Enhanced Cyclic Stability of Composite-Modified Iron-Based Oxygen Carriers in Methane Chemical Looping Combustion: Mechanistic Insights from Chemical Calculations

Featured Application

This study presents a high-performance, low-cost design strategy for composite-modified iron-based oxygen carriers in Chemical Looping Combustion technology. It significantly improves carbon capture efficiency and reduces industrial application costs, thereby providing a theoretical foundation for the clean utilization of fossil fuels and the realization of carbon neutrality goals.

Abstract

Chemical Looping Combustion (CLC) technology has emerged as a promising approach for carbon capture owing to its CO₂ separation capability, which addresses the pressing challenge of global climate change. Although iron-based oxygen carriers offer economic advantages owing to their abundance and low cost, their limited cyclic stability restricts their industrial deployment. This study focused on optimizing the performance of iron-based oxygen carriers through composite modification with Al₂O₃ and TiO₂. Using Cantera (2.5.0) software and the minimum Gibbs free energy principle, conversion rates and product distributions of Fe₂O₃, Fe₂O₃/Al₂O₃, and Fe₂O₃/TiO₂ were systematically analyzed under varying temperatures (800–950 °C), oxygen carrier-to-fuel molar ratios (O/C = 1–15), and pressures (0.1–1.0 MPa). The optimal conditions were identified as 900 °C, O/C = 8, and 0.1 MPa. After 50 simulation cycles, Fe₂O₃/Al₂O₃ and Fe₂O₃/TiO₂ achieved average total reaction counts of 503 and 543, respectively, substantially exceeding 296 cycles for Fe₂O₃. The results indicated that Al₂O₃ and TiO₂ improved cyclic stability via physical support and structural regulation mechanisms, thereby offering a practical carrier composite modification strategy. This study provides a theoretical basis for the development of high-performance oxygen carriers and supports the industrial application of CLC technology for efficient carbon capture and emission mitigation.

1. Introduction

Global temperatures have risen by more than 1 °C compared with pre-industrial temperatures [1]. Since 1998, record-breaking global temperatures have coincided with annual CO₂ emissions reaching approximately 33 billion tons, far surpassing the absorption capacity of natural carbon sinks [2]. Despite advances in energy efficiency and renewable energy adoption, CO₂ emissions from fossil fuel combustion and land-use changes continue to rise, largely driven by the energy, industrial, and transportation sectors [3,4,5]. Industrial activity remains a dominant emission source, and projections suggest that global CO₂ emissions could double by 2100 compared to the 2000 levels [6,7]. In this context, Carbon Capture, Utilization and Storage (CCUS) technology is essential for achieving greenhouse gas reduction targets, particularly in hard-to-abate sectors such as power generation and steel production [8,9]. Although energy efficiency and renewables are central to decarbonization, CCUS serves as a critical complementary strategy, enabling cleaner fossil fuel use during the transition to low-carbon systems, reducing environmental impacts, and bridging current and future energy economies [10]. Since 2017, CCUS projects have expanded at a compound annual rate exceeding 32% with significant progress across various regions. However, achieving the 2030 target of approximately 1 billion tons of annual storage will require overcoming significant technical, economic, and policy challenges [11,12].

The CCUS technology framework comprises three core components [13]: capture stage [14,15,16], utilization stage [17,18], and storage stage [19]. CO₂ capture can be achieved through three main approaches: pre-combustion, post-combustion, and oxy-fuel combustion carbon capture. Pre-combustion carbon capture involves converting fossil fuels into synthesis gas composed of H₂ and CO₂ via gasification reactions during the fuel pre-treatment stage, enabling effective CO₂ separation prior to combustion. This technological pathway offers the significant advantage of high capture efficiency. However, it requires extensive modifications to the existing industrial infrastructure, posing substantial capital investment barriers [20]. Post-combustion carbon capture is currently the most mature and widely applied capture method for removing CO₂ from flue gases after combustion [21]. Advanced materials such as polyethylene imine, zeolitic imidazolate frameworks, and pillar [5] arenes provide technical support [22,23,24]. However, high energy consumption associated with solvent regeneration and degradation during long-term operation remains a key bottleneck hindering large-scale applications [20]. Oxy-fuel combustion carbon capture replaces air with high-purity oxygen during fossil fuel combustion, generating flue gas primarily composed of CO₂ and water vapor, which simplifies the subsequent CO₂ separation. In addition to its efficient capture performance, oxy-fuel combustion can achieve near-zero emissions of nitrogen oxides (NO_x). Despite its advantages, this technology remains in the industrial demonstration stage, with the high cost of oxygen separation being the primary constraint limiting its commercial deployment [25,26].

As an emerging carbon capture technology, CLC offers an energy-efficient and environment-friendly alternative to conventional carbon capture methods. This technology eliminates the high energy consumption and cost barriers of traditional pre- and post-combustion approaches through its inherent CO₂ separation mechanism [27]. As illustrated in Figure 1, a typical dual-reactor configuration includes an air reactor (AR) and a fuel reactor (FR). In the fuel reactor, the oxygen carrier undergoes a reduction reaction (R1) with fuel (e.g., coal, natural gas, or biomass), generating CO₂ and H₂O. In the air reactor, the reduced oxygen carrier reacts with atmospheric oxygen during the oxidation reaction (R2), completing its regeneration. The regenerated oxygen carrier is then recirculated to the fuel reactor for the next reduction cycle with fresh fuel. After water vapor condensation and separation, a high-purity CO₂ stream is obtained [28].

