energies

Article
Pressurized Chemical Looping for Direct Reduced Iron
Production: Carbon Neutral Process Configuration
and Performance
Nicole Bond *, Robert Symonds and Robin Hughes

Natural Resources Canada, CanmetENERGY-Ottawa, Ottawa, ON K1A 1M1, Canada;

robert.symonds@nrcan-rncan.gc.ca (R.S.); robin.hughes@nrcan-rncan.gc.ca (R.H.)
* Correspondence: nicole.bond@nrcan-rncan.gc.ca

Abstract: To achieve net-zero iron and steel production by 2050, many iron and steel producers are
turning to direct reduced iron (DRI)—electric arc furnace (EAF) steel production as an opportunity
to achieve significant CO2 emissions reductions relative to current levels. However, additional
innovations are required to close the gap between DRI and net-zero steel. Pressurized chemical
looping-DRI (PCL-DRI) is a novel technology explored to meet this target, in which the reformer
firebox and fired process gas heaters are replaced with PCL combustion units. Captured CO2 is
conditioned and compressed for pipeline transportation and storage/utilization. The performance of
two different PCL-DRI configurations relative to traditional DRI processes was explored via process
simulation: a Midrex-type process and an Energiron-type process. The PCL-DRI processes were
shown to have equivalent or lesser total fuel consumption (8% reduction) compared to the base
cases, and greater process water production (170–260% increase), with minimal or no loss in thermal
efficiency. PCL-DRI is a strong competitor to alternative methods of reaching net-zero DRI due to
lower energy penalties for carbon capture, no required changes to stream chemistry in or out of the
EAF, and no requirement for hydrogen infrastructure.
Citation: Bond, N.; Symonds, R.;
Hughes, R. Pressurized Chemical Keywords: pressurized chemical looping combustion; syngas production; carbon neutral ironmaking;
Looping for Direct Reduced Iron direct reduced iron; CO2 capture; decarbonization
Production: Carbon Neutral Process
Configuration and Performance.
Energies 2022, 15, 5219. https://
doi.org/10.3390/en15145219 1. Introduction
Academic Editor: Arturo Cabello 1.1. Background

Received: 24 May 2022

Ironmaking is one of Canada’s top CO2 emitting industries. Canada ranked eighth in
Accepted: 8 July 2022
global iron ore production, producing 52.4 Mt of iron ore and 11 Mt of crude steel in 2020 [1].
Published: 19 July 2022
Globally, iron and steel production results in 4–8% of anthropogenic CO2 emissions, with
specific emissions ranging from 1.4–3.5 t CO2 /t liquid steel [2]. Canadian iron and steel
Publisher’s Note: MDPI stays neutral
producers have committed to produce net-zero iron and steel by 2050, as a part of the global
with regard to jurisdictional claims in
target to limit Earth’s temperature rise to 1.5 ◦ C [3,4].
published maps and institutional affil-
There are two main processes dominating the steel production industry; integrated
iations.
steel mills (ISM), which use a blast furnace (BF) and basic oxygen furnace (BOF), and a mini-
mill which uses direct reduced iron (DRI) and steel scrap to charge an electric arc furnace
(EAF). The BF-BOF route represents the majority of steel production, with nearly 14-fold
Copyright: © 2022 by the authors.
more steel produced in this manner globally compared to the DRI-EAF route in 2017 [5].
Licensee MDPI, Basel, Switzerland. Transitioning to DRI-EAF can reduce the CO2 emission intensity by 61–68% relative to
This article is an open access article BF-BOF, but further innovation is required to achieve net zero [6,7].
distributed under the terms and The two largest market shares for current DRI production are the Midrex and Energiron
conditions of the Creative Commons processes [8]. Process flow diagrams for both of these processes are shown in Figure 1.
Attribution (CC BY) license (https:// In both cases, solid iron ore is charged into the top of the shaft furnace. A reducing gas
creativecommons.org/licenses/by/ containing primarily CO, H2 , and CH4 is injected from the side and reacts with the ore to
4.0/). reduce it to solid metallic iron. The depleted reducing gas exits the top of the shaft furnace

Energies 2022, 15, 5219. https://doi.org/10.3390/en15145219 https://www.mdpi.com/journal/energies

