IECM Technical Documentation: Chemical Looping Combustion for Pre

IECM Technical Documentation:

Chemical Looping Combustion

for Pre-combustion CO2 Capture

September 2012

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IECM Technical Documentation:

Chemical Looping Combustion for Pre-combustion CO2 Capture

RES Activity No. 0004000.2.672.241.003:

The Role of Simulation and Modeling in Accelerating CO2 Capture Technology

Prepared for:

National Energy Technology Laboratory

Department of Energy

Pittsburgh, PA 15236

www.netl.doe.gov

Prepared by:

Hari C. Mantripragada

Edward S. Rubin

Carnegie Mellon University

Pittsburgh, PA 15213

www.iecm-online.com

September 2012

Integrated Environmental Control Model - Technical Documentation Table of Contents v

Table of Contents

Chemical Looping Combustion 1

Objectives of this Report ........................................................................................................... 1 Introduction ............................................................................................................................... 2

Oxygen Carriers .......................................................................................................... 3 Reactor Design ............................................................................................................ 4 CLC Applications ........................................................................................................ 5

Thermodynamic Analysis of CLC Reactions .......................................................................... 10 Equilibrium Constant Solution Method ..................................................................... 10 Adiabatic Temperature .............................................................................................. 12 Air Reactor Calculations ........................................................................................... 13 Fuel Reactor Analysis................................................................................................ 19 Methods for Calculating Solids Inventory ................................................................. 30 Reactor Design .......................................................................................................... 32

IGCC power plant using CLC ................................................................................................. 34 Gas turbine and compressor calculations .................................................................. 34 CO2 Purification Unit ................................................................................................ 35 Calculation procedure ................................................................................................ 38

Case Studies ............................................................................................................................. 40 Base Case .................................................................................................................. 40 Results ....................................................................................................................... 41

Conclusion ............................................................................................................................... 48 References ............................................................................................................................... 48

Integrated Environmental Control Model - Technical Documentation List of Figures vi

List of Figures

Figure 1. Chemical looping combustion process: Oxygen carrier is oxidized in the air reactor and reduced in the fuel

reactor. .......................................................................................................................................................................... 2

Figure 2. Typical process flow diagram of CLC process with interconnected fluidized beds (1) high velocity riser air

reactor, (2) cyclone for particle separation and (3) fuel reactor (Fang et al, 2009) ...................................................... 5

Figure 3. Example of application of CLC to power generation – a simple cycle power plant where hot exhaust from air

(oxidation) and fuel (reduction) reactors are expanded in air and CO2 turbines, respectively (Brandvoll and Bolland,

2004). ............................................................................................................................................................................ 6

Figure 4. Example of a CLC combined cycle power plant. Steam is produced using the hot exhaust from air and CO2

turbines in a heat recovery steam generator (HRSG) (Naqvi and Bolland, 2007). ....................................................... 7

Figure 5. Conceptual design of combined gasification and combustion of solid fuels using chemical looping (Rizeq et al,

2003). ............................................................................................................................................................................ 8

Figure 6. ALSTOM's Ca-based CLC process (Andrus et al, 2009). ..................................................................................... 8

Figure 7. Different applications of ALSTOM's chemical-looping concept using solid fuels (Andrus, 2009) ..................... 9

Figure 8. Chemical looping with oxygen-uncoupling (Mattisson et al, 2009). .................................................................... 9

Figure 9. Schematic of chemical looping combustion ........................................................................................................ 14

Figure 10. Variation of equilibrium constant for the oxidation reaction Ni + 0.5O2 <=> NiO, with temperature ............. 14

Figure 11. Inlet and outlet flows of the air reactor.............................................................................................................. 14

Figure 12. Variation of heat of reaction of Ni + 0.5O2 NiO, with temperature ............................................................. 16

Figure 13. Variation of adiabatic reaction temperature of AR as a function of inlet OC temperature at different inlet air

temperatures. Pure NiO is assumed. x=0, z=0. ........................................................................................................... 18

Figure 14. Variation of adiabatic reaction temperature of AR as a function of inlet air temperature and inlet OC

temperatures. Pure NiO is assumed. x=0, z=0. ........................................................................................................... 18

Figure 15. Variation of adiabatic temperature of AR as a function of inlet air temperature and inlet OC temperature. 40%

NiO in the OC, x=0, z=0............................................................................................................................................. 18

Figure 16. Variation of adiabatic temperature in AR as a function of excess air ratio (x) and excess NiO ratio (z) at fixed

inlet temperature of 400oC and inlet OC temperature of 900

oC ................................................................................. 19

Figure 17. Equilibrium constants for reactions between CO and NiO and H2 and NiO as a function of temperature ....... 22

Figure 18. Variation of specific heat (Cp) of different components with temperature ....................................................... 27

Figure 19. Variation of equilibrium constant expressions with temperature ...................................................................... 27

Figure 20. Variation of equilibrium constant expressions with temperature ...................................................................... 27

Figure 21. Adiabatic FR temperature as a function of inlet fuel temperature and inlet OC temperature. z=0, pure NiO as

OC .............................................................................................................................................................................. 28

Integrated Environmental Control Model - Technical Documentation List of Figures vii

Figure 22. Adiabatic FR temperature as a function of inlet OC temperature and inlet fuel temperatures. z=0, pure NiO as

OC .............................................................................................................................................................................. 29

Figure 23. Adiabatic fuel reactor temperature as a function of excess NiO (z) at inlet OC temperature = 1200oC. xMeO =

40% ............................................................................................................................................................................. 29

Figure 24. Regression lines for solids inventory for different fuels at Xox values very close to 1. Data from Abad et al

(2007) and Mattison et al (2007) were used in estimating these. ................................................................................ 31

Figure 25. Schematic of CLC process applied to IGCC ..................................................................................................... 34

Figure 26. Flow diagram of CO2 purification unit (CPU) .................................................................................................. 35

Figure 27. Figure showing the energy requirement per tonne of CO2 cpatured as a function of the amount of inert

material (expressed in terms of N2) present in the feed gas. This figure is adopted from Aspelund and Jordal (2007).

.................................................................................................................................................................................... 37

Figure 28. Diagram showing IECM screen with performance results of a CLC reactor system ........................................ 42

Figure 29. Diagram showing IECM screen with performance results of a CLC gas turbine power block ......................... 42

Figure 30. Diagram showing IECM screen with performance results of a CLC steam cycle power block ........................ 42

Figure 31. Fuel reactor temperature as a function of air reactor temperature for excess NiO ratio at FR outlet = 0.3 ....... 43

Figure 32. Required excess air as a function of air reactor temperature for excess NiO ratio at FR outlet = 0.3 ............... 43

Figure 33. Net plant power output as a function of air reactor temperature for excess NiO ratio at FR outlet = 0.3 ......... 44

Figure 34. Net plant efficiency as a function of air reactor temperature for excess NiO ratio at FR outlet = 0.3 .............. 44

Figure 35.FR temperature as function of excess NiO ratio at FR outlet and AR temperature ............................................ 45

Figure 36. Required excess air ratio as function of excess NiO ratio at FR outlet and AR temperature ............................ 45

Figure 37. Net plant power output as function of excess NiO ratio at FR outlet and AR temperature ............................... 45

Figure 38. Net plant efficiency as function of excess NiO ratio at FR outlet and AR temperature .................................... 46

Figure 39. Mass flow rate of oxygen carrier as a function of excess NiO ratio at FR outlet and air reactor temperature .. 46

Figure 40. Total capital requirement for the CLC system as a function of air reactor temperature .................................... 47

Figure 41. Overall plant capital cost as a function of air reactor temperature .................................................................... 47

Figure 42.Total capital requirement of the CLC system as a function of excess NiO ratio at FR outlet and air reactor

temperature ................................................................................................................................................................. 47

Integrated Environmental Control Model - Technical Documentation List of Tables viii

List of Tables

Table 1. Thermodynamic properties of different species relevant for CLC applications ................................................... 13

Table 2. Equilibrium constants of different FR reactions at different temperatures ........................................................... 21

Table 3. Mass balance equations for FR ............................................................................................................................. 24

Table 4. Specific heats and other coefficients of different components used for calculation of FR adiabatic temperature 26

Table 5. Solids inventory for different fuels as a function of conversion variation. Estimated from Abad et al (2007) and

Mattisson et al (2007). ................................................................................................................................................ 30

Table 6. Design parameters of GE 7EA turbine ................................................................................................................. 34

Table 7. Input assumptions in IECM for base case ............................................................................................................ 40

Table 8.CLC Capital cost assumptions for base case ......................................................................................................... 41

Table 9. Syngas composition before and after the fuel reactor ........................................................................................... 41

Integrated Environmental Control Model - Technical Documentation Acknowledgements ix

Acknowledgements

This work is supported by the U.S. Department of Energy under Activity No.

0004000.2.672.241.003 from the National Energy Technology Laboratory (DOE/NETL). Any

opinions, findings, and conclusions or recommendations expressed in this material are those of the

authors and do not reflect the views of any agency.

We thank Haibo Zhai (Project Manager for IECM) and Karen Kietzke (Research Programmer for

IECM) of Carnegie Mellon University for their months of tireless efforts in converting the analytical

models into the user-friendly interface of IECM.

Integrated Environmental Control Model - Technical Documentation Chemical Looping Combustion 1

Chemical Looping Combustion

Objectives of this Report This report deals with the use of chemical looping combustion (CLC) technology for pre-

combustion CO2 capture from an integrated gasification combined cycle (IGCC) power plant. A

brief literature review of the CLC process is provided along with the description of the process.

In the latter part of the report, the thermodynamic principles of CLC process are explained and a

thermodynamic performance model, including reactor design, is developed which is

implemented in to the Integrated Environmental Control Model (IECM). Cost models were

developed based on data from open literature to estimate the cost of CLC system and its effect on

the overall plant costs. The models are illustrated with case studies towards the end.