The reaction occurring in the fuel reactor is as follows:

$$(2\mathrm{n}+\mathrm{m}){\mathrm{M}\mathrm{e}}_{\mathrm{x}}{\mathrm{O}}_{\mathrm{y}}\left(\mathrm{s}\right)+{\mathrm{C}}_{\mathrm{n}}{\mathrm{H}}_{2\mathrm{m}}\to (2\mathrm{n}+\mathrm{m}){\mathrm{M}\mathrm{e}}_{\mathrm{x}}{\mathrm{O}}_{\mathrm{y}-1}+{\mathrm{m}\mathrm{H}}_2\mathrm{O}+{\mathrm{n}\mathrm{C}\mathrm{O}}_2$$

(R1)

The reaction occurring in the air reactor is as follows:

$${\mathrm{M}\mathrm{e}}_{\mathrm{x}}{\mathrm{O}}_{\mathrm{y}-1}\left(\mathrm{s}\right)+1/2{\mathrm{O}}_2\left(\mathrm{g}\right)\to {\mathrm{M}\mathrm{e}}_{\mathrm{x}}{\mathrm{O}}_{\mathrm{y}}\left(\mathrm{s}\right)$$

(R2)

The overall reaction equation is:

$${\mathrm{C}}_{\mathrm{n}}{\mathrm{H}}_{2\mathrm{m}}+\left(\mathrm{n}+1/2\mathrm{m}\right){\mathrm{O}}_2\left(\mathrm{g}\right)\to {\mathrm{m}\mathrm{H}}_2\mathrm{O}+{\mathrm{n}\mathrm{C}\mathrm{O}}_2$$

(R3)

Through the cyclic oxygen transfer mechanism of the oxygen carrier, CLC directly produces a gas stream containing a high concentration of CO₂ without the need for complex gas separation systems, thereby fundamentally reducing capital investment. Compared with traditional post-combustion capture technologies, CLC can reduce CO₂ capture-related energy consumption by 30–50%, significantly enhancing the overall system energy efficiency. The indirect contact between fuel and air during the CLC process effectively inhibits the formation of nitrogen oxides, ensuring compliance with stringent emission standards [29].

The development of CLC technology from a conceptual origin to engineering applications in power generation and hydrogen production has undergone a progressive technological evolution. In 1983, Richter and Knoche formally proposed a CLC process for energy-efficient combustion, clearly defining the dual-reactor architecture and application of circulating oxygen carriers, thereby providing a theoretical paradigm for subsequent research [30]. From the 1990s to the 2000s, research has predominantly focused on innovations in oxygen carrier materials. Early studies mainly employed nickel-based oxides, which exhibit high reactivity under laboratory conditions. However, the toxicity of nickel and its high material cost limit its industrial application [31,32,33]. Consequently, iron-based oxides have gradually emerged as the mainstream choice, driven by the considerations of economic feasibility and environmental compatibility. Surface modification and composite structural design have been adopted to enhance the redox performance, laying the groundwork for pilot-scale applications [34]. Recent research has increasingly focused on Chemical Looping with Oxygen Uncoupling (CLOU) [35]. By releasing gaseous active oxygen, CLOU improves combustion kinetics and enables fuel conversion efficiencies exceeding 95%, thereby significantly boosting system energy performance [29]. This developmental trajectory demonstrates that CLC is rapidly progressing, from fundamental research to commercial deployment. Among the various materials investigated, iron-based carriers have attracted significant attention owing to their low cost, sulfur resistance, and sintering durability [36].

Systematic research on CLC has established an end-to-end technology chain from fundamental studies to engineering applications through integrated efforts in theoretical modeling, experimental analysis, and material development. In modeling, numerical simulations provide essential guidance for reactor design and process optimization [37]. In the experimental analysis, reactor configuration optimization is pivotal for efficient carrier circulation. The circulating fluidized bed (CFB) is the mainstream reactor design [38]. In material development, the enhancement of oxygen carrier performance remains the primary bottleneck for CLC industrialization. Tijani et al. systematically evaluated four transition metals (Cu, Co, Fe, and Ni) supported on various carrier oxides (Al₂O₃, CeO₂, TiO₂, and ZrO₂) as oxygen carriers in methane CLC. Research has identified Cu/Al₂O₃, Co/CeO₂, Fe/ZrO₂, and Ni/ZrO₂ as high-efficiency CLC oxygen carrier candidates, owing to their high oxygen transfer capacity, low activation energy, and resistance to coking [39]. Karimi et al. synthesized oxygen carriers using Al₂O₃ and TiO₂ as supports and Fe, Mn, Co, and Cu as active metals. Comprehensive characterization by X-ray diffraction, scanning electron microscopy, and energy-dispersive X-ray spectroscopy was conducted to examine the effects of oxygen carrier type, support material, and reaction temperature on the methane conversion rate, H₂ yield, and cyclic stability. The results demonstrated that iron-based oxygen carriers achieved the highest hydrogen production efficiency among all tested materials [40]. Previous studies have shown that Al₂O₃ and TiO₂ supports improve reducibility, thermal stability, and resistance to carbon deposition while promoting favorable metal–support interactions that enhance oxygen transfer capacity and cyclic stability [39]. Specifically, Al₂O₃ effectively suppresses sintering, whereas TiO₂ provides strong redox stability [40]. These findings underscore the suitability of Al₂O₃ and TiO₂ over other supports for developing high-efficiency, coke-resistant, and cyclable oxygen carriers [41].