Energies 2022, 15, 5219 2 of 17

(referred to herein as “top gas”). A portion of the top gas is recycled back for re-use as
reducing gas, while the remainder is combusted as fuel to provide heat for reforming and
sensible heating of the reducing gas. There are a few key differences between the Midrex
and Energiron processes: (1) the shaft furnace operates at 100–200 kPa for Midrex, while
Energiron operates at 600–800 kPa; (2) Midrex operates a dry reformer, while Energiron
uses steam methane reforming and/or gases from other parts of the steel mill; (3) Energiron
includes a CO2 absorber to purge CO2 from the recycled top gas, though this stream is
typically vented, and (4) Energiron includes a fired gas heater that creates a second stream
of CO2 emissions [9–12].
Chemical looping combustion (CLC) is a means of heat production from fossil fuels
with inherent separation and capture of CO2 . Through the use of a solid oxygen carrier,
the combustion of the fuel is isolated from air without requiring a costly air separation
unit [13–15]. The typical set-up requires two fluidized bed reactors: (1) the air reactor, where
air oxidizes the oxygen carrier and (2) the fuel reactor, where fuel reacts with the oxygen in
the oxygen carrier, producing a pure stream of CO2 and water vapor [16]. Pressurization of
the process (pressurized chemical looping, or PCL), provides several additional advantages,
including reduced equipment size and capital costs, increased reaction rates, and the
potential for latent heat recovery with process water recovery [17,18]. This work proposes
novel configurations for net-zero DRI production through the integration of PCL wherever
fossil fuels are being consumed to produce heat. The performance and opportunities for
PCL-DRI are compared and contrasted with the most common DRI facilities currently
in operation.

1.2. Design Specification for DRI

Given the emergence of DRI as a promising option for the iron and steel industry to
substantially reduce CO2 emissions, the scale of future DRI facilities must be comparable to
that of existing and future iron and steel facilities. Thus, this work considers a DRI facility
with a capacity of producing 2 million tonnes per annum (MTPA) of DRI to be comparable
to the Voestalpine Midrex plant constructed in Texas, USA in 2017 [19]. To arrive at hourly
flow rates, an annual availability of 91.3% was applied. This is based on the performance
guarantee for Midrex facilities currently on the market [11]. For ease of comparison, the
availability of the PCL-DRI facilities was considered to be the same. Given that Midrex
and Energiron are the two DRI technologies with the largest market share, both of these
configurations were used as a basis for comparison to potential PCL-DRI configurations.
The key units of a DRI facility are described in the following subsections.

1.2.1. Shaft Furnace

The shaft furnace is the heart of the DRI process. Iron ore (usually in pellet form) is fed
to the top of the furnace, while reducing gas is fed from the side and flows counter-currently
to the ore. The reducing gas enters the furnace at 850–900 ◦ C for Midrex or 920–1050 ◦ C for
Energiron III [12]. The upper region of the furnace is the reducing zone, where iron oxide
is reduced upon interaction with the gas, as shown in Equations (1)–(6) below. Methane
reforming and water-gas shift reactions (Equations (7) and (8)) also occur in this region [12].
◦
3Fe2 O3 +H2 → 2Fe3 O4 +H2 O ∆Hr = 0.7 kJ/mol H2 (1)
◦
Fe3 O4 +H2 → 3FeO + H2 O ∆Hr = 7.8 kJ/mol H2 (2)
◦
FeO + H2 → Fe + H2 O ∆Hr = 24.0 kJ/mol H2 (3)
◦
3Fe2 O3 +CO → 2Fe3 O4 +CO2 ∆Hr = −40.4 kJ/mol CO (4)
◦
Fe3 O4 +CO → 3FeO + CO2 ∆Hr = −36.9 kJ/mol CO (5)
◦
FeO + CO → Fe + CO2 ∆Hr = −17.2 kJ/mol CO (6)
◦
CH4 +H2 O
CO + 3H2 ∆Hr = −206.2 kJ/mol CH4 (7)
 Energies 2022, 15, 5219 3 of 17

Energies 2022, 15, x FOR PEER REVIEW 3 of 18

◦
CO + H2 O
CO2 + H2 ∆Hr = −41.2 kJ/mol CO (8)

(a)

(b)
Figure 1. Simplified process configurations for (a) a Midrex process and (b) an Energiron III process.
Figure 1. Simplified process configurations for (a) a Midrex process and (b) an Energiron III process.
For ease of viewing, not all heat exchangers or heat integration are shown. Stream properties are
Forin
defined ease
theof viewing, not all
Supplementary heat exchangers
Materials (Tables S1or heat
and S2).integration are shown. Stream properties are
defined in the Supplementary Materials (Tables S1 and S2).
1.2.1. Shaft Furnace
The shaft furnace is the heart of the DRI process. Iron ore (usually in pellet form) is
fed to the top of the furnace, while reducing gas is fed from the side and flows counter-
Energies 2022, 15, 5219 4 of 17

The depleted reducing gas exits the top of the furnace as top gas. In the lower region,
the iron cools. Depending on the type of DRI product from the facility (hot DRI, cold
DRI, or hot briquette iron), there may be additional cooling gas supplied to this region.
This is also where carbon deposition onto the iron occurs via methane decomposition and
the reverse Boudouard reaction [12,20]. The carbon contained within the DRI product
improves the efficiency of the downstream EAF [6]. DRI products from the shaft furnace
typically meet the following specifications: 90–94% iron, 92–97% metallization, and 0.5–4%
carbon [10,11].