Introduction Chemical looping combustion (CLC) is an indirect combustion process in which fuel is

combusted without direct contact with air. Transfer of oxygen between air and fuel takes place

with the aid of an oxygen-carrier (OC). The oxygen-carrier extracts O2 from air in one reactor

and then transfers it to fuel in a subsequent reactor. Since the fuel does not come in direct contact

with air, the products of combustion contain only carbon dioxide (CO2) and water (H2O). CO2

stream of very high purity can be obtained by condensing the water vapor. In a normal

combustion process where fuel and air react by direct contact, the flue gases contain a large

fraction of nitrogen (N2) and a lot of additional energy is required to separate the CO2 from flue

gas. The advantage of chemical looping combustion is that, no additional energy is required for

CO2 separation. Another inherent advantage of CLC is that very less NOx is formed because of

low temperature in the air reactor. A schematic of the process is shown in Figure 1.

Oxygen carriers are usually oxides of transient metals (Cu, Cd, Ni, Mn, Fe, Co). The metal

oxides supply oxygen to the fuel. The reduction reaction by which O2 is supplied to the fuel is as

follows:

(2n + m)MexOy + CnH2m nCO2 + mH2O + (2n + m)MexOy-1 (1)

The reduced form of metal oxide is sent to the air reactor where the oxidation reaction by which

oxygen is extracted from air by the O2-carrier takes place:

MexOy-1 + 1/2O2 MexOy (2)

Figure 1. Chemical looping combustion process: Oxygen carrier is oxidized in the air reactor and reduced in the fuel reactor.

The oxidation reaction in the air reactor is always exothermic, releasing a significant quantity of

heat. Thus, the high temperature products from the air reactor possess high enthalpy, which can

be used for electricity production in a gas turbine. However, since the combined overall reaction

is the same as a normal combustion reaction, CLC does not lead to any enthalpy gains.

The use of two reactor vessels and circulation of solids between the reactors are the main

operational disadvantages of CLC process. The other drawback is that the exhaust gases are at

lower temperatures and pressures than are usually witnessed in gas turbine applications.

Air

Reactor

Air (N2, O2)

N2, O2

Fuel

Reactor

Fuel

(CnH2m)

CO2, H2O

MexOy

MexOy-1

Air

Reactor

Air (N2, O2)

N2, O2

Fuel

Reactor

Fuel

(CnH2m)

CO2, H2O

MexOy

MexOy-1


Oxygen Carriers

The oxygen carrier is the most important component of a CLC system. The role of an OC is to

efficiently transfer O2 between air and fuel. Hence, it must have high reactivity with both air and

fuel. Since CLC is a cyclic process, the OC should also possess mechanical strength to sustain

for a long time. Apart from that, in order to be used in a fluidized bed reactor, OC should also

possess good fluidizing properties. So, the three main requirements of an OC are:

Reactivity – should have high reactivity with both air and fuel

Durability – should be able to sustain long cyclic oxidation and reduction processes without

fragmentation or attrition

Fluidizability – should be fluidizable and stable in cyclic processes

Besides these, the OC must also be abundantly available, environmentally benign and cost less.

Oxides of transition metals such as Cu, Cd, Ni, Mn, Fe and Co are the preferred oxygen carriers

for CLC.

Though the pure metal oxides have reactive properties, they are often alloyed with support

materials to enhance the properties. Having a porous support increases the surface area for

reaction, acts as a binder to improve the mechanical strength and improves resistance to attrition.

A few typical metal oxides and porous support materials are discussed here.

Ni-based OCs

Ni-based OCs have all the desirable characteristics for CLC mainly because they can operate at

high temperatures (> 900oC) and have good methane conversion rates. The disadvantage is that

Ni is conducive to the formation of CO and H2.

It is widely observed that NiO supported over an inert material has higher reactivity and

mechanical strength than pure NiO. Examples of support materials include Yttria-stabilized-

Zirconia (YSZ), Al2O3, NiAl2O4, MgAl2O4, oxides of Si, Ti, Zr and so on. Of these, NiO coated

with NiAl2O4 has the highest potential to be an OC for a CLC process for applications using

natural gas or syngas as the fuel because of its high reactivity in oxidation and reduction, high

selectivity of CH4 to CO2 and H2O and high mechanical strength and durability over a wide

temperature range. It is relatively inexpensive too.

Cu-based OCs

CuO/Cu has the advantage of having the highest oxygen carrying capacity among the metal

oxides. High O2-carrying capacity leads to low solids flow rate. Its reduction reaction with fuel is

also exothermic which means that no additional heat is required for the reaction to take place.

Thermodynamically also, CuO/Cu is favored for complete conversion in the reduction reaction.

Besides, Cu is one of the cheapest metals. However, reactivity of this metal oxide reduces after a

few cycles and it cannot operate at high temperatures.

In order to improve the reactivity of CuO/Cu systems, Al2O3 and SiO2 have been tried as binder

materials. Though reactivity can be improved, high-temperature operation still is an area of

concern.

Fe-based OCs


Iron oxides are abundantly available in nature in various forms – magnetite (Fe3O4), hematite

(Fe2O3) and wustite (FeO). Hence, iron based oxygen carriers have high environmental

compatibility. Because of their abundance, their prices are lower than the oxides of Cu or Ni. Of

these, Fe2O3/Fe3O4 combination has the fastest reduction rate, and hence is more favored than

Fe3O4/FeO or FeO/Fe combinations.

Though pure Fe2O3 has excellent chemical stability at lower temperatures (<800oC), it is prone to

agglomeration at higher temperatures. Addition of supports such as Al2O3 and MgAl2O4

improves reactivity and mechanical strength but there is scope for further improvement in high-

temperature operation.

However, Fe2O3 can be used as on oxygen carrier for CLC of solid fuels like coal because it

enhances gasification reactions, which need to occur before the actual combustion begins.

Mn-based

Though it is a promising metal oxide, it has not been widely tested so far. The disadvantage of

manganese oxides is their incompatibility with common support materials like Al2O3 and SiO2.

But Mn2O3 with ZrO2 as the support material has very high reactivity and mechanical strength.

Its feasibility as an oxygen carrier will be known only after wide-scale testing.

Based on the above brief description, Ni-based oxygen carriers are likely to be favored for

gaseous fuel CLC.

Mixed metal oxides

Mixed metal oxides, which are a combination of oxides of two or more metals, are novel

materials which aim to avoid the disadvantages and combine the advantages of individual metal

oxides. For example, though CuO cannot operate at high temperatures and NiO leads to the

formation of CO and H2 in the products, the combination of CuO and NiO can be operated at

high temperatures without CO-formation. Though mixed metal oxides have not been widely

tested, they do have potential to become efficient oxygen carriers for CLC.

Others

Besides the metal oxides discussed above, other oxygen carriers have also been tried for their

applicability to CLC.

Perovskites are materials in which metal oxides are combined in a non-stoichiometric ratio,

thereby introducing structural and electronic defects that can be used to enhance reactivity.

Though some of them (eg. La0.8Sr0.2Co0.2Fe0.2O3-δ) have been tested, a lot of improvement is still

needed for them to be used as oxygen carriers for CLC.

CaSO4 has also been tried as an oxygen carrier. Though it exhibits good reactivity and stability,

the formation of SO2 at high temperatures is a definite disadvantage.

Reactor Design

The reactors in which CLC reactions take place should ensure good contact between reactants

(fuel or air) and oxygen carriers so that maximum conversion is achieved. Also, since there are

two reactors – air reactor and fuel reactor – solid flow occurs between the two reactors

continuously. There should also be minimum loss of oxygen carrier in the circulation loop.

Based on these requirements, interconnected fluidized bed reactors seem to be the ideal design


configuration (Lyngefelt et al., 2001). Though there are variations in the specific design details

of the reactors, they are all based on the principle of interconnected fluidized bed reactors (see

Fig 2).

The air reactor is usually a high velocity riser and the fuel reactor is a bubbling fluidized bed

design. The oxidized solid particles leaving the riser are circulated to the fuel reactor where they

are reduced due to their reaction with the fuel. The particles in the reduced form are again

circulated to the air reactor, forming a loop. High velocity is preferred for air reactors since the

volumetric flow rate of air (because of the presence of N2) is about 10 times than that in the fuel

reactor and without high enough velocity, the size of the reactor becomes too large.

Particle separation systems in the circulation loops also form an important component in the

circulation loop. They must ensure that the maximum amount of solids is recovered from the gas

streams. Gas leakage from the loop has to be avoided. Therefore, gas seals to prevent gas leakage

also form an important component in the overall system design.

Figure 2. Typical process flow diagram of CLC process with interconnected fluidized beds (1) high velocity riser air reactor, (2)

cyclone for particle separation and (3) fuel reactor (Fang et al, 2009)

A dual circulating fluidized bed (DCFB) design, in which even the fuel reactor is operated in

turbulent regime, was used in a 120 kW experimental reactor at the Vienna University of

Technology (Kolbitsch et al, 2010). CH4, CO, H2 and C3H8 were used as fuels with different OC

materials. In preliminary experiments, Ilmenite (FeTiO3) and Ni-based carriers were found to be

promising materials for CLC with high gas conversions. To date, this is the biggest CLC system

to be experimentally tested.

CLC Applications

CLC is still in the early stages of its development. The main area of research has been towards

the improvement in oxygen carriers. Few experimental studies are available that test the

operational feasibility of CLC processes on a large scale. Hence, only theoretical literature exists

for the feasibility studies of CLC in large-scale applications.

The main application of CLC is going to be in power plants burning gaseous or solid fuels

(Ishida and Jin, 1994). An example of a natural gas power plant using the CLC concept is shown

in Fig 3. This design of power plant is basically a variation of a natural gas combined cycle

(NGCC) power plant with the gas turbine combustion chamber replaced with a chemical looping


reactor system and the gas turbine replaced by air and CO2 turbines. Pre-heated compressed air

enters the air reactor to react with the reduced form of oxygen carrier. The oxidized form of

oxygen carrier exits the air reactor to be sent to the fuel reactor. The product air, which is at high

temperature and pressure and lean in oxygen is sent for expansion in a gas turbine to generate

electricity. Fuel (natural gas) reacts with the oxidized form of oxygen carrier in the fuel reactor.