This study utilized experimental data to explore oxygen carrier performance enhancement through numerical simulations. A multiphase thermodynamic equilibrium model was developed based on the Gibbs free energy minimization principle. The effects of temperature, oxygen carrier/fuel molar ratio, and pressure on fuel reactor reactions were simulated to identify the optimal conditions. In addition, the degradation behaviors of the three oxygen carriers over 50 cycles were comparatively assessed. The primary research objectives were to (1) identify optimal reaction conditions, (2) validate the modification effects of inert supports, (3) reveal the underlying reaction mechanisms, and (4) demonstrate that the cyclic durability of composite-modified oxygen carriers surpasses that of unmodified iron-based carriers. The results confirmed that the composite doping of metal carriers effectively enhanced both the reactivity and cyclic stability of iron-based oxygen carriers, thereby providing a theoretical foundation to support the industrial viability and economic feasibility of CLC technology.

2. Materials and Methods

2.1. Materials

A comparative analysis of commonly used oxygen carriers indicates that iron-based systems offer robust structural stability at high temperatures and moderate intrinsic reactivity, and their performance can be further improved through composite modification. Moreover, their abundance and low cost make them highly suitable for large-scale industrial applications. In this study, three oxygen carrier compositions were developed: Fe₂O₃, Fe₂O₃/Al₂O₃ (1:1), and Fe₂O₃/TiO₂ (1:1). A total of fifty redox cycles were conducted under optimal conditions determined through simulation-based optimization.

2.2. Software

Cantera is an open-source toolkit developed by Professor David G. Goodwin’s team at the California Institute of Technology for numerical simulation of chemical kinetics, thermodynamics, and multicomponent transport [42]. Saccone et al. developed a simplified framework mechanism, derived from the detailed kinetic scheme of Cantera, to model high-pressure, non-diluted CH₄/O₂ combustion in liquid rocket engines. This mechanism showed good agreement with high-pressure shock tube experimental data in terms of ignition delay time [43]. In CLC studies, Cantera enables accurate simulation of redox kinetics and integrates the NASA polynomial database to compute thermodynamic properties for over 8000 species from 298 to 3000 K with errors within ±2% [44]. Currently, Cantera (2.5.0) is widely used in combustion science, chemical engineering, and aerospace applications.

2.3. Principles

The simulation process adopted in this study was based on the principle of minimum Gibbs free energy [45,46,47,48]. This principle serves as the fundamental theoretical framework for thermodynamic equilibrium analysis and determines the final state of spontaneous system evolution under constant temperature and pressure conditions. In this state, the Gibbs free energy (G) of the system reaches its global minimum [49].

The classical definition of Gibbs free energy is

$$\mathsf{\Delta }\mathrm{G}=\mathsf{\Delta }\mathrm{H}-\mathrm{T}\mathsf{\Delta }\mathrm{S}$$

(1)

where $\mathrm{H}$ is the enthalpy of the system, which represents the total heat content (J/mol); $\mathrm{T}$ is the thermodynamic temperature (K); and $\mathrm{S}$ is the entropy (J/(mol·K)).

At equilibrium, the change in Gibbs free energy satisfies the condition $\mathsf{\Delta }\mathrm{G}$ = 0, corresponding to an extremum point in the system energy profile at which macroscopic properties, such as component concentrations and phase distributions, no longer vary with time. When the Gibbs free energy change for a reaction satisfies $\mathsf{\Delta }\mathrm{G}$ < 0, the reaction proceeds spontaneously. In contrast, when $\mathsf{\Delta }\mathrm{G}$ = 0, the forward and reverse reaction rates are equal, and the system attains dynamic equilibrium. Under constant temperature and pressure with zero non-expansion, chemical reaction systems spontaneously evolve toward decreasing Gibbs free energy ($\mathrm{G}$), ultimately reaching a global minimum and establishing a thermodynamic equilibrium state. This principle forms the theoretical basis for predicting the equilibrium compositions of multiphase systems via energy extremum criteria and is particularly applicable to the coupled analysis of complex multi-path reactions, including fuel pyrolysis, gasification, and redox processes in CLC systems.

The equilibrium state can be mathematically formulated as an optimization problem for Gibbs free energy:

$$\min \mathrm{G}={\sum }_{\mathrm{i}=1}^{\mathrm{n}}{\mathsf{\mu }}_{\mathrm{i}}{\mathrm{N}}_{\mathrm{i}}$$

(2)

where ${\mathsf{\mu }}_{\mathrm{i}}$ is the chemical potential of component i; and ${\mathrm{N}}_{\mathrm{i}}$ is the molar amount of component i.

2.4. Steps

2.4.1. Definition of Reaction System and Components

Definition of Reactants:

Fuel: methane (CH₄);

Oxygen carriers: Fe₂O₃, Fe₂O₃/Al₂O₃, and Fe₂O₃/TiO₂.

All possible reactants:

Gas phases: CO, CO₂, H₂, H₂O, CH₄, and O₂;

Solid phases: Fe₂O₃, Fe₃O₄, FeO, Al₂O₃, FeO·Al₂O₃, TiO₂, Fe₂O₃·TiO₂, FeO·TiO₂, and C;

Liquid phase: typically ignores in high-temperature CLC (>600 °C) but needs consideration in low-temperature scenarios (<100 °C).