1.2.2. Reformer
The reformer supplies fresh syngas makeup, which is sent to the shaft furnace in combi-
nation with recycled top gas for the reduction of iron ore. In the Midrex process (Figure 2a),
dry reforming reactions dominate, and the reformer operates at 200–300 kPa(g) [9]. All
steam required for reforming is contained within the top gas recycled to the reformer
feed, while additional natural gas is supplied for conversion. In the Energiron III process
(Figure 2b), the reformer functions as a more traditional steam-methane reformer (SMR),
and the recycled top gas is combined with the produced syngas further downstream. In this
work, syngas exits the reformer at 925 ◦ C and 220 kPa(a) (Midrex-type processes) or 880 ◦ C
and 945 kPa(a) (Energiron-type processes).

1.2.3. Top Gas Scrubber

The top gas scrubber is a wet scrubber that cleans and removes moisture from the top
gas as it exits the shaft furnace [21]. In this work, the scrubber was operated at an outlet
temperature of 50–51◦ C to achieve adequate water knock-out from the top gas recycled as
reducing gas.

1.2.4. Top Gas Compressor

The top gas compressor re-pressurizes the fraction of the top gas that is recycled
for incorporation into the reducing gas. In this work, the discharge pressure from the
top gas compressor is 300 kPa(a) (Midrex-type processes) or 1000 kPa(a) (Energiron-type
processes). In the base DRI configurations, the portion of the top gas used as fuel in the
reformer firebox do not pass through this compressor, as it is already at sufficient pressure
for passage through the firebox and downstream heat exchangers.

1.2.5. CO2 Removal Unit (Energiron III-Type Configurations)

The CO2 removal unit in the Energiron process is modeled as an amine absorber
using MEA as the solvent. It removes CO2 from the recycled top gas stream to prevent
build-up of CO2 in the system and ensure that the bustle gas re-entering the shaft furnace
has sufficient partial pressure of reducing species. The amine solvent is regenerated in
a regeneration tower that consumes medium-pressure steam. The majority of the heat
required for regeneration is collected from the top gas in the heat recuperator. The operating
conditions and performance of the CO2 removal unit is identical in both the Energiron III
base case and PCL-DRI E case. In this work, the amine system was modeled using a solvent
concentration of 35 wt % MEA and a solvent loading of 0.25 mol CO2 /mol MEA. Top gas
enters the absorber at 1000 kPa(a) and 109 ◦ C. The regenerator operates at 180 kPa(a), with
separated CO2 exiting at 45 ◦ C.

1.2.6. Syngas Treatment

Syngas treatment varies depending on whether it is a Midrex-type process or an
Energiron-type process. For Midrex-type processes, the syngas exiting the reformer requires
further heating to reach the desired injection temperature to the shaft furnace. This is
achieved by direct addition of oxygen to the syngas in a duct burner. Natural gas is also
added as a final adjustment to the reducing gas composition as it enters the shaft furnace.
Energies 2022, 15, x FOR PEER REVIEW 5 of 18

Energies 2022, 15, 5219 cesses). In the base DRI configurations, the portion of the top gas used as fuel in the re- 5 of 17

former firebox do not pass through this compressor, as it is already at sufficient pressure
for passage through the firebox and downstream heat exchangers.

(a)

(b)

Figure 2. Simplified process configurations for PCL-DRI processes following: (a) a Midrex-type
flowsheet (PCL-DRI—M) and (b) an Energiron III-type flowsheet (PCL-DRI—E). For ease of view-
ing, not all heat exchangers or heat integration are shown. Stream properties are defined in the
Supplementary Materials (Tables S3 and S4).
Energies 2022, 15, 5219 6 of 17

For Energiron-type processes, the syngas exiting the reformer is cooled and then
shifted (via a high-temperature water-gas shift reactor) to enrich its hydrogen concentration.
It is then further cooled to remove excess water, before being combined with the recycled
top gas exiting the CO2 absorber. The syngas must then be reheated to enter the shaft
furnace. This is done first in a fired heater, and then finished in a duct burner with the
addition of oxygen and natural gas.