The products consisting of only CO2 and H2O are also at high temperature and they can be

expanded in another gas turbine to generate electricity (Brandvoll and Bolland, 2004).

In a variation of this power plant design, Naqvi et al (2006) introduce a heat recovery steam

generator that produces steam from the hot exhaust from the air and CO2 turbines (Fig 4). Being

a combined cycle configuration, this power plant design is more thermally efficient than a simple

cycle design without the steam turbine. The efficiency can further be improved by incorporating

multi chemical-looping reactors which will enable reheating of air in the intermediate stages of

expansion in the air turbine (Naqvi and Bolland, 2007). Though different variations exist in the

particular conceptual plant designs, the common basis is similar to a NGCC power plant with

combustion chamber replaced with a CLC reactor system.

Thermodynamic analyses of CLC combined cycles using natural gas or syngas as fuels found

that the overall efficiencies range from 48 – 55% (lower heating value). Considering that CO2

capture is inherent in this process, this efficiency range is much higher compared to that obtained

from conventional power plants with CCS. However, it must be reiterated that these are only

theoretical studies and it will take much more experimentation with the reactor designs and

oxygen carrier combinations to arrive at a definite conclusion about the feasibility of CLC for

power plant applications. Nevertheless, CLC is a promising technology for power plants with

CO2 capture.

Figure 3. Example of application of CLC to power generation – a simple cycle power plant where hot exhaust from air

(oxidation) and fuel (reduction) reactors are expanded in air and CO2 turbines, respectively (Brandvoll and Bolland, 2004).


Figure 4. Example of a CLC combined cycle power plant. Steam is produced using the hot exhaust from air and CO2 turbines in

a heat recovery steam generator (HRSG) (Naqvi and Bolland, 2007).

Syngas (CO and H2) from coal gasification can also be used as a fuel in CLC (Mattisson et al,

2007). Usually in an integrated gasification combined cycle (IGCC) power plant, oxygen from

air separation unit (ASU) is used to gasify coal, which produces syngas. Syngas can be subjected

to a water gas shift (WGS) reaction to convert CO to CO2, which is separated from H2 in a gas-

separation process like Selexol or Rectisol. The H2-rich gas is then combusted in a combined

cycle power plant to generate electricity. By directly burning syngas from gasifier in a CLC

reactor, inherent CO2 separation can be obtained, without the use of WGS and subsequent gas

separation steps. However, it is not clear if this is thermodynamically and economically more

efficient than an IGCC with CO2-capture.

The other application of CLC is to directly combust solid fuels like coal and petroleum coke

using the CLC concept, thereby avoiding the intermediate step of producing syngas in a coal

gasifier which is further combusted in a CLC reactor.

CLC can also be used for coal gasification applications – either direct gasification of coal or use

of syngas obtained from conventional coal gasification. The reaction between solid fuel and solid

oxygen carrier is usually very slow. Hence, when CLC is used for direct combustion of solid

fuels, steam is used as the fluidizing medium so that it can gasify the solids first and the gaseous

products can react with oxygen carriers (Leion et al, 2008). GE-EER put forward the concept of

a combined gasification-combustion technology (Rizeq et al, 2003). Here three circulating

fluidized beds with CO2 sorbent material (CaO) and oxygen carriers together to gasify and

combust coal (Fig 5).


Figure 5. Conceptual design of combined gasification and combustion of solid fuels using chemical looping (Rizeq et al, 2003).

ALSTOM also developed a 65 kWt pilot scale plant using a similar principle, using Ca-based

oxides and sulfides as oxygen and heat carriers as well as CO2-sorbents (Abdulally et al, 2010).

This concept is shown in Fig 6 (Andrus et al, 2009). CaS/CaSO4 acts as the oxygen carrier from

air reactor to the fuel reactor. In the fuel reactor, CaO/CaCO3 is used as a CO2-sorbent. CO2 is

recovered by heating CaCO3 in a calciner.

Figure 6. ALSTOM's Ca-based CLC process (Andrus et al, 2009).

By varying the operating parameters, different products such as syngas and hydrogen can be

obtained from this concept (Andrus, 2009). When less than stoichiometric amount of oxygen is

provided, this concept can be used to produce syngas from coal or other solids. If CaO loop is

added to this gasification process, CO and CO2 can be separated, thus producing predominantly

H2. The three options – combustion, gasification and H2-production – are schematically shown

in Fig 7.

Another method of direct combustion of solid fuels using CLC is called CLC with oxygen-

uncoupling (CLOU) where the oxygen carrier releases O2 in the fuel reactor and the gaseous O2

reacts with the solid fuel causing combustion reactions. The process and reactions are shown in

Fig 8. This avoids the intermediate step of gasification (Mattison et al, 2009).


Figure 7. Different applications of ALSTOM's chemical-looping concept using solid fuels (Andrus, 2009)

Figure 8. Chemical looping with oxygen-uncoupling (Mattisson et al, 2009).

Other applications include using CLC for hydrogen production and for natural gas reforming

(Fang et al, 2009). Though the products are different for different applications, the inherent

separation of CO2 is the common feature of all CLC applications.


Thermodynamic Analysis of CLC Reactions Thermodynamic analysis of chemical reactions evaluates the extent of conversion of reactants at

given operating conditions like the composition of reactants, temperature and pressure of the

reaction. When applied to CLC, thermodynamic analysis evaluates the maximum conversion of

fuel or air in FR and AR respectively possible by different OCs. The main objectives of

thermodynamic analysis are:

To evaluate the conversion of metal to metal oxide in the air reactor and reduction of

metal oxide in the fuel reactor at different temperatures and pressures

To evaluate the amount of OC required to maintain heat balance in the CLC system

Methodology

The basic concept of chemical equilibrium thermodynamics is the minimization of Gibbs free

energy in the reaction system. The objective is to find the product composition, given a reactant

composition and operating temperature and pressure.

To calculate the equilibrium gas composition, two different approaches can be used:

Calculation of equilibrium constant – this approach can be used if the specific reactions

are known and when the number of reactions is small

Gibbs free energy minimization – this approach is used when the likely products are

known but the exact reaction mechanism is not known. This can also be used when the

number of reaction is high, when it becomes computationally difficult to calculate

product composition from equilibrium reaction constants of each reaction

Equilibrium Constant Solution Method

Equilibrium constant for a reaction is given by

(1)

where, is the Gibbs free energy change of the reaction, calculated using the following

reaction:

(2)

where, τ = T/To, To = 298.15K. Note that the temperature scale in the above calculations is

Kelvin.

The MATLAB function used to calculate Gibbs free energy of individual species is given below:


For different reactions, Gibbs free energy change is calculated subtracting the Gibbs free energy

of formation of the reactants from that of the products. For elements, Gibbs free energy is set to

zero. Gibbs free energy of different species is calculated from their formation reactions. The

MATLAB routine used to calculate Gibbs free energy of formation is given below:

where, g_i is the Gibbs free energy of the species calculated using Eqn 2.

For a typical reaction of the form described below, equilibrium concentrations can be calculated

from the equilibrium constant and extent of reaction.

(3)

is the stoichiometric coefficient of the component Ai. The stoichiometric coefficient is

negative for reactants and positive for products.

For a system with multiple reactions, the mole fraction of a given component is given by

(4)

Subscript ‘i' denotes species (total N species) and ‘j’ denotes a particular reaction. The numerator

of the above equation is the number of moles of species ‘i' in the equilibrium mixture and

denominator is the total number of moles of all the species combined in the equilibrium mixture.

(5)

is the extent of reaction ‘j’.

G_CH4 = g_CH4 - G_C-2*G_H2; %C+2H2 -> CH4

G_CO = g_CO-G_C-0.5*G_O2; %C+0.5O2->CO

G_CO2 = g_CO2-G_C-G_O2; %C+O2->CO2

G_H2O = g_H2O-G_H2-0.5*G_O2; %H2+0.5O2->H2O

G_NiO = g_NiO-G_Ni-0.5*G_O2; %Ni+0.5O2->NiO

function G = gibbs(props,T)

%Calculate the Gibbs free energy of a species based on properties

and

%temperature.

%prop_CH4 = [-50460 -74520 1.702 9.081e-3 -2.164e-6 0];

g = props(1);

h = props(2);

A = props(3);

B = props(4);

C = props(5);

D = props(6);

G = -8.314*T*(-g/(8.314*298.15)+h/(8.314*298.15)*(1-298.15/T)+...

A*(log(T/298.15)-1+298.15/T)+...

(B/2)*298.15*(T/298.15-1)^2/(T/298.15)+ ...

(C/6)*298.15^2*(T/298.15-1)^2*(T/298.15+2)/(T/298.15) ...

+(D/2)*(T/298.15-1)^2/T^2);


Equilibrium constant of each reaction can be defined as

(6)

K can be calculated at any given temperature using eqns 1 and 2. By writing these equations for

all reactions and species, we get a series of linear equations as a function of extent of reaction

( ). These can be solved numerically and the equilibrium concentrations of each species can be

calculated.

Adiabatic Temperature

When the reaction occurs under adiabatic conditions, no heat is lost to the surroundings and

hence, the temperature of the products increases for an exothermic reaction. The process can be

thought of as occurring at standard temperature (298.15K) and then heating of the products from

standard temperature to the final temperature. Since no heat is lost, the whole heat of reaction is

used to heat up the products.

(7)

(8)

(9)

The values of constants A, B, C and D for different species are obtained from standard property

tables. Properties of some of the materials relevant to CLC are shown in Table 1.

The MATLAB function to calculate specific heat is given below:

The overall Cp of the mixture of gases is obtained by multiplying the individual specific heats

with the respective mole fractions of each species. Cp also depends on temperature, so an

iterative solution is required to calculate the adiabatic reaction temperature. Also, the heat carried

into any reactor depends on the flow rates and temperatures of different inlet components.

function C_p = spec_heat(props,T)

%Calculate the specific heat based on properties

%prop_CH4 = [-50460 -74520 1.702 9.081e-3 -2.164e-6 0];

A = props(3);

B = props(4);

C = props(5);

D = props(6);

C_p = 8.314*(A+B*298.15*(T/298.15+1)/2+ ...

