon to the air reactor. It gives an indication about the efficiency of the separation in the carbon stripper. Test 2 clearly indicated the role of the carbon stripper in the carbon capture. The values obtained are typically 60–75% for the tests with this fuel. This is rather low but can be explained by the low reactivity of the petroleum coke making the residence time in the fuel reactor too short to achieve high conversion of the fuel. Thus, higher solids inventory is needed for this fuel. Redimensioning of the carbon stripper is also very likely to improve the carbon capture. Gas conversion in the fuel reactor: the degree of gas conversion shows if the circulation between the air reactor and the fuel reactor is sufficient to maintain the reactions in the fuel reactor. However, even with sufficient circulation, unconverted gases are present in the stream from the fuel reactor. Indeed CH4, CO and H2 are responsible for an oxygen demand of about 25%. The reason is inadequate contact between reducing gases released from particles and oxygen carrier. The reactor design could be changed to improve the contact between the oxygen carrier and the fuel particles. Other choice of metal oxide can also be considered. A possibility to address the reducing gases in the exhaust from the fuel reactor is to add an ‘‘oxygen polishing’’ step downstream: after the fuel reactor cyclone, pure oxygen is
injected to the gas flow to fully oxidize combustibles to CO2 and H2O.
4.
Conclusion
A chemical-looping combustor was run for a total of 11 h under stable tests conditions using pet coke as the fuel. The CO2 capture ranged from 60 to 75% which is lower than for the previous tests with a South African coal (Lyngfelt et al., 2004) due to a less reactive fuel. Thus, the residence time in the fuel reactor was not sufficient to reach high CO2 capture for this fuel. The carbon stripper effect on the CO2 capture was verified. The oxygen demand, with H2 included, was around 25%. The oxygen demand can be addressed for instance by a downstream ‘‘oxygen polishing’’ step, or a better fuel reactor design providing a better contact between the gas released from the fuel and the oxygen carrier particles. The solid fuel conversion reached higher figures of 66 to 78% thanks to an improved FR cyclone but still this means a high loss of unconverted char because of poor efficiency of the small cyclone. Particle size distribution of ilmenite particles recovered from the air reactor gives estimated fractional losses of fines from 0.01 to 0.03%/h. The low tendency for attrition/fragmentation of this material and its low market price make it an interesting option for use in solid fuel CLC.
Acknowledgement This work has been made in the EU financed research project Enhanced Capture of CO2 (ENCAP), SES6-2004-502666.
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