1.3. PCL Configuration for DRI

Two novel configurations were explored in this work to provide a concept-level
assessment of how PCL can be applied to DRI processes to capture CO2 emissions. PCL
reactors have previously been investigated for other applications in which industrial heat
is needed to produce electricity, steam, hydrogen, and syngas [17]. In the case of DRI, the
PCL reactors serve as a replacement for the reformer fireboxes and process gas heaters in
the typical Midrex and Energiron process configurations, thereby capturing the flue gas
emissions that would result from those fossil fuel-fired units. The high-pressure flue gas
exiting the fuel reactor allows for latent heat recovery, while the O2 -depleted air stream
from the air reactor (which shall henceforth be referred to as “vitiated air”) can be directed
to a turbine to offset the power requirement for the main air compressor. Midrex-type and
Energiron III-type PCL process configurations are shown in Figure 2a,b, respectively. New
or modified subunits for PCL-DRI are described in the sections below.

1.3.1. Air and Fuel Reactors

The air and fuel reactors comprise the core of the PCL process, whereby the heat
supplied to the reformer and to heat process gas streams is produced. Ilmenite ((Fe, Ti)2 O3 )
was used as the oxygen carrier due to its low toxicity, availability, and low cost [22,23],
though numerous other oxygen carriers can be considered [24]. The fuel reactor is fired
with top gas from the DRI shaft furnace, containing CH4 , CO, and H2 as combustible
components. Natural gas is added as supplemental fuel, as needed. Equations (9)–(14)
show the reactions that occur in the fuel reactor between ilmenite and the fuel, while
Equations (15) and (16) show the reactions between ilmenite and air in the air reactor [17].
◦
3Fe2 O3 +CO → 2Fe3 O4 +CO2 ∆Hr = −40.4 kJ/mol CO (9)
◦
3Fe2 O3 +H2 → 2Fe3 O4 +H2 O ∆Hr = 0.7 kJ/mol H2 (10)
◦
Fe2 TiO5 +CO + TiO2 → 2FeTiO3 +CO2 ∆Hr = −65.9 kJ/mol CO (11)
◦
Fe2 TiO5 +H2 +TiO2 → 2FeTiO3 +H2 O ∆Hr = −24.6 kJ/mol H2 (12)
◦
12Fe2 O3 +CH4 → 8Fe3 O4 +CO2 + 2H2 O ∆Hr = 170.0 kJ/mol CH4 (13)
◦
4Fe2 TiO5 +CH4 +4TiO2 → 8FeTiO3 +CO2 +2H2 O ∆Hr = 66.0 kJ/mol CH4 (14)
◦
4FeTiO3 +O2 → 2Fe2 TiO5 +2TiO2 ∆Hr = −434.0 kJ/mol O2 (15)
◦
4Fe3 O4 +O2 → 6Fe2 O3 ∆Hr = −486.0 kJ/mol O2 (16)
Due to the nature of the reactants, the reactions in both the air and fuel reactor are net
exothermic, though the majority of the heat is released in the air reactor and is removed
both in-bed and in the freeboard. The air and fuel reactor bed temperatures were fixed at a
maximum of 950 ◦ C to avoid the use of exotic metals for internal components. The preheat
temperature of the fuel reactor was controlled such that the small exothermic nature of the
reactions was consumed in sensible heating of the reactants, thus requiring no net heat
removal from the fuel reactor. The operating pressure of both reactors was set to 800 kPa(a)
based upon the optimum determined in previous work [25].
Energies 2022, 15, 5219 7 of 17

1.3.2. Reformer
For both PCL-DRI cases, the air reactor of the PCL unit provides the heat of reforming
to the catalyst-filled tubes, acting as a replacement to a traditional reformer firebox. This
offers many advantages, including reduced thermal stresses on the reformer tubes due to a
lower temperature difference between the reformed gas and air reactor, lower mechanical
stress due to a smaller pressure differential between the inside and outside of the reformer
tubes (for the Energiron-type configuration), and lower external tube temperature (due to
improved heat transfer). These factors allow for the selection of less costly tube materials.

1.3.3. Air Compressors and Power Recovery

The compressor supplying the air reactor has two stages. The first stage is directly
coupled to an expansion turbine, and is driven by the expansion of the vitiated air stream
exiting the air reactor. The second stage is electrically driven. The air discharge pressure
from the second stage is 971 kPa(a). Previous work has investigated options for power
recovery in PCL, including simple and high-temperature turbo expanders, incorporation of
a duct burner for additional power recovery, or omission of power recovery [17].