C*298.15^2*((T/298.15)^2+(T/298.15)+1)/3 ...

+D/(T*298.15));


Table 1. Thermodynamic properties of different species relevant for CLC applications

Component ∆Go,f,298 (J/mol) ∆Ho,f,298 (J/mol) A B x 1e03 C x 1e06 D x 1e-05

CH4 -50,460 -74,520 1.702 9.081 -2.164

CO -137,169 -110,525 3.376 0.557 0.000 -0.031

CO2 -394,359 -393,509 5.457 1.045 0.000 -1.157

H2O -228,572 -241,818 3.470 1.450 0.000 0.121

N2 0 0 3.280 0.593 0.000 0.040

H2 0 0 3.249 0.422 0.000 0.083

C 0 0 1.771 0.771 0.000 -0.867

Cu 0 0 2.678 0.816 0.000 0.035

CuO -127,966 -155,719 5.780 0.973 0.000 -0.874

Fe 0 0 -0.111 6.111 0.000 1.150

Fe2O3 -742,697 -824,642 11.812 9.697 0.000 -1.976

Fe3O4 -1,015,779 -1,118,918 9.594 27.112 0.000 0.409

Ni 0 0 2.552 1.165 0.000 1.579

NiO -211,698 -239,858 4.802 1.488 0.000 2.633

The procedure used to calculate adiabatic temperature is as follows:

Assume an initial value of Tadiabatic

Calculate Cp,mean at that temperature

Calculate a new value for Tadiabatic using Eqn 8

Calculate Cp,mean at the new temperature and repeat the steps till the difference between

successive steps is very small

Figure 9 shows the schematic of a CLC process. Thermodynamic analysis of this system is

described in the following sections.

Air Reactor Calculations

For the case study of NiO/Ni system, the following reaction occurs in the air reactor (AR):

(10)

where, ‘x’ is the excess air ratio.

Equilibrium constant for the oxidation reaction of Ni, KAR, was calculated at different

temperatures for the oxidation reaction and the variation is depicted in Fig 10. The figure shows

that the equilibrium reaction constant decreases with increasing temperature, which is typical of

exothermic reactions, but the reaction constant is still very high.


Figure 9. Schematic of chemical looping combustion

Figure 10. Variation of equilibrium constant for the oxidation reaction Ni + 0.5O2 <=> NiO, with temperature

Figure 11. Inlet and outlet flows of the air reactor


Figure 11 shows the mass flows for the air reactor. Inlet air and the reduced form of OC

(predominantly Ni) enter the air reactor where the oxidation reaction takes place. The extent of

the reaction depends on the temperature at which it takes place. The amount of O2 at equilibrium

conditions can be calculated using the equilibrium constant in the following way

(11)

yO2 is the mole fraction of O2 at equilibrium conditions. Since Ni and NiO are solids, their

amounts are included in the equilibrium constant equation. Since the reaction constant is

inversely proportional to the mole fraction of oxygen (eqn 11), it can be safely assumed that the

oxidation reaction goes to completion, i.e almost zero amount of oxygen is present under

equilibrium conditions. Hence, the mass balance equations can be written using simple

stoichiometry for the air reactor. The following equations, expressed in molar terms for

simplicity, are used to calculate the mass balance in the AR.

Inlet flows

Mair,in = (1+x)Mair,stoichiometric (12)

MO2,in = 0.21Mair,in (13)

MN2,in = 0.79Mair,in (14)

The oxygen carrier consists of Ni, NiO and an inert material used to support the metal oxide as

well as increase the surface area of reaction. The support material considered here is Al2O3.

Oxygen carrier consisting of 40% MeO and 60% Al2O3 has been widely tested at lab-scale

[Abad et al, 2007]. This inert material does not undergo any reaction and so its mass does not

change throughout the system. The mass fraction of MeO in the OC is denoted by xMeO. Ideally,

stoichiometric amount of NiO is needed for the combustion reaction of fuel and hence no NiO is

present at the outlet of the fuel reactor and hence in the inlet to the air reactor. But extra NiO

may be present in order to carry sufficient heat between the two reactors. The excess NiO at the

AR inlet (which is FR outlet) is represented as the fraction ‘z’, which is the molar ratio of NiO to

OC.

z = MNiO/(MNiO + MNi) (15)

When ‘z’ is zero, there is no NiO in the inlet. When z = 1, which is not a possible condition,

there is no Ni in the inlet stream of the AR.

MNi,in = MNi,out,FR

MNiO,in = MNiO,out,FR = z/(1-z) MNi,in (16)

xMeO = MNiO x MWNiO/ (MNiO x MWNiO + MAl2O3 x MWAl2O3)

MAl2O3 = (1-xMeO)/xMeO x (MWNiO/MWAl2O3) x MNiO (17)

It has to be noted that xMeO is the fraction of NiO in fully oxidized OC.

Outlet flows

One mole of Ni reacts with 0.5 moles of O2 to form one mole of NiO. Hence outlet flow of O2

can be calculated as

MO2,out = MO2,in – 0.5MNi,in (18)


Nitrogen does not participate in the reactions. Hence, its inlet and outlet flows are the same.

MN2,out = MN2,in

Mair,out = MO2,out + MN2,out (19)

Since in most practical applications, there will be some amount of excess air, it can be assumed

that all of Ni is oxidized to NiO. One mole of Ni is converted to one mole of NiO.

MNi,out = 0 (20)

MNiO,out = MNiO,in + MNi,in = 1/(1-z)MNi,in (21)

Adiabatic temperature of AR

Assuming one mole of Ni at the inlet to AR, all other flows are expressed per mole of Ni. O2

mole flow is also expressed in terms of excess air ratio ‘x’ and NiO mole flow is a function of

‘z’. For CLC applications in IGCC, the most suitable approach will be to operate the CLC

system at the gas turbine pressure since depleted air from the AR will be the motive fluid in the

gas turbine. Hence, inlet air to the AR needs to be compressed to that pressure. Inlet temperature

of air to AR is the same as the compressor exit temperature. Reduced OC coming from the FR is

of the same temperature as the FR. Tair and TMeO are the temperatures of inlet air and MeO,

respectively. Thus, adiabatic temperature can be calculated as a function of ‘x’, ‘z’, Tair and

TMeO.

The heat released in the reaction is called heat of reaction. The heat of reaction for the oxidation

reaction of Ni at different temperatures is shown in Fig 12. It can be seen that the heat of reaction

decreases with temperature, but only marginally.

Figure 12. Variation of heat of reaction of Ni + 0.5O2 NiO, with temperature


The MATLAB routine to calculate the adiabatic temperature is given below:

Figure 13 shows the variation of adiabatic reaction temperature of AR as a function of inlet OC

temperature at different inlet air temperatures, when there is no excess air or NiO. Similarly, Fig

14 shows the variation with inlet air temperature at different OC temperatures. Pure NiO, without

any support material, is considered for these calculations. For no excess air or NiO, adiabatic

temperature in the AR is extremely high, much higher than the melting points of most of the

metals. Thus, there is a need for support materials in order to enhance the thermal stability of the

OCs. Excess air and MeO also help in carrying away some heat of reaction thereby reducing the

adiabatic temperature.

Figure 15 shows the same variation with a supported OC (40% NiO, 60% Al2O3). It can be

observed that the adiabatic temperature has decreased considerably compared to a pure NiO

oxygen carrier. This temperature is still higher than the melting point of NiO (1200oC).

%%%%% Calculating adiabatic reaction temperature

H_reactants = H_R_sens+H_Reaction_298;

for m=1:length(T_air)

for n=1:length(T_MeO)

num=H_reactants(m,n);

T_adia(1) = 600; %K, initial guess

for k=2:100

T = T_adia(k-1);

props = prop_Ni;

C_p_Ni = spec_heat(props,T);

props = prop_O2;

C_p_O2 = spec_heat(props,T);

props = prop_N2;

C_p_N2 = spec_heat(props,T);

props = prop_NiO;

C_p_NiO = spec_heat(props,T);

props=prop_Al2O3;

C_p_Al2O3=spec_heat(props,T);

denom = 1000*(M_Ni_out*C_p_Ni+M_O2_out*C_p_O2+...

M_N2_out*C_p_N2+M_NiO_out*C_p_NiO+M_Al2O3_out*C_p_Al2O3);

T_adia(k)=298.15-num/denom;

err = abs(T_adia(k)-T_adia(k-1));

T_out(m,n)=T_adia(k)-273.15;

if err<1, break; end

end

end

end


Figure 13. Variation of adiabatic reaction temperature of AR as a function of inlet OC temperature at different inlet air

temperatures. Pure NiO is assumed. x=0, z=0.

Figure 14. Variation of adiabatic reaction temperature of AR as a function of inlet air temperature and inlet OC temperatures.

Pure NiO is assumed. x=0, z=0.

Figure 15. Variation of adiabatic temperature of AR as a function of inlet air temperature and inlet OC temperature. 40% NiO in

the OC, x=0, z=0.


Figure 16 shows the variation of adiabatic temperature at different values of excess air and NiO

ratios, at fixed inlet air and OC temperatures. AR temperature further reduces with increasing

excess air ratio and also with increasing excess NiO ratio at the inlet to AR.

Figure 16. Variation of adiabatic temperature in AR as a function of excess air ratio (x) and excess NiO ratio (z) at fixed inlet

temperature of 400oC and inlet OC temperature of 900oC

Hence it is clear that the adiabatic AR temperature is a strong function of inlet air temperature,

inlet OC temperature, excess air ratio and also excess NiO ratio. Most of the variations are either

linear or nearly linear. Regression analysis was performed for adiabatic AR temperature as a

function of these variables and a linear fit was obtained with a fairly high R2 value. The relation

is given by:

TAR = 772 + 0.281 Tair,in + 0.660 TOC,in – 111x – 376z (R2 = 90.1%) (22)

Details of regression are given below:

This equation was adopted for use in IECM.