2.4.2. Construction of Thermodynamic Database

The thermodynamic parameters of the species were obtained to form mechanism files.

NASA polynomials were used to describe temperature-dependent properties [50].

Selection of phase models:

Gas phase: ideal gas model (applicable for P < 1 MPa, T > 600 °C);

Solid phase: stoichiometric solid model (assuming oxygen carriers without lattice defects).

2.4.3. Set Model Constraints

The total thermodynamic function of the system represents the sum of thermodynamic functions of each phase:

$${\sum }_{\mathsf{\alpha }}\mathrm{d}\mathrm{G}\left(\mathsf{\alpha }\right)=-{\sum }_{\mathsf{\alpha }}\mathrm{S}\left(\mathsf{\alpha }\right)\mathrm{d}\mathrm{T}+{\sum }_{\mathsf{\alpha }}\mathrm{V}\left(\mathsf{\alpha }\right)\mathrm{d}\mathrm{p}+{\sum }_{\mathsf{\alpha }}{\sum }_{\mathrm{B}}{\mathsf{\mu }}_{\mathrm{B}}\left(\mathsf{\alpha }\right)\mathrm{d}{\mathrm{n}}_{\mathrm{B}}\left(\mathsf{\alpha }\right)$$

(3)

For any phase φ in the system,

$$\mathrm{d}\mathrm{G}\left(\mathsf{\alpha }\right)=-\mathrm{S}\left(\mathsf{\alpha }\right)\mathrm{d}\mathrm{T}+\mathrm{V}\left(\mathsf{\alpha }\right)\mathrm{d}\mathrm{p}+{\sum }_{\mathrm{B}}{\mathsf{\mu }}_{\mathrm{B}}\left(\mathsf{\alpha }\right)\mathrm{d}{\mathrm{n}}_{\mathrm{B}}\left(\mathsf{\alpha }\right)$$

(4)

where $\mathrm{G}\left(\mathsf{\alpha }\right)$, $\mathrm{S}\left(\mathsf{\alpha }\right)$, $\mathrm{V}\left(\mathsf{\alpha }\right)$, ${\mathsf{\mu }}_{\mathrm{B}}\left(\mathsf{\alpha }\right)$, and ${\mathrm{n}}_{\mathrm{B}}\left(\mathsf{\alpha }\right)$ represent the Gibbs function, entropy, volume, chemical potential of component B in phase φ, and the amount of substance of component B, respectively.

The partial molar Gibbs function of component B in the mixture is defined as the chemical potential of B:

$${\mathsf{\mu }}_{\mathrm{B}}={\mathrm{G}}_{\mathrm{B}}$$

(5)

The standard Gibbs free energy of any species at temperature T is given by

$${\mathrm{G}}_{\mathrm{B}}^{\mathrm{o}}={\mathrm{H}}_{\mathrm{B}}^{\mathrm{o}}-\mathrm{T}\times {\mathrm{S}}_{\mathrm{B}}^{\mathrm{o}}$$

(6)

where ${\mathrm{H}}_{\mathrm{B}}^{\mathrm{o}}$ and ${\mathrm{S}}_{\mathrm{B}}^{\mathrm{o}}$ represent the enthalpy and entropy of species B, respectively, at temperature $\mathrm{T}$. These are obtained through parameterization of NASA polynomial coefficients.

$${\displaystyle \frac{{\widehat{\mathrm{h}}}^{\circ }\left(\mathrm{T}\right)}{\stackrel{-}{\mathrm{R}}\mathrm{T}}}=-{\mathrm{a}}_0{\mathrm{T}}^{-2}+{\mathrm{a}}_1{\displaystyle \frac{\ln \mathrm{T}}{\mathrm{T}}}+{\mathrm{a}}_2+{\displaystyle \frac{{\mathrm{a}}_3}{2}}\mathrm{T}+{\displaystyle \frac{{\mathrm{a}}_4}{3}}{\mathrm{T}}^2+{\displaystyle \frac{{\mathrm{a}}_5}{4}}{\mathrm{T}}^3+{\displaystyle \frac{{\mathrm{a}}_6}{5}}{\mathrm{T}}^4+{\displaystyle \frac{{\mathrm{a}}_7}{\mathrm{T}}}$$

(7)

$${\displaystyle \frac{{\widehat{\mathrm{s}}}^{\circ }\left(\mathrm{T}\right)}{\stackrel{-}{\mathrm{R}}}}=-{\displaystyle \frac{{\mathrm{a}}_0}{2}}{\mathrm{T}}^{-2}-{\mathrm{a}}_1{\mathrm{T}}^{-1}+{\mathrm{a}}_2\ln \mathrm{T}+{\mathrm{a}}_3\mathrm{T}+{\displaystyle \frac{{\mathrm{a}}_4}{2}}{\mathrm{T}}^2+{\displaystyle \frac{{\mathrm{a}}_5}{3}}{\mathrm{T}}^3+{\displaystyle \frac{{\mathrm{a}}_6}{4}}{\mathrm{T}}^4+{\mathrm{a}}_8$$

(8)

$${\displaystyle \frac{{\widehat{\mathrm{c}}}_{\mathrm{p}}^{\circ }\left(\mathrm{T}\right)}{\stackrel{-}{\mathrm{R}}}}={\mathrm{a}}_0{\mathrm{T}}^{-2}+{\mathrm{a}}_1{\mathrm{T}}^{-1}+{\mathrm{a}}_2+{\mathrm{a}}_3\mathrm{T}+{\mathrm{a}}_4{\mathrm{T}}^2+{\mathrm{a}}_5{\mathrm{T}}^3+{\mathrm{a}}_6{\mathrm{T}}^4$$