1.3.4. CO2 Processing

The CO2 processing requirements will be site- and application-specific. Depending on
the location of the facility, the captured CO2 may be compressed and transported as a gas
by pipeline, or may be converted to a liquid for transportation by rail, ship, or truck [26,27].
Here, we consider supercritical CO2 transported by pipeline. Prior to compression, heat is
recovered from the flue gas using various heat recovery heat exchangers. The gas is then
further cooled in a direct contact cooler, to condition it for entry to the CO2 compression
train. CO2 compression occurs in five stages with intercooling. The first stage has a
compression ratio of 4, and is followed by a dryer to remove moisture down to the pipeline
specification. The remaining stages are equally sized with compression ratios of 1.44, to
arrive at a final product pressure of 12,060 kPa(a), which is within the specified range for
the Alberta Carbon Trunk Line [28].
The required purity of the CO2 is dependent upon the transportation pipeline specifi-
cations and on the final destination of the CO2 (storage or utilization). For this comparative
concept-level study, no additional purification is assumed to be required. Options for re-
moving combustibles and inerts have been reviewed in literature and include pure oxygen
injection, cryogenic fractionation, use of solvents, and/or membranes [29,30].

1.3.5. Top Gas Compressor

For PCL-DRI, the top gas compressor is larger than the base case, since all top gas
is directed through it, rather than just the fraction that is recycled as reducing gas. The
discharge pressure is maintained the same as the base case.

1.3.6. Fuel Gas Compressor (Midrex-Type Configuration)

The Midrex reformer operates at a lower pressure than the required fuel gas supply
pressure to the PCL units. Thus, for the PCL-DRI M process, the fuel portion of the top
gas is sequentially pressurized to 971 kPa(a) in a second compressor following the top gas
compressor (the fuel gas compressor) to meet the supply pressure requirements of the fuel
reactor.

1.3.7. Syngas Treatment

All aspects of the syngas treatment are the same except for one substitution: the
heating that would have been done in a fired heater in the Energiron base case is now done
inside the air reactor of the PCL unit. The heating duty and outlet temperature is kept
constant between the two cases.
Energies 2022, 15, 5219 8 of 17

2. Materials and Methods

2.1. Process Simulation Methodology
All four configurations were simulated in Aspen HYSYS V11 (HYSYS). Since this was
a preliminary study, a detailed shaft furnace model was not developed. Instead, the top gas
exiting the top of the shaft furnace was treated as the inlet boundary of the models, and the
reducing gas exiting the duct burner was treated as the outlet boundary. The conditions
and compositions of these two streams were informed by literature. For the Midrex base
case (Base—M), data for the Contrecoeur plant presented by Hamadeh was used [12]. For
the Energiron III base case (Base—E), the data was taken from Zugliano et al. [31]. The
former produces a cold DRI product, while the latter produces hot DRI. PCL-DRI M was
built up using the Midrex shaft furnace data, while PCL-DRI E used the Energiron III data.
Before using either set of data, the mass and energy balances around each shaft furnace
were verified in independent models in HYSYS. These models consisted of simplified
conversion reactors and heat exchangers, applying the reactions identified in Section 1.2.1.
As is discussed further in Section 3.1.1, the mass balance on the Midrex data did not close
and required minor adjustments to the top gas stream. This corrected top gas, scaled to the
desired DRI production rate, was used as the input to the models for Base—M and PCL-
DRI—M. There were no errors identified in the Energiron data, and so the literature top
gas composition and flow, after scaling up to the target operating flow, was used directly as
the input to Base—E and PCL-DRI—E. The operating conditions of the remaining process
units (such as the top gas scrubber and the reformer) and required flows of raw materials
were adjusted until the composition and mass flow of the reducing gas exiting the duct
burner of the models were in agreement with the literature values for the reducing gas
entering the shaft furnace.

2.2. Case Descriptions

There are a total of five process configurations that are considered in this work, as
summarized in Table 1. There are two base cases—one for a Midrex process (Base—M)
and one for an Energiron III process (Base—E), as shown in Figure 1. These cases assume
that all CO2 is vented to the atmosphere. There is limited public operating data for DRI
facilities; other more optimal operating points may exist. For the purpose of comparison to
PCL-DRI, the available data are sufficient, as the performance of the shaft furnace is kept
constant for all Midrex-type cases and for all Energiron-type cases.

Table 1. Summary of cases simulated in HYSYS.

Case Base—M Base—E Base—E + Comp PCL-DRI—M PCL-DRI—E

DRI process type Midrex Energiron Energiron Midrex Energiron
PCL employed No No No Yes Yes
CO2 compression train No No Yes Yes Yes
CO2 fate Vented Vented Partial capture Full capture Full capture

There are two PCL-DRI configurations, as shown in Figure 1. PCL-DRI M is the

Midrex-type case, while PCL-DRI E is the Energiron-type case. Both of these cases employ
a CO2 compression train to condition the captured CO2 for pipeline transportation.
Energiron DRI facilities include a CO2 absorber in their standard flowsheet. Case
“Base—E + Comp” explores a partial CO2 capture scenario in which the CO2 stream from
the absorber is directed to a CO2 compression train for transportation to a sequestration or
utilization site.