Fuel Reactor Analysis

Syngas as a fuel contains mainly CO and H2, but also a little CH4. The reactions of these fuels

with NiO are given by:

1. CH4 + 4NiO CO2 + 2H2O + 4Ni, ∆H = 134 kJ/mol

It is also known that when CH4 reacts with NiO, some amount of CO is also formed. This can be

attributed to the incomplete combustion of CH4, given by the reaction:

Predictor Coef SE Coef T P

Constant 771.61 13.82 55.83 0.000

T-air_2_1 0.28100 0.01136 24.73 0.000

T-MeO_2_1 0.66034 0.01136 58.11 0.000

x_2_1 -111.137 1.607 -69.16 0.000

z_2_1 -375.793 4.423 -84.96 0.000

S = 53.3004 R-Sq = 90.1% R-Sq(adj) = 90.1%


2. CH4 + 3NiO CO + 2H2O + 3Ni

3. CO + NiO CO2 + Ni, ∆H = – 43.3 kJ/mol

4. H2 + NiO H2O + Ni, ∆H = – 2.1kJ/mol

For a generic syngas which is a combination of CO, H2 and CH4, the combustion reaction can be

written as:

xCO + yH2 + zCH4 + (x+y+4z)NiO (x+z)CO2 + (y+2z)H2O + (x+y+4z)Ni (22)

When CO and H2O are present in a mixture, water gas shift reaction also takes place (Fan, 2010):

5. CO + H2O CO2 + H2

Since there are multiple reactions occurring in FR, the solution using Gibbs’ free energy

minimization can be used over the equilibrium constant method. Though solving 5 reactions is

complicated, it was found that the number of reactions can be brought down to two, which makes

the calculations much simpler. Hence, the equilibrium constant method of solution by

equilibrium constants has been used here. The method is explained below.

Equilibrium reaction constants

The equilibrium constants for different reactions were calculated over a range of temperatures, as

explained previously and the values are given in table 2. It can be seen from the table that the

equilibrium constants for the reactions of CH4 are orders of magnitude higher than for others.

For each reaction, equilibrium constant can be expressed in terms of the mole fractions as

From the expressions, the following equations can be derived:

Thus only K3 and K4 are important in a system involving syngas reacting with NiO. Further,

because K1 and K2 are exceptionally high compared to others, it can safely be assumed that yCH4

is nearly zero, meaning that there is hardly any CH4 present in the equilibrium mixture.


Table 2. Equilibrium constants of different FR reactions at different temperatures

Temperature (oC) 400 600 800 1000 1200

CH4-NiO 6.85E+06 2.48E+09 8.81E+10 9.29E+11 4.82E+12

CH4-NiO-incomplete 1.78E+03 4.45E+06 5.37E+08 1.32E+10 1.27E+11

CO-NiO 3841.12 558.05 164.08 70.48 38.00

H2-NiO 311.10 205.94 148.66 114.19 91.83

WGS 12.35 2.71 1.10 0.62 0.41

Mass balance of FR

If ε is considered the extent of a reaction, the mole flows of different species under equilibrium

conditions can be expressed as the following equations:

Mtotal = Mtotal,in + 2 ε1 + 2 ε2 (23)

MCH4 = MCH4,in – ε1 – ε2 (24)

MCO = MCO,in + ε2 – ε3 – ε5 (25)

MH2 = MH2,in – ε4 + ε5 (26)

MCO2 = MCO2,in + ε1 + ε3 + ε5 (27)

MH2O = MH2O,in + 2 ε1 + 2 ε2 + ε4 – ε5 (28)

Mole fraction of each species can be expressed as yi = Mi/Mtotal.

If yCH4 is assumed to be 0, it follows that MCH4 = 0 and hence

ε1 + ε2 = MCH4,in (29)

From equation for K3, the following expression can be derived:

ε1 + ε3 + ε5 =

(30)

From equation for K4, the following can be derived:

ε4 – ε5 =

(31)

The right hand side expressions in the above equations (30 – 31) are constants and can be called

as C2 and C3 for simplicity.

From the above equations, MCH4, MH2, MH2O and MCO2 can be obtained as follows:

MCH4 = 0 (32)

MH2 = MH2,in – C3 (33)

MH2O = MH2O,in +2MCH4,in + C3 (34)

MCO2 = MCO2,in + C2 (35)

Also, it can be derived that


MCO2 + MCO = MCO2,in + MCO,in + MCH4,in – C3

Hence, expression for MCO can also be obtained:

MCO = MCO,in + MCH4,in – C2 (36)

Mtotal = Mtotal,in + 2MCH4,in (37)

Hence mole flows of all the reacting components can be obtained by the above algebraic

expressions (32 – 37). The constants C2 and C3 are functions of equilibrium constants K3 and K4,

which are a function of temperature. Thus, mass balance can be solved for any given

temperature. The mole flows of other components in the syngas such as N2, Ar etc remain

unchanged since they don’t take part in any reaction.

It is known that natural logarithms of equilibrium constants relate linearly with 1/T, where T is

temperature expressed in K (Smith et al, 2005). K3 and K4 were plotted in this way and the graph

is shown in Fig 17, including the regression equations. From the regression equations, the

following expressions can be derived for K3 and K4, T is expressed in Kelvin:

ln(K3) = 5726/T – 0.2407 (38)

ln(K4) = 1520/T + 3.5467 (39)

Thus, the mass balance equations have been reduced to pure algebraic expressions.

Figure 17. Equilibrium constants for reactions between CO and NiO and H2 and NiO as a function of temperature

The inlet flows of solids to the FR are the same as the outlet flows of AR. Thus,

MNi,in,FR = MNi,out,AR

MNiO,in,FR = MNiO,out,AR

The quantity of Al2O3 remains unchanged.

Whatever quantity of NiO comes into the FR, only stoichiometric amount is consumed. From

Eqn 22, the stoichiometric molar flow of NiO can be expressed as:

MNiO,stoic = MCO,fuel + MH2,fuel + 4MCH4,fuel

One mole of NiO is consumed to form one mole of Ni. Thus, the outlet flow rates of solids can

be calculated as:


MNiO,out,FR = MNiO,in,FR – MNiO,stoic (40)

MNi,out,FR = MNi,in,FR + MNiO,stoic (41)

Thus, both the solids and gases’ mass flow has been accounted for.

Mass balance equations for FR are summarized in Table 3.

Adiabatic temperature of FR

Because of temperature dependence of specific heat, an iterative procedure is required to

calculate adiabatic temperature from Eqn 8. However, a simplified solution is developed here

which avoids the use of interative calculations.

Heat of reaction at 298.15K can be written as

Using expressions from the above table and substituting expressions for MNiO,stoic from Eqn

22 and C2 and C3 from eqns 30 and 31, heat of reaction can be expressed in terms of inlet molar

flows or mole fractions.

The reactants (fuel and OC) are at a higher temperature than 298.15K. For calculation, it is

assumed that the reactants are cooled to 298.15K, at which temperature the reaction takes place

and the products are heated to the adiabatic temperature. So the sensible heat of cooling of

reactants has to be accounted for.


Table 3. Mass balance equations for FR

Component Molar flow into FR Molar flow out of FR Change in molar flow

Al2O3 MAl2O3,in =

MAl2O3,in 0

Ni MNi,in (assumed to be 0) MNiO,stoic MNiO,stoic

NiO MNiO,in =

MNiO,in – MNiO,stoic - MNiO,stoic

CH4 MCH4,in MCH4,out = 0 - MCH4,in

CO MCO,in MCO,in + MCH4,in – C2 MCH4,in – C2

H2 MH2,in MH2,in – C3 - C3

CO2 MCO2,in MCO2,in + C2 C2

H2O MH2O,in MH2O,in + 2MCH4 + C3 2MCH4,in + C3

N2, O2, Ar … Mi,in Mi,in 0

Enthalpy of reactants is the sum of Hsens and heat of reaction. When divided by the molar flow

rate of fuel can be expressed as a function of mole fractions instead of molar flow rates:

Specific heat of products can also be similarly written in terms of the flow rates of fuel gas

components.


where, To = 298.15K.

From Eqn 8, adiabatic temperature is

In the above expressions, specific heats and reaction constants K3 and K4 are dependent on

temperature. Figures 18 – 20 show the variation of specific heats and the expressions 1/(1+K3),

K3/(1+K3), 1/(1+K4) and K4/(1+K4) as a function of temperature. It can be seen that specific

heats of almost all the components, except CH4, CO2 and H2O are fairly constant over the range

of temperature. Assuming an average specific heat for each component will not lead to

significant errors in calculations. The fuel reactor operating temperature is always in the range of

900 – 1200oC. In this narrow range, Cp of even CH4, CO2 and H2O can be considered constant.

Fuel temperature in IGCC cases is always constant around 50oC. If the fuel is natural gas (CH4),

it may have to be preheated to a temperature of 400oC. In this range of fuel temperatures also, Cp

of all the components can be considered to be constant. Thus, Cp for all the components except

CH4, CO2 and H2O is considered as constant over the whole range of temperatures. For CH4,

CO2 and H2O, two values are taken depending on whether they appear in the reactants or

products.

Similarly, the terms K3/(1+K3) and K4/(1+K4) are fairly constant and close to 1 in the FR

operating range of temperature. Average values of 0.9851 and 0.9913 are considered for these

two variables respectively.

Both 1/(1+K3) and 1/(1+K4) vary with temperature as shown in Fig 20. Linear regression

equations were obtained for these two variables as shown in Fig 20. These equations can be used

in the calculation of adiabatic temperature. However, because of the difference in magnitude of

terms involving formation enthalpies (Hf) of different components and 1/(1+K3) and 1/(1+K4),

the variation in temperature does not affect the overall calculation by much. So, in places where

the magnitude of these variables is negligible, average values are used but when the values are

significant, the linear regression equations are used. It was found that using average values for

these two variables (0.0149 and 0.0087 respectively) does not cause a significant error. Hence

the average values are used.