(9)

where ${\widehat{\mathrm{h}}}^{\circ }\left(\mathrm{T}\right)$ represents the molar enthalpy under the reference state (J·mol⁻¹); ${\widehat{\mathrm{s}}}^{\circ }\left(\mathrm{T}\right)$ represents the absolute molar entropy (J·mol⁻¹·K⁻¹); ${\widehat{\mathrm{c}}}_{\mathrm{p}}^{\circ }\left(\mathrm{T}\right)$ represents the molar heat capacity at a constant pressure (J·mol⁻¹·K⁻¹), all of which vary with temperature T; R is the universal gas constant, typically 8.314 J·mol⁻¹·K⁻¹; and ${\mathrm{a}}_0,{\mathrm{a}}_1...{\mathrm{a}}_8$ are the fitting coefficients of the NASA polynomials to describe the variation in thermodynamic properties with temperature. They were obtained from the NASA thermodynamic database built into Cantera software, which was fitted based on experimental data [51].

2.4.4. Selection of Solution Algorithm

• The AR reaction temperature was set to 800 °C. This optimized temperature ensured stable regeneration of oxygen carriers while maintaining structural integrity throughout cyclic operation [52].
• Selection of optimal temperature: The combustion conversion rates of three oxygen carriers were calculated for the first five cycles at 800 °C, 850 °C, 900 °C, and 950 °C under 0.1 MPa and $\mathrm{O}/\mathrm{C}$ = 4 (oxygen carrier in large excess relative to fuel). The relationship between cycle number and conversion rate at different temperatures was then output.
• Selection of optimal $\mathrm{O}/\mathrm{C}$ ratio: The molar fraction of cyclic products of Fe₂O₃ oxygen carrier was calculated within the $\mathrm{O}/\mathrm{C}$ range of 0–15 at 0.1 MPa. The relationship between product distribution and $\mathrm{O}/\mathrm{C}$ ratio was then determined.
• Selection of optimal pressure: The molar fraction of cyclic products of Fe₂O₃ oxygen carrier was calculated within the pressure range of 0.1–1.1 MPa at the optimal temperature and $\mathrm{O}/\mathrm{C}$ ratio. The relationship between product distribution and pressure was output.
• Degradation simulation: The combustion conversion rates of three oxygen carriers over 50 cycles were calculated under optimal temperature, optimal O/C ratio, optimal pressure, and with degradation considered. The relationship between cycle number and conversion rate at different $\mathsf{\phi }\mathrm{B}$ values was then analyzed.
• Termination conditions: The cycle was terminated when the oxygen carrier was insufficient (i.e., lack of reactant in the Fuel Reactor) or when $\mathrm{E}\mathrm{i}\mathrm{n}$ < $\mathrm{E}\mathrm{o}\mathrm{u}\mathrm{t}$ (i.e., the reduction reaction in the Fuel Reactor could not proceed spontaneously).

• Definition of key parameters:

$\mathrm{M}\mathrm{o}\mathrm{x}$: weight of metal oxide in the oxidized state (g);

$\mathrm{M}\mathrm{r}\mathrm{e}$: weight of metal oxide in the reduced state (g);

$\mathrm{M}\mathrm{n}$: weight of metal oxide after the nth cycle (g);

$\mathrm{X}\mathrm{n}$: oxygen carrier conversion rate after the nth cycle (%);

$$\mathrm{X}\mathrm{n}=(\mathrm{M}\mathrm{o}\mathrm{x}-\mathrm{M}\mathrm{n})/(\mathrm{M}\mathrm{o}\mathrm{x}-\mathrm{M}\mathrm{r}\mathrm{e})$$

(10)

$\mathrm{O}/\mathrm{C}$: molar ratio of oxygen carrier to fuel;

$\mathsf{\phi }\mathrm{B}$: degradation rate of oxygen carrier B (%).

With degradation considered:

$$\mathrm{X}\mathrm{n}=[\mathrm{M}\mathrm{o}\mathrm{x}-\mathrm{M}\mathrm{n}(1-\mathsf{\phi }\mathrm{B}\left)\mathrm{n}\right]/(\mathrm{M}\mathrm{o}\mathrm{x}-\mathrm{M}\mathrm{r}\mathrm{e})$$

(11)

$\mathrm{E}\mathrm{i}\mathrm{n}$: Total input energy to the system (kJ).

$\mathrm{E}\mathrm{o}\mathrm{u}\mathrm{t}$: Total output energy from the system (kJ).

$\mathrm{E}\mathrm{i}\mathrm{n}$ and $\mathrm{E}\mathrm{o}\mathrm{u}\mathrm{t}$ satisfy:

$$\mathrm{E}={\sum }_{\mathrm{i}}{\mathrm{n}}_{\mathrm{i}}\times ∆{\mathrm{H}}_{\mathrm{i}}\left(\mathrm{T}\right)$$

(12)

where

${\mathrm{n}}_{\mathrm{i}}$ is the molar amount of reactant or product;

$∆{\mathrm{H}}_{\mathrm{i}}\left(\mathrm{T}\right)$ is the enthalpy of the i substance at temperature T.