2.3. Fluid Package and Thermodynamic Data

The Peng–Robinson fluid package was used for all streams except for the modeling
of the CO2 absorber in Energiron-type configurations. For those streams, the Acid Gas—
Chemical Solvents package was used. Peng–Robinson is an equation of state model and is
Energies 2022, 15, 5219 9 of 17

the best optimized model in HYSYS for gases at high temperature. For the reactions with
the solid iron ore and ilmenite ore species, hypothetical compounds were created in HYSYS
using thermodynamic data from FactSage [17,32].

2.4. Battery Limits

Inputs to the process simulations are ilmenite ore (the oxygen carrier in the PCL
reactors), oxygen, natural gas, air, and top gas from the shaft furnace. Properties of these
streams are provided in Table 2. Oxygen is assumed to be provided as a utility at the
pressure required for the process. Iron ore is an input to the DRI shaft furnace; however,
since the furnace is not modeled directly, the ore is not considered here.

Table 2. Properties of reactants in base case and PCL-DRI process simulations.

Natural Gas Ilmenite a Oxygen Air Top Gas—Midrex Top Gas—Energiron

Temperature (◦ C) 25 25 25 25 285 412
Pressure (kPa(a)) 400 101 300|750 b 101 142 661
Species (mol frac)
CH4 0.9594 0 0 0 0.0300 0.0586
H2 O 0 0 0 0 0.2000 0.2847
CO 0 0 0 0 0.1900 0.1391
CO2 0.0073 0 0 0 0.1600 0.0728
O2 0 0 0.9000 0.2100 0 0
H2 0 0 0 0 0.4000 0.4332
N2 0.0169 0 0.1000 0.7900 0 0.0117
C2 H6 0.0160 0 0 0 0 0
C3 H8 0.0004 0 0 0 0 0
TiO2 0 0.4124 0 0 0 0
Fe2 TiO5 0 0.2918 0 0 0 0
Fe2 O3 0 0.2958 0 0 0 0
a PCL-DRI cases only. b A|B are for A = Midrex-type process and B = Energiron-type process.

Outputs at the simulation boundaries common to all cases are the reducing gas and
produced condensate. From the base case simulations, additional product streams are the
flue gas from the reformer/fired gas heater and purge CO2 from the absorber (Base—E
only). For the PCL-DRI cases, these streams are replaced by the compressed CO2 product
and the vitiated air stream from the air reactor. Spent ilmenite is the final product from
the PCL-DRI process. The properties of these streams are summarized in Table 3, while
case-specific stream information can be found in the Supplementary Materials.

Table 3. Properties of products in base case and PCL-DRI process simulations.

Reducing CO2 from Vitiated Compressed

Condensate Flue Gas a Ilmenite c
Gas Absorber b Air c CO2 Product d
Temperature (◦ C) 957–962 50–67 106–351 45 191–425 17–170 49
Pressure (kPa(a)) 200–725 122–621 101–121 180 800 101 12 060
Species (mol frac)
CH4 0.083–0.091 0 0 0.001 0 0 0.015
H2 O 0.030–0.043 0.999–1.000 0.162–0.216 0.054 0 0 0
CO 0.169–0.320 0 0 0 0 0 0
CO2 0.022–0.050 0–0.001 0.091–0.111 0.943 0 0 0.935
O2 0 0 0.009–0.070 0 0 0.013 0
H2 0.509–0.662 0 0 0.002 0 0 0
N2 0.013–0.016 0 0.658–0.718 0 0 0.987 0.050
Energies 2022, 15, 5219 10 of 17

Table 3. Cont.

Reducing CO2 from Vitiated Compressed

Condensate Flue Gas a Ilmenite c
Gas Absorber b Air c CO2 Product d
C2 H6 0 0 0 0 0 0 0
C3 H8 0 0 0 0 0 0 0
TiO2 0 0 0 0 0.412 0 0
Fe2 TiO5 0 0 0 0 0.292 0 0
Fe2 O3 0 0 0 0 0.296 0 0
aBase—M, Base—E, and Base—E + comp cases only. b Base—E + comp case only. c PCL-DRI cases only. d

PCL-DRI and Base—E + comp cases only.