Since both Hreactants and Cp,products can now be expressed in terms of constants and inlet mole

fractions, the expression for adiabatic temperature becomes a linear relation and can be easily

solved without iterations.


The constants used for solution are given in the Table 4.

Using these constants the following expressions can be derived:

T = To – (A + 14.15yCO2T +2.42yH2OT)/B

Hence,

T = (BTo – A)/(B + 14.15yCO2 + 2.42yH2O)

Where,

A =

B =

, where bi are the coefficients for each component

shown in Table 4.

In most cases, the value of 14.15yCO2 + 2.42yH2O is negligible compared to B. Hence, the

expression for adiabatic temperature can be further simplified to:

T (oC) = To – A/B, where To =25

oC.

This procedure is used for inclusion in IECM.

Table 4. Specific heats and other coefficients of different components used for calculation of FR adiabatic temperature

Component Specific heat Cp,i (J/mol-K) Coefficient (bi)

CH4 44.86 256

CO 31.03 82.26

H2 29.55 70.91

CO2 44.44 50.37

H2O 37.11 39.02

N2 30.62 30.62

O2 32.4 32.4

Ar 20.79 20.79

Ni 31.9

NiO 55.24

Al2O3 134.76


Figure 18. Variation of specific heat (Cp) of different components with temperature

Figure 19. Variation of equilibrium constant expressions with temperature

Figure 20. Variation of equilibrium constant expressions with temperature


Alternative method to calculate adiabatic temperature of FR

The adiabatic temperature of FR can also be calculated in a manner similar to that explained for

AR. Adiabatic temperature depends on the inlet composition and temperature of fuel, OC and

their mass flow rates, by using regression analysis for TFR in terms of inlet composition and

temperatures. However, since the number of components in inlet syngas is much higher than in

air, regression equations can be developed for fixed syngas compositions.

As case studies, three different fuel compositions are considered – syngas derived from Illinois#6

coal using GE, E-Gas and Shell gasification technologies (NETL, 2007a).

When CLC is applied to an IGCC power plant, it is likely to be place downstream of the sulfur

capture unit, which operates typically at a temperature of 50oC. Thus, the fuel temperature can be

fixed at 50oC. Hence, the adiabatic reaction temperature depends on the temperature and flow

rate of incoming OC, assuming one mole of fuel. The flow rate of OC can be expressed in terms

of the variable ‘z’, the excess NiO in the incoming OC.

Inlet temperature of OC will be the same as the temperature at which AR operates because the

OC comes from the AR.

For a particular syngas composition, Fig 18 shows the variation of adiabatic FR temperature with

inlet fuel temperature, as a function of inlet OC temperature. Pure NiO is assumed as the OC.

Figure 21. Adiabatic FR temperature as a function of inlet fuel temperature and inlet OC temperature. z=0, pure NiO as OC

Similarly, adiabatic FR temperature is shown as a function of inlet OC temperature at different

inlet fuel temperatures in Fig 19. It can be seen that the temperature is a strong function of inlet

OC and fuel temperatures.


Figure 22. Adiabatic FR temperature as a function of inlet OC temperature and inlet fuel temperatures. z=0, pure NiO as OC

Figure 20 shows the variation in FR temperature with excess NiO in the inlet (z) for three

different gasification technologies. OC inlet temperature is 1200oC. The temperature is a strong

function of excess NiO for all three cases.

Figure 23. Adiabatic fuel reactor temperature as a function of excess NiO (z) at inlet OC temperature = 1200oC. xMeO = 40%

Regression analysis was performed for FR adiabatic temperature as a function of variables inlet

OC temperature and excess NiO. The results showed a fairly strong linear relation with R2 values

nearly equal to 1. The equations for different gasification cases are as follows:

TFR = 61.3 + 56z + 0.891TOC, for GE case (R2 = 99.7%) (41)

TFR = 19.4 + 95.7z + 0.891TOC, for E-Gas case (R2 = 99.6%) (42)

TFR = 102 + 8.06z + 0.9TOC, for the Shell case (R2 = 99.7%) (43)

Thus, adiabatic temperature of FR can be calculated once AR temperature and z are fixed. This

method, however, has not been used in IECM since the earlier method is more robust and can be

adapted to any inlet composition or temperatures.


Methods for Calculating Solids Inventory

Two important variables in the design of a CLC system are the solids circulation rate and solids

inventory. Solids circulation rate is determined by the amount of oxygen that needs to be

transferred between the reactors and also to cater to the heat balance of the system. The mass

balance analysis described before calculates the mass and molar flow rate of solids between the

two reactors. Two methods are presented here to calculate solids inventory in the system.

Method 1

Inventory of solids in each reactor depends on the reactivity of the OC with fuel and air, the

conversion of solids in each reactor and the times of residence in each reactor. Abad et al (2007)

describe a detailed methodology of calculating solids inventory for CH4 as fuel, given the type of

OC and its reaction kinetic parameters obtained from experimental investigation. Mattisson et al

(2007) describe a similar methodology for syngas as fuel. A parameter called ‘conversion’ is

defined as

Xox = (m – mred)/(mox – mred), at the exit of the AR, and

Xred = 1 – (m – mred)/(mox – mred), at the exit of FR and the ‘conversion difference’ as

∆X = Xox – Xred

In the above equations, ‘m’ is the mass of OC, mox is the mass of OC if it was completely

oxidized and mred is the mass of OC if it were completely in the reduced state. The two papers

cited above calculate the solids inventory per MW of each fuel at different values of ∆X and Xox.

In the present report, it is assumed that the OC coming out of AR is in completely oxidized state,

i.e, it contains only NiO and no Ni. Thus, Xox =1. The graphs for different fuels depicted in the

two studies were used to estimate the overall solids inventory at Xox values very close to 1 and

for different values of ∆X. Table 3 shows the values estimated from these papers. These values

for used to develop regression equations for overall solids’ inventory for different fuel types.

Figure 21 shows the excellent fit offered by a 3rd

order polynomial equation for all the three

fuels. These equations, applicable only when Xox is very close to 1, are used for estimation of

solids inventory.

Table 5. Solids inventory for different fuels as a function of conversion variation. Estimated from Abad et al (2007) and

Mattisson et al (2007).

Overall solids inventory mOC (kg/MWf)

For CH4 fuel For CO fuel For H2 fuel

∆X mOC ∆X mOC ∆X mOC

0.01 32 0.01 30 0.01 25

0.03 40 0.03 35 0.03 30

0.06 50 0.05 40 0.07 40

0.12 70 0.09 50 0.09 50

0.2 100 0.25 100 0.23 100

0.33 150 0.4 150 0.4 150

0.54 250 0.61 250 0.6 250


Figure 24. Regression lines for solids inventory for different fuels at Xox values very close to 1. Data from Abad et al (2007) and

Mattison et al (2007) were used in estimating these.

From the terminology used in this report, mole flow of OC at any point (excluding Al2O3) in the

loop is:

MOC = MNiO + MNi

Mass flow rate of OC is given by

m = MNiOMWNiO + MNiMWNi

Since ‘z’ is the molar fraction of NiO in the OC,

z = MNiO/MOC

Therefore,

MOC = MNi/(1-z) = MNiO/(1-z)

To get the mass flow of the theoretically fully oxidized form of OC at any point in the system,

the molar flow rate of OC can be multiplied by the molecular weight of NiO, since one mole of

Ni corresponds with one mole of NiO. Similarly, the molar flow rate of OC can be multiplied by

the molecular weight of Ni in order to get the mass flow rate of the theoretically fully reduced

form. Thus,

mox = MOCMWNiO

mred = MOCMNi

Xox is assumed to be 1. Xred can be derived as

Xred = 1 – MNiO/MOC = 1 – z (44)

Hence ∆X = z.

Thus, the regression equations of solids’ inventory for different fuels based on their lower

heating value can be expressed as follows:

mOC,CH4 = 264.05z3 – 28.904z

2 + 347.91z + 28.968 (kg/MW, LHVCH4=800kJ/mol)

mOC,CO = 298.72z3 – 38.899z

2 + 278.79z + 26.467 (kg/MW, LHVCO=283.2kJ/mol)


mOC,H2 = 704.5z3 – 451.67z

2 + 403.12z + 18.147 (kg/MW, LHVH2=241.5kJ/mol)

Overall solids inventory can be expressed as

msolids = (yCH4 mOC,CH4 + yCO mOC,CO + yH2 mOC,H2)LHVfuel (45)

This is the solids inventory for both the reactors combined. Individual solids inventories can be

calculated if the mean residence time of solids in either reactor is known. The solids inventory in

the other reactor can be calculated from the overall solids inventory.

msolids,AR = mOC,in,AR tAR = msolids – msolids,FR (46)

msolids,FR = mOC,in,FR tFR = msolids – msolids,AR (47)

Since this method uses calculations based on figures from published literature, it can prone to

cause errors. Hence, an alternative method is used for inclusion in IECM.

Method 2

Solids inventory can also be calculated if the residence time needed for complete conversion in

each reactor are known. Mean residence time for a NiO/Al2O3 system can be approximated to

be 4s for AR and 60s for FR [Wolf, 2004]. The overall solids inventory can be calculated by

adding equations 46 and 47.

Reactor Design

The most likely design for a CLC system will be comprised of circulating fluidized bed (CFB)

design for AR and FR (Fig 2). For reactor design, the methodology described in Lyngfelt et al

(2001) is used.

The first step is to calculate the area of air reactor, which is a function of the air flow rate

AAR (m2) = Mair,in x 22.4x(TAR+273.15)/(PAR x 273.15 x Ug) (48)

where, PAR is the pressure at which AR operates, Ug is the superficial gas velocity. The most

common value of Ug for these kind of applications is 7m/s, though a range of 5 – 10m/s is

possible [Fan, 2010].

Diameter of vessel D (m) = (4xA/3.14)½.

Maximum diameter of a vessel is 8m [Fan, 2010]. If D > 8m, the air flow is divided in to

multiple vessels.