2.4.5. Predicted Operating Condition Reactions

The thermodynamic simulations in this study employed Fe₂O₃, Fe₂O₃/Al₂O₃, and Fe₂O₃/TiO₂ as oxygen carriers with CH₄ as the fuel. The FR primarily generated C, H₂, CO, CO₂, CH₄, and H₂O, which subsequently underwent redox reactions with oxygen carriers to drive the combustion process.

Table 1. Reactions between fuel and oxygen carriers.

FR Products	Reaction	ΔH (kJ/mol)
C	C + 6Fe2O3 → CO2 + 4Fe3O4	95.98	(R4)
	C + 2Fe2O3 → CO2 + 4FeO	164.88	(R5)
H2	H2 + 3Fe2O3 → H2O + 2Fe3O4	−2.01	(R6)
	H2 + Fe2O3 → H2O + 2FeO	32.44	(R7)
CO	CO + 3Fe2O3 → CO2 + 2Fe3O4	−37.67	(R8)
	CO + Fe2O3 → CO2 + 2FeO	−3.22	(R9)
	CO + Fe2O3 + 2Al2O3 → 2FeO·Al2O3 + CO2	122.52	(R10)
	CO + Fe2O3·TiO2 + TiO2 → 2FeO·TiO2 + CO2	−1.57	(R11)
CH4	CH4 + 12Fe2O3 → 2H2O + CO2 + 8Fe3O4	221.76	(R12)
	CH4 + 4Fe3O4 → 12FeO + 2H2O + CO2	365.93	(R13)
	CH4 + 4Fe2O3 → 2H2O + CO2 + 8FeO	317.87	(R14)
	CH4 + 4Fe2O3 + 8Al2O3 → 8FeO·Al2O3 + 2H2O + CO2	233.93	(R15)
	CH4 + 12Fe2O3·TiO2 → 8Fe3O4 + 12TiO2 + CO2 + 4H2O	−347.13	(R16)
	CH4 + 4Fe2O3·TiO2 + 4TiO2 → 8FeO·TiO2 + CO2 + 2H2O	323.35	(R17)
	CH4 + 4Fe3O4 + 12TiO2 → 12FeO·TiO2 + 2CO2 + 2H2O	23.23	(R18)

recorded reactions between fuel and oxygen carriers in FR;

Table 2. Reactions between gases and carbon.

Reaction	ΔH (kJ/mol)
C + H2O → CO + H2	131.46	(R19)
CO + H2 → C + H2O	−131.31	(R20)
CO + H2O → H2 + CO2	−41.14	(R21)
C + CO2 → 2CO	172.67	(R22)
2CO → C + CO2	−172.46	(R23)
CH4 + H2O → CO + 3H2	223.70	(R24)
CH4 + CO2 → 2H2 + 2CO	247.29	(R25)
2CO + 2H2 → CH4 + CO2	−247.29	(R26)
CO2 + 4H2 → CH4 + 2H2O	−165.00	(R27)
C + 2H2 → CH4	−74.87	(R28)
CH4 → C + 2H2	75.00	(R29)

recorded reactions between gases and carbon in FR;

Table 3. Reactions in AR.

Reaction	ΔH (kJ/mol)
4Fe3O4 + O2 → 6Fe2O3	−483.07	(R30)
4FeO + O2 → 2Fe2O3	−554.68	(R31)
4FeO·TiO2 + O2 → 2Fe2O3 + 4TiO2	−562.82	(R32)
4FeO·Al2O3 + O2 → 2Fe2O3 + 4Al2O3	−562.82	(R33)

recorded reactions in AR.

3. Results and Discussion

3.1. Selection of Optimal Reaction Conditions

3.1.1. Selection of Optimal Temperature

Figure 2 shows that Fe₂O₃ conversion increased over cycles 1–5 before reaching a plateau across temperatures. At 900 °C, the conversion was markedly higher than that at 800–850 °C and approached that at 950 °C. Notably, in cycles 4–5, the Fe₂O₃ conversion at 950 °C (~91.7% and 92.8%) fell below that at 900 °C (~93.9% and 94.8%), consistent with high-temperature sintering of the oxygen carrier [53].

Figure 3 shows a similar trend for Fe₂O₃/Al₂O₃, but without the 900–950 °C decrease observed for Fe₂O₃. This confirms that Al₂O₃ promotes the formation of stable alumina–ferrite phases at elevated temperatures, suppresses sintering, preserves surface area, and thereby sustains higher activity [54].

Figure 4 shows that Fe₂O₃/TiO₂ exhibited the same overall behavior and consistently maintained high conversion from 900 to 950 °C, indicating excellent thermal stability [55].

A comprehensive analysis of the conversion rate performance of the three oxygen carriers at different temperatures (Figure 5) revealed that all exhibited significantly higher conversion rates at 900 °C than at 800 °C and 850 °C, with levels approaching those at 950 °C. While the conversion rate of Fe₂O₃ decreased at 950 °C compared to 900 °C, the two composite oxygen carriers (Fe₂O₃/Al₂O₃ and Fe₂O₃/TiO₂) showed slight improvements, preliminarily verifying their superior thermal stability. Although the conversion rate at 900 °C was slightly lower than that at 950 °C (by approximately 1%), operation at 950 °C resulted in approximately 11.5% greater energy loss than that at 900 °C. Considering economic feasibility, operating the fuel reactor at 900 °C could effectively reduce the operational costs. Therefore, 900 °C is the optimal temperature for balancing conversion efficiency, cyclic stability, and energy consumption. At this temperature, all three iron-based oxygen carriers demonstrated excellent reactivity and thermal stability, representing the best balance between energy input and economic returns and meeting the technical requirements of CLC for both performance and economic viability [56].