2.5. Calculation of Thermal Efficiency

The thermal efficiency of the process is calculated as the percent of useful heat input
to the process compared to the total heat input to the process, as shown in Equation (17).

useful heat (total heat input) − (waste heat)

Thermal efficiency (%) = ∗ 100 %= ∗ 100 % (17)
total heat input total heat input
The total heat input is calculated as the summation of the electric power for all rotating
equipment, the higher heating value multiplied by flow rate of all reactant streams, any
hot utility stream duties, and the sensible heat of all reactants above a reference condition
of 25 ◦ C and 101 kPa(a). Waste heat was calculated as the summation of all cooling utility
duties, and the sensible heat of waste streams (e.g., condensate, flue gas that is vented,
spent oxygen carrier) above a reference condition of 25 ◦ C and 101 kPa(a).

3. Results
3.1. Base Case Agreement to Literature Values
3.1.1. Base—M
A mass balance on the literature data provided in [12] for the shaft furnace in the
Contrecoeur facility showed that the reported flow of top gas exiting the shaft furnace was
incorrect. To achieve a correct mass balance, the mass flow of top gas was increased by 7%
relative to the literature data and the composition was slightly adjusted. This discrepancy is
believed to be a result of errors in measurement of the high-temperature process gas stream.
Note that this additional mass flow all ends up in the excess purge stream (stream 14)
and thus has no impact on the idealized process performance. A comparison between the
reducing gas composition generated by the model and the target literature value shows
good agreement (Table 4).

Table 4. Comparison of Base—M model boundaries at shaft furnace to literature.

Top Gas from Shaft Furnace Reducing Gas to Shaft Furnace

Literature Model Literature Model
Temperature (◦ C) 285 285 957 957
Pressure (kPa(a)) 142 142 n.d. 200
Mass flow (t/h) 299.8 1 319.8 229.5 1 229.5
Species (mol frac)
CH4 0.0295 0.3000 0.0908 0.0909
H2 O 0.1903 0.2000 0.0428 0.0426
CO 0.1958 0.1900 0.3271 0.3197
CO2 0.1709 0.1600 0.0240 0.0216
H2 0.4028 0.4000 0.4966 0.5094
N2 0.0102 0.0200 0.0176 0.0157
1 Scaled to 2 MTPA DRI production rate.
Energies 2022, 15, 5219 11 of 17

While no operating data for the balance of the plant is provided in [12], the operating
conditions of the reformer in Base—M can be compared to typical ranges that would be
expected for a Midrex reformer, where dry reforming reactions dominate. Table 5 shows
that all model parameters fall within the expected range except for the steam to carbon
ratio. In the model, this ratio is lower than the expected value, however, it falls within the
same magnitude and results in dry reforming. It is not possible to increase the steam to
carbon ratio in the model without exceeding the specified steam content for the reducing
gas in Table 4. Overall, the model provides a good representation of the Midrex reformer,
and given that the same conditions are used in PCL-DRI—M, will not impact differential
performance observations between Base—M and PCL-DRI—M.

Table 5. Comparison of Base—M reformer operating conditions to expected ranges.

Operating Condition Base—M Typical Midrex Reformer 1

Reformer feed pre-heat temperature (◦ C) 500 400–500
Reformed gas outlet temperature (◦ C) 925 925
Reformed gas outlet pressure (kPa(a)) 220 200–300
CO2 /Carbon ratio 0.80 0.80
Steam/Carbon ratio 0.48 0.65
1 Typical values as reported in [9].

3.1.2. Base—E
Verification of the mass balance around the shaft furnace did not show any major
discrepancies for the data provided in Zulgiano et al. [31], thus no adjustments to the
model’s input top gas composition were required. The reducing gas composition calculated
by the model for the Base—E case shows good agreement with the literature values, as
shown in Table 6.

Table 6. Comparison of Base—E model boundaries at shaft furnace to literature.

Top Gas from Shaft Furnace Reducing Gas to Shaft Furnace

Literature Model Literature Model
Temperature (◦ C) 412 412 962 962
Pressure (kPa(a)) 661 661 725 725
Mass flow (t/h) 293.71 293.7 193.0 1 192.7
Species (mol frac)
CH4 0.0586 0.0586 0.0830 0.0831
H2 O 0.2881 0.2881 0.0296 0.0296
CO 0.136 0.136 0.1620 0.1691
CO2 0.0728 0.0728 0.0498 0.0502
H2 0.4332 0.4332 0.6629 0.6623
N2 0.0113 0.0113 0.0127 0.0129
1 Scaled to 2 MTPA DRI production rate.