Pressure drop in the AR is a function of the bed material (solids) and area of AR, which can be

calculated from Eqn 46:

∆pAR (Pa) = msolids,AR x 9.81/AAR (49)

Height of AR is a function of the pressure drop, density of oxygen carrier and void space in the

reactor:

hAR = ∆pAR / (dOC x 9.81 (1 – v)) (50)

where, dOC is the apparent density of solid particles (3446 kg/m3) [Abad et al, 2007].


Wolf (2004) uses the following equation for solids fraction in the dense bed of AR. Void space

is 1 – solids fraction.

v = 1- ε = -0.053 ln (Ug) + 0.2306

For Ug in the range of 5 – 10 m/s, this value is between 0.11 and 0.15. A value of 0.13 can be

taken as the mean solids fraction in AR.

Volume of the dense bed of AR is calculated as the product of AR height and area

VAR,dense = AAR x hAR (m3)

Volume can also be calculated directly using the equation:

VAR,dense = msolids,AR/(dOC x v) (m3)

This is the volume of dense bed. The overall height of AR riser is about 5 times the height of the

height of dense bed [Wolf, 2004]. Hence the volume of AR can be calculated as;

VAR = 5VAR,dense =5 msolids,AR/(dOC x v) (m3) (51)

Area and height of FR are calculated using similar procedure. For FR with large residence time,

Ug is very low, on the order of 0.05 m/s. From Wolf (2004), the solids fraction in FR (bubbling

bed) can be estimated as:

ε = (Ug/Ut)0.4

,where Ut is the terminal velocity. The typical ratio Ug/Ut is 0.25 and hence solids

fraction (1-v) is about 0.6.

The pressure drop in FR is calculated using this area and the bed material in FR (mOC,FR).

∆pFR (Pa) = msolids,FR x 9.81/AFR (52)

hFR = ∆pFR / (dOC x 9.81 x v)) (53)

VFR = AFR x hFR (m3)

or simply,

VFR = msolids,FR/(dOC x v) (m3) (54)


IGCC power plant using CLC Figure 22 shows a schematic of the CLC process applied to IGCC. The depleted air from the air

reactor, where the oxidation reaction is exothermic, is at a high temperature and pressure. The air

stream is expanded in a gas turbine, generating electricity. The heat from the exhaust of the gas

turbine can be recovered in a heat recovery steam generator (HRSG) to generate steam which is

further expanded in a steam turbine, generating more electricity. The outlet from the fuel reactor

can also be expanded and cooled in a similar combination of gas turbine and steam turbine.

Water is condensed from the products of FR after the HRSG and the CO2-rich stream is sent for

purification and compression.

Figure 25. Schematic of CLC process applied to IGCC

Gas turbine and compressor calculations

Thermodynamic models for gas turbine and compressor are already available in IECM. Hence,

new models are not developed. However, for the low temperature CLC applications, the gas

turbines used in the current version of IECM, which are based on GE 7FA and GE 7FB

technologies will not be suitable. GE 7EA turbine was found more acceptable to the low

temperature applications such as CLC. Design parameters of GE 7EA are given in Table # [GE

7EA factsheet].

Table 6. Design parameters of GE 7EA turbine

Parameter Value

Pressure ratio 12.7:1

Design rating (with natural gas) 85.1 MW

Turbine inlet temperature [Chaker and Meher-Homji, 2002] 1110oC

Exhaust temperature 537oC

Design flow rate 297 kg/s


Using these design parameters, gas turbine and compressor were calibrated using procedures

previously followed in IECM [IGCC tech document]. Gas turbine efficiency is varied to match

the exhaust gas temperature and compressor efficiency is varied to match the overall power

output. Following this calibration procedure, gas turbine and compressor isentropic efficiencies

are fixed at 88.7% and 82.1%, respectively.

It can be noted that in a conventional combined cycle power plant, the flow rate of air that passes

through the compressor is lower than the flow rate of gas passing through the gas turbine. In

contrast, in a CLC system, the amount of air that is compressed is more than the amount that

passes through the gas turbine. Hence, net power output from a CLC system is lower than from a

corresponding conventional gas turbine power plant.

CO2 Purification Unit

After the outlet stream from FR is expanded to atmospheric pressure and cooled in HRSG (to a

temperature of roughly 120oC), it goes through a condenser and drying unit in order to remove

all the H2O from the gas. The dried gas consists only of CO2 and other components like N2, Ar,

CO, H2 etc, which together can be called ‘inerts’. In the CO2 purification unit, inerts are

separated from the CO2-rich stream in order to get a high purity (~99%) CO2 stream to be

compressed. The process is shown in Fig 23. The mass balance and energy requirements are

described in this section.

yCO2,out,FR = mole fraction of CO2 in the gas from HRSG, which is the same as at FR outlet

yH2O,out,FR = mole fraction of H2O in the gas from HRSG

Mout,FR = molar flow of gas after HRSG

The inlet flows into CPU are:

Mgas,dry = Mout,FR – MH2O,out,FR (57)

Minert,in = Mout,FR – MH2O,out,FR – MCO2,out,FR (58)

MCO2,in = MCO2,out,FR (59)

yCO2,in = MCO2,in/(MCO2,in + Minert,in)

Minert,in =

MCO2,in (60)

Condenser

and Drying

CO2 Purification and

Compression Unit

From

HRSG

Dry gas

into CPU

(1 bar)

High purity CO2 to

pipeline

(150 bar)

H2O Purge

(inerts, CO2)

Figure 26. Flow diagram of CO2 purification unit (CPU)


Assume the molar ratio of CO2 and inerts in the purge gas is ‘a’. This value is determined by

thermodynamic calculations and is not available at this point of time. So, this needs to be

assumed from empirical data.

a = MCO2,purge/Minert,purge

Purity of CO2 in the stream to pipeline is given by

yCO2,out = MCO2,out/(MCO2,out + Minert,out)

Minert,in = Minert,out + Minert,purge

Minert,in = Minert,out + MCO2,purge/a

= Minert,out + (MCO2,in – MCO2,out)/a

This can be rearranged in terms of mole fraction of CO2:

or

(61)

If ‘a’ is known and yCO2,out is fixed, MCO2,out can be calculated.

Mass of CO2 captured is given by

mCO2,captured = 44MCO2,out

(62)

Efficiency of capture is given by:

ηCO2,capture = MCO2,out / MCO2,in =

(63)

For energy requirements, the estimation given in Aspelund and Jordal (2007) is used. Figure 24

shows a result from that paper, from which line P2 is applicable to CLC purpose. The figure

shows energy requirement to clean and compress flue gas from 1 bar to 150 bar, as a function of

the amount of inert material (in the form of Nitrogen). The mole fraction of inerts is 1 – mole

fraction of CO2.


Figure 27. Figure showing the energy requirement per tonne of CO2 cpatured as a function of the amount of inert material

(expressed in terms of N2) present in the feed gas. This figure is adopted from Aspelund and Jordal (2007).

A rough correlation is given by the following linear equation

WCPU,total (kWh/tonne CO2) = 289.25 – 183.33yCO2,in (R2 = 98.12%) (64)

This power includes purification and compression to 150 bar.

Cost Estimation

Capital cost of fluidized bed reactor for CLC is based on the cost of similar gasifiers (NETL,

2007b). The base cost is $16.15M (2000$) for a CFB capacity of 180m3. The cost of AR and FR

are estimated based on that. An exponent of 0.6 is used. Apart from the reactors, there is a set of

cyclones to separate solids from gases. Their cost can be estimated to be roughly 5% of reactor

costs.

DCAR = (no. of trains) x 16.15 x (VAR/180)0.6

($M, 2000) (65)

DCFR = (no. of trains) x 16.15 x (VFR/180)0.6

($M, 2000) (66)

DCCyclone = 0.05(DCAR + DCFR)

Oxygen carrier (NiO/Al2O3) unit cost COC = $7.75/kg

Initial OC cost = mOC x COC ($)

Degradation rate = 0.027%/hr (Mattisson et al, 2007)

Make up OC = 0.027 x mOC/100 (kg/hr)

Annual make-up OC cost = (capacity factor) x 0.0027 x (mOC/100) x COC x 8760

System Design

In an IGCC process, the limiting variable is the mass flow rate of flue gases into the gas turbine.

Once a gas turbine design is chosen, its pressure ratio and the power output at a design mass flow

of gases are fixed. For e.g. a GE 7FB design gas turbine with a pressure ratio of 18.5 has a design

output of 180MW with an air flow rate of 448kg/s. This flow rate has been chosen for

calculations.

mdesign,GT = 448 (kg/s)


Mdesign,GT = mdesign,GT/29 (kmol/s)

The air flow into the air reactor of the CLC system should be such that the outlet flow should be

equal to the design value of the gas turbine.

From Eqn 18, MO2,out = MO2,in – 0.5MNi,in

Mair,out = MO2,out + MN2,out = MO2,in – 0.5MNi,in + MN2,in = Mair,in – 0.5MNi,in

The Ni coming into AR is equal to the stoichiometric amount with respect to the fuel. The flow

rate of air at the outlet of AR should be equal to the design flow rate. Therefore,

Mair,in = Mdesign,GT + 0.5MNiO,stoic = Mdesign,GT + 0.5sNiO,fMfuel

(67)

For NiO, stoichiometric ratio with respect to fuel is half of the stoichiometric ratio of O2 in air

with respect to fuel. Thus,

Mair,in = Mdesign,GT + sO2,fMfuel (68)

The actual air can be expressed in terms of stoichiometric air and excess air ratio, as in Eqn 69.

(1+x)Mair,stoic = Mdesign,GT + sO2,fMfuel (69)

Stoichiometric air can be expressed in terms of fuel flow and stoichiometric ratio of O2 with

respect to fuel. One mole of O2 is equivalent to 4.76 moles of air.

(1+x)(4.76 sO2,f)Mfuel = Mdesign,GT + sO2,fMfuel

From the above equation, fuel flow requirement of the CLC system can be calculated for a

choice of gas turbine.

(70)

Once the required fuel flow is known, required air flow at the inlet to AR can be calculated from

Eqn 68.