3.1.2. Selection of Optimal O/C Ratio

Analysis of the molar fraction variation curves of solid products during CLC of iron-based oxygen carriers with methane at different O/C ratios (Figure 6a) revealed that when the O/C ratio was low (1–2), the mole fraction of FeO approached 100%, whereas that of Fe₂O₃ was nearly zero, indicating extensive reduction in the carrier to low-valence oxides. As the O/C ratio increased (2–8), the mole fraction of FeO rapidly decreased because of the participation of more oxides in the reaction, promoting the transformation of FeO into higher-valence states. With a further increase in O/C (8–15), FeO continued to decrease and stabilize at a low level. Within the O/C range of 2–12, Fe₃O₄ gradually increased and became the dominant solid product, reflecting progressive partial oxidation. Beyond this range, Fe₃O₄ diminished, whereas unreacted Fe₂O₃ increased.

Analysis of the mole fraction curves of gaseous products (Figure 6b) demonstrated that as the O/C ratio increased from 0, the mole fraction of H₂O increased rapidly. When O/C was within the range of 0–8, the mole fractions of H₂ and CO initially increased and then declined. This indicated that hydrogen and carbon in methane were not fully oxidized during this stage. With increasing O/C, H₂ and CO dropped to near zero as sufficient oxide drove complete oxidation to H₂O and CO₂. When O/C > 8, the mole fraction of H₂O stabilized at approximately 66.7%, and that of CO₂ stabilized at approximately 33.3%. These findings demonstrate that under sufficient oxygen conditions, the H and C in methane are fully oxidized to H₂O and CO₂.

Combining the solid and gas-phase trends, when O/C < 8, H and C in methane were incompletely oxidized, resulting in residual H₂ and CO and leading to energy loss. Within the O/C range of 8–12, Fe₃O₄ was the predominant solid product. Fe₃O₄ exhibits both oxidizing and reducing capabilities, enabling effective oxygen release during the CLC cycle [57]. Simultaneously, the mole fractions of H₂ and CO approached zero, indicating nearly complete methane combustion, which met the requirements for efficient CLC and facilitated CO₂ separation. When O/C > 12, the proportion of Fe₂O₃ increased significantly. Although the product purity improved slightly, this led to the excessive consumption of oxygen carrier resources and increased operational costs. Overall, a comprehensive evaluation suggested that at O/C = 8, the redox behavior of the oxygen carrier achieved an optimal balance, ensuring complete methane conversion while minimizing oxygen consumption, thereby balancing conversion efficiency and resource utilization.

3.1.3. Selection of Optimal Pressure

Analysis of the mole-fraction variation in solid products during CLC of iron-based oxygen carriers with methane at different pressures (Figure 7a) shows that as pressure increased from 0.1 to 1.0 MPa, the Fe₃O₄ fraction continuously rose. This indicated a lower overall reduction degree of the carrier at higher pressures, favoring the formation of intermediate-valence oxides (Fe₃O₄). Conversely, the FeO fraction decreased with increasing pressure, consistent with the reduced extent of carrier reduction; at 0.1 MPa, FeO was relatively abundant, implying deeper reduction under low pressure. Across the entire range, Fe₂O₃ declined only slightly with increasing pressure, suggesting that more of the carrier participates in the reaction at elevated pressures but converts predominantly to Fe₃O₄ rather than to FeO.

Analysis of the mole-fraction variation in gaseous products (Figure 7b) indicates that with increasing pressure, H₂O gradually decreased; at 0.1 MPa, H₂O reached its maximum, indicating that lower pressure favors complete oxidation of hydrogen in methane. CO₂ also decreased slightly with pressure, whereas CO and H₂ increased, reflecting a progressive shift toward incomplete oxidation. At 0.1 MPa, CO and H₂ remained extremely low, confirming more complete methane conversion and minimal formation of partial oxidation products. As pressure increased, the reaction shifted toward CO and H₂, consistent with the lower reduction degree of the oxygen carrier.

Overall, pressurization from 0.1 to 1.0 MPa slightly increased carrier participation but compromised combustion completeness. At 0.1 MPa, the gas phase contained a higher proportion of CO₂ with negligible CO and H₂, whereas the solid phase contained more FeO, indicating a more complete methane conversion under this condition. From an operational perspective, 0.1 MPa (near atmospheric pressure) avoids complex compression, thereby reducing energy demand and equipment costs, enhancing safety and operational stability, and simplifying the process. Considering oxidation degree, carrier-reduction behavior, and process economics, 0.1 MPa was identified as the optimal reaction pressure.

3.2. Degradation Simulation

An analysis of the wear characteristics of the three iron-based oxygen carriers during 50 cycles of CLC with methane revealed that the Fe₂O₃ oxygen carrier (Figure 8a) exhibited a rapid increase in the conversion rate, reaching approximately 94.8% during the initial reaction stages, followed by a gradual decline as the number of cycles increased. By the 50th cycle, a significant decrease in the conversion rate was observed, indicating a progressive decline in the reactivity of Fe₂O₃ oxygen carriers due to wear, carbon deposition, and structural degradation. The magnified view shows that the conversion rate decay trend remained consistent across all three simulations, with the average curve clearly illustrating the declining pattern of conversion rate with cycling, thereby reflecting the limited cyclic stability of Fe₂O₃ oxygen carriers [58].