Calculation procedure

The calculation procedure for the CLC system can be listed as follows:

1. Choose the gas turbine. This fixes the operating pressure (PAR), design air flow

(Mdesign,GT) and the turbine output power. This also fixes the pressure ratio of the

compressor and hence the temperature of air (Tair) going into AR, which is calculated

using the compressor model equations.

2. Fuel composition is fixed by the choice of gasifier. From the fuel composition, calculate

the stoichiometric O2 (sO2,f)and NiO (sNiO,f) ratios. Fuel temperature is fixed at 50oC

because it comes from the sulfur cleanup process which operates at 50oC.

3. Choose the air reactor temperature (TAR). Usual operating temperatures using an NiO-

based OC range from 900 – 1200 oC.

4. Choose the excess NiO ratio (z)

5. Since TAR and z are fixed, fuel reactor temperature (TFR) can be calculated from the

regression Eqns 41 – 43, depending on the type of gasifier


6. The OC temperature (TMeO) going into the AR is the same as TFR.

7. Since TMeO, Tair, TAR and z are known, excess air ratio (x) can be calculated from Eqn 22.

This denotes the excess air required to maintain the air reactor temperature.

8. Flow rate of fuel (Mfuel) can be calculated since Mdesign,GT and x are known.

9. FR calculations - based on fuel flow rate and composition, flow rates of individual

components can be obtained. Based on the thermodynamic model for FR described

before, the flow rates of different components, including solids and gases, at the outlet of

FR can be calculated.

10. AR calculations – based on the fuel flow rate and excess air ratio, inlet air flow can be

calculated. Based on the thermodynamic model for AR described before, the flow rates of

different components, including solids and gases, at the outlet of AR can be calculated.

11. Overall solids inventory can be calculated based on z and Eqn 45. Assuming a residence

time of solids in the FR which is typically around 60s [Wolf, 2010], the solids inventory

in the FR can be calculated using Eqn 46 and hence solids inventory in the AR

12. Based on the calculations above, reactor dimensions can be estimated using the procedure

described in the reactor design section

13. From the reactor volume, capital cost can be estimated using eqn#. O&M costs can be

calculated based on the solids flow rates.

14. Energy output from the gas turbine and expander and energy required for compressor can

also be calculated


Case Studies The models developed in the preceding sections were coded into IECM. This section presents a

few case studies of this model in IECM.

Base Case

The following assumptions are considered for the base case configuration:

GE (quench) gasifier

Selexol H2S control system

Chemical looping CO2 capture system

Wet cooling tower

2 GE 7EA gas turbines

Illinois# bituminous coal

Performance and operating conditions of the CLC system, which are the defaults in IECM, are

given in Table 7. All costs are calculated in 2011 constant USD. Capital cost assumptions for the

CLC system are shown in Table 8. Note that the values for project and process contingencies are

fixed at 30% and 50%, respectively, since this technology has not been tested at a commercial

scale. IECM defaults are used for other parameters.

Table 7. Input assumptions in IECM for base case

Metal Oxide (MeO) Type NiO

Inert Support Material Al2O3

MeO Fraction in Oxygen Carrier (OC) (%) 40

CO2 Product Stream

Max CO2 Compressor Capacity (tonne/hr) 299.4

No. of Operating CO2 Compressors 1

No. of Spare CO2 Compressors 0

Air Reactor

Air Reactor Temperature (oC) 1100

Inlet Excess MeO (mol MeO/mol OC) 0.3

Excess Air Ratio* 3.204

Superficial Gas Inlet Velocity (m/s) 7

Residence Time of Solids (s) 5

Fuel Reactor

Fuel Reactor Temperature (oC) 1045

Stoichiometric Ratio (mol O2/mol fuel)* 0.3842

Combustion Efficiency (%)* 99.22

Residence Time of Solids (s) 60

OC Degradation Rate (%/hr) 0.0272

Maximum Train Diameter (m) 8

*Calculated based on other inputs


Table 8.CLC Capital cost assumptions for base case

Construction Time (years) 3

General Facilities Capital (%PFC) 10

Engineering & Home Office Fees (%PFC) 7

Project Contingency Cost (%PFC) 30

Process Contingency Cost (%PFC) 50

Royalty Fees (%PFC) 0.5

Results

Performance results

The performance model calculates the mass and energy balances of the system. Figures 28 – 30

show the mass flows of different streams in the CLC system and power block for the base case.

Table 9 shows the syngas composition in and out of the fuel reactor. Almost all of CO and H2 are

converted to CO2 and H2O. As mentioned earlier, 100% conversion is not possible with Ni-based

oxygen carriers. Depleted air coming out of the air reactor has about 17% O2.

Table 9. Syngas composition before and after the fuel reactor

Syngas Components Syngas In (%vol) Syngas Out (%vol)

Carbon Monoxide (CO) 39.17 0.28

Hydrogen (H2) 35.38 0.31

Methane (CH4) 0.57 0.00

Hydrogen Sulfide (H2S) 0.01 0.01

Ammonia (NH3) 0.01 0.01

Hydrochloric Acid (HCl) 0.01 0.01

Carbon Dioxide (CO2) 12.88 51.73

Water Vapor (H2O) 10.15 45.83

Nitrogen (N2) 0.93 0.92

Argon (Ar) 0.89 0.88

Oxygen (O2) 0.00 0.00

Total 100 100


Figure 28. Diagram showing IECM screen with performance results of a CLC reactor system

Figure 29. Diagram showing IECM screen with performance results of a CLC gas turbine power block

Figure 30. Diagram showing IECM screen with performance results of a CLC steam cycle power block


Effect of air reactor temperature

Figures 31 – 34 show the effect of varying the air reactor temperature on different performance

variables, for a fixed excess NiO content of 0.3. Figure 31 shows that FR temperature increases

with AR temperature but is always less than the AR temperature. Figure 32 shows that the excess

air ratio required to meet the mass flow requirements and AR temperature decreases with an

increase in AR temperature, which means the compressor energy requirement also decreases. On

the other hand, because of higher temperatures, more output is produced from the turbines,

which is the reason for increased net plant power output, as shown in Fig 33. Net pant efficiency

also increases with AR temperature.

Figure 31. Fuel reactor temperature as a function of air reactor temperature for excess NiO ratio at FR outlet = 0.3

Figure 32. Required excess air as a function of air reactor temperature for excess NiO ratio at FR outlet = 0.3


Figure 33. Net plant power output as a function of air reactor temperature for excess NiO ratio at FR outlet = 0.3

Figure 34. Net plant efficiency as a function of air reactor temperature for excess NiO ratio at FR outlet = 0.3

Effect of excess NiO

Figures 35 – 39 show the effect of excess NiO at FR outlet (AR inlet) on different performance

variables, at different AR temperatures. Figure 35 shows that the FR temperature increases and

approaches AR temperature, with increasing excess NiO fraction. The required excess air ratio

also decreases with increasing NiO fraction (Fig 36) because the excess NiO is sufficient to

maintain the AR temperature. Figure 37 shows that the net power output does not strongly

depend on the excess NiO fraction since the power output is determined by air flow rate and

temperature. However, net plant efficiency decreases with excess NiO, as shown in Fig 38.

Figure 39 shows the variation in the mass flow rate of OC into the AR, which increases

exponentially beyond an excess NiO fraction of about 0.5.


Figure 35.FR temperature as function of excess NiO ratio at FR outlet and AR temperature

Figure 36. Required excess air ratio as function of excess NiO ratio at FR outlet and AR temperature

Figure 37. Net plant power output as function of excess NiO ratio at FR outlet and AR temperature


Figure 38. Net plant efficiency as function of excess NiO ratio at FR outlet and AR temperature

Figure 39. Mass flow rate of oxygen carrier as a function of excess NiO ratio at FR outlet and air reactor temperature

Cost results

Figure 40 shows the total capital requirement (TCR) of the CLC system (which includes air

reactor, fuel reactor, cyclones and CO2 purification unit) as a function of air reactor temperature,

for the base case excess NiO fraction. For the base case (1100oC), TCR is about $119 million.

TCR increases with increasing AR temperature. Figure 41 shows the overall plant capital cost as

a function of AR temperature. For the base case (1100oC), the overall plant capital cost is about

$3300/kW-net. Because higher AR temperature leads to higher power output and efficiency, the

specific plant capital cost decreases significantly with AR temperature. It was found that the

CLC system contributes 12-13% of the overall plant capital cost. Figure 42 shows that TCR of

the CLC system depends strongly on the excess NiO fraction.

Cost of CO2 avoided

To calculate the cost of CO2 avoided using the CLC system, reference plant CO2 flow rate is

considered to be 0.795 kg/kWh and cost of electricity was assumed to be $75.34/MWh, based on

an IGCC plant with two GE 7FA turbines. The cost of CO2 avoided using the CLC system was

found to be $34.5/tonne of CO2.


Figure 40. Total capital requirement for the CLC system as a function of air reactor temperature

Figure 41. Overall plant capital cost as a function of air reactor temperature

Figure 42.Total capital requirement of the CLC system as a function of excess NiO ratio at FR outlet and air reactor temperature


Comparison with other configurations

In the current version of IECM, a base plant without CCS using GE 7EA turbine technology

does not exist. Hence, a direct comparison of the plant with CCS (using CLC) to a plant without

CCS is not possible. However, the user may change the design parameters (pressure ratio,

temperature, isentropic efficiency) for a plant using GE 7FA/7FB turbine technologies to match

the output with a 7EA turbine, in order to get a sense of the additional performance and cost

penalties of CCS with CLC.

To compare the CLC system with a Selexol process, the IECM user should adjust the number of

turbines so that the amount of CO2 captured in both the plants is comparable.

Conclusion

Thermodynamic and cost models were developed for pre-combustion CO2 capture using

chemical looping combustion technology, and implemented into IECM. The use of IECM to

estimate the performance and cost of an IGCC power plant with CLC-based CO2 capture was

also demonstrated in this report.

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Documents

IECM Technical Documentation: Chemical Looping Combustion for Pre