For the Fe₂O₃/Al₂O₃ oxygen carrier (Figure 8b), the conversion rate increased rapidly and remained high during the initial cycles. Compared with Fe₂O₃, it exhibited a significantly smaller amplitude of decay in the conversion rate throughout the cycling process. Even after 50 cycles, the conversion rate remained relatively high, indicating that the incorporation of Al₂O₃ improved oxygen carrier performance. In the magnified view, the decay rate curves exhibited smooth trends across all simulations, suggesting that the Fe₂O₃/Al₂O₃ oxygen carrier possesses excellent cyclic stability. The addition of Al₂O₃ can provide structural support and protection, thereby mitigating degradation induced by wear [59].

Similarly, the Fe₂O₃/TiO₂ oxygen carrier (Figure 8c) exhibited a rapid increase and early stabilization during conversion. Over 50 cycles, the decay was less pronounced than that of Fe₂O₃, and the overall conversion remained consistently higher. The magnified view showed a moderate and uniform decay with low dispersion, suggesting that TiO₂ improves resistance to wear and structural degradation, thereby sustaining reactivity during cycling [60].

A comparative analysis of the three oxygen carriers (Figure 8d) showed that both Fe₂O₃/Al₂O₃ and Fe₂O₃/TiO₂ consistently maintained higher conversion than Fe₂O₃, with the advantage becoming more pronounced in later cycles. The magnified comparison highlights their smaller decay amplitudes, confirming that the addition of Al₂O₃ and TiO₂ effectively mitigated conversion loss. Compared with Fe₂O₃, the composites sustained higher reactivity over extended cycling, demonstrating superior resistance to wear and structural degradation [61,62,63,64].

It is worth noting that in the simulation, changes in the carrier’s chemical properties during long-term reactions and the presence of impurities make it difficult to completely reproduce the conversion rate variations observed under actual working conditions. In the simulation, a simplified particle wear model was used, which likely underestimated long-term performance degradation [61]. Although the simulation reliably predicts initial reactivity, it cannot capture long-term structural instability because phase separation and sintering are difficult to model at the microscopic scale [65]. Furthermore, the simulation neglects the actual impurities present in the reactants, resulting in performance estimates that are higher than those in actual applications.

3.3. Predicted Total Number of Cycles

Figure 9 shows the degradation trends of Fe₂O₃, Fe₂O₃/Al₂O₃, and Fe₂O₃/TiO₂ oxygen carriers during the first 50 cycles of CLC with methane. The total number of expected redox cycles was predicted based on the established termination conditions.

The Fe₂O₃ oxygen carrier (Figure 9a) demonstrated an average total reaction count of 296 across 50 simulations, indicating relatively poor cyclic stability and high sensitivity to fluctuations in the reaction conditions.

The Fe₂O₃/Al₂O₃ oxygen carrier (Figure 9b) achieved an average total reaction count of 503, which was significantly higher than that of Fe₂O₃, indicating its superior cyclic performance and enhanced durability under repeated redox conditions.

The Fe₂O₃/TiO₂ oxygen carrier (Figure 9c) reached the highest average reaction count of 543 among the three carriers, suggesting the slowest degradation rate and excellent long-term cyclic stability.

Both composite oxygen carriers (Fe₂O₃/Al₂O₃ and Fe₂O₃/TiO₂) exhibited significantly higher average total reaction counts than Fe₂O₃, reflecting their greater resistance to wear, sintering, and structural degradation during cyclic operation. This enhanced durability enables the maintenance of high reactivity across extended cycling periods, thereby satisfying the long-term operational requirements of CLC systems. These findings are consistent with those of previous studies demonstrating that composite oxygen carriers improve structural integrity during repeated redox cycles.

4. Conclusions

This study optimized the cyclic performance of iron-based oxygen carriers by investigating the effects of reaction conditions and metal oxide doping. A single-variable control strategy combined with a thermodynamic equilibrium simulation based on the principle of minimum Gibbs free energy was adopted. Using Cantera software, the conversion rates and product distributions of three oxygen carriers (Fe₂O₃, Fe₂O₃/Al₂O₃, and Fe₂O₃/TiO₂) were systematically analyzed under varying temperatures (800–950 °C), oxygen carrier-to-fuel molar ratios (O/C = 0–15), and pressures (0.1–1.0 MPa). The degradation performance of over 50 cycles was evaluated, and the total number of cycles was predicted according to the established termination criteria. The results indicated that optimal comprehensive reactor performance was achieved at 900 °C, O/C = 8, and 0.1 MPa. The incorporation of Al₂O₃ and TiO₂ significantly enhanced cyclic stability and wear resistance. These findings not only clarified the mechanisms by which Al₂O₃ and TiO₂ improved the cyclic performance of iron-based oxygen carriers through structural support and stabilization but also provided a practical “carrier composite modification” strategy for designing high-performance oxygen carriers. From an industrial integration perspective, the optimal conditions identified in this study are highly favorable. Operating at atmospheric pressure simplifies plant design and improves operational safety. Furthermore, a moderate temperature of 900 °C facilitates easier thermal integration with steam cycle systems. The abundance and low cost of iron oxide further strengthen the commercial viability of these materials. Overall, by combining high cyclic stability, low cost, and compatibility with existing infrastructure, these materials represent a critical step toward the commercialization of CLC technology and its integration into a sustainable, low-carbon industrial future.