Table 7 shows the reformer operating conditions for Base—E in comparison to a typical
SMR. The reformed gas outlet temperature and the steam to carbon ratio in the feed are both
within expected ranges. The steam to carbon ratio was set near the minimum of the range
because, unlike for H2 production, DRI requires a larger yield of CO to act as a reductant
in the shaft furnace. SMRs are typically operated at high pressure to limit downstream
compression requirements at the hydrogen production facility, however, the reactions are
less thermodynamically favorable as the pressure increases [33,34]. Given the Energiron
shaft furnace operates at 600–800 kPa [12], there is no need to increase the reformer pressure
beyond what is required as allowance for pressure drop across downstream processing
units between the reformer and the shaft furnace.
Energies 2022, 15, 5219 12 of 17

Table 7. Comparison of Base—E reformer operating conditions to typical steam methane reforming.

Operating Condition Base—E Typical SMR

Reformed gas outlet temperature (◦ C) 880 800–900 1
Reformed gas outlet pressure (kPa(a)) 945 1500–3000 1
Steam/Carbon ratio 1.6 1.5–5 2
1 As described in [35]. 2 As described in [34–36].

3.2. Comparison of Process Performance

The PCL-DRI cases are equivalent or superior to the base cases in many aspects of DRI
process performance as shown in Table 8. First, comparing Base—M to PCL-DRI—M, the
overall natural gas consumption of the two processes, which serves only to supplement the
feed gas to the reformer tubes and provide direct contact heating of the reducing gas in the
duct burner, is identical. This is because the chemistry and the conditions of the recycle gas
loop are maintained constant between both cases. Likewise, the oxygen consumption is
identical. The required fuel consumption (in the form of top gas) to the fuel reactor of the
PCL unit compared to the reformer firebox is less in the PCL-DRI—M case due to the lower
operating temperature of the PCL unit, thus requiring less sensible heating of the gases.
Since heat losses in the recycled gas loop are not considered in these idealized cases, there
is an excess top gas stream that is required neither as a fuel for heating, nor as a recycle
feed to the reformer. For the purpose of these simulations, this excess gas is directed to a
purge stream and it is excluded from emissions considerations or its impact on the thermal
efficiency of the plant in this work. When factoring in practical operating considerations,
this stream would not exist, because all of the top gas would be required to be combusted
as fuel in the Midrex firebox to account for heat losses, with further supplementation of
natural gas. To avoid introduction of additional uncertainty through the estimate of heat
losses for the PCL-DRI and Base—E configurations, no heat losses in the recycle gas and
reforming loop are considered for any of the cases. Thus, fuel consumption and CO2
production rates represent the minimum possible for these configurations.

Table 8. Comparison of key performance indicators between base cases and PCL-DRI cases.

Performance Base— PCL-DRI— Base—E +

Unit Base—E PCL-DRI—E
Indicator M M Comp
Total natural gas consumption GJ/t DRI 11.9 11.9 10.7 10.7 9.80
Fuel consumption for heating
GJ/t DRI 4.64 4.26 4.54 4.54 3.65
(reforming + fired heaters/fuel reactor)
Oxygen consumption Nm3 /t DRI 27.7 27.7 16.4 16.4 16.4
Water produced t H2 O/t DRI 0.135 0.347 0.231 0.236 0.397
Nitrogen production Nm3 /t DRI 0 714 0 0 692
Utility cooling duty MWh/t DRI 0.243 0.258|0.542 1 0.686 0.721 0.660|0.723 1
Net power import MWhe /t DRI 0.055 0.114|0.148 1 0.056 0.082 0.085|0.130
Thermal efficiency % 91.7 91.9 92.0 91.6 90.9
CO2 produced t CO2 /t DRI 0.393 0.354 0.450 0.450 0.401
% CO2 captured % 0 100 0 47 100
% CO2 emitted % 100 0 100 53 0
1Values A|B are: A = excluding contribution of DCC + CO2 compression; B = including contribution of DCC +
CO2 compression.

Comparing the fuel consumption for Base—E and PCL-DRI—E, again, there is a
reduction in the fuel consumption for the PCL fuel reactor due to its lower operating
temperature. In these configurations, all of the available top gas plus additional natural gas
is required to fulfill the heating requirement of the PCL unit/firebox; thus, this translates
directly to an 8% reduction in the battery limit natural gas requirement for the process.
The PCL-DRI cases have potential value-added by-products. Process water is pro-
duced at a rate of 2.6 times greater for the Midrex-type cases and 1.7 times greater for the
Energies 2022, 15, 5219 13 of 17

Energiron-type cases. This water is recovered from the fuel reactor flue gas during heat
recovery/cooling, compression, and drying. This additional water stream can be evaluated
for use as boiler feed water, for cooling and scale removal in the hot rolling mill, or for
other applications at the iron and steel facility [37]. Additionally, there is production of a
nitrogen stream from the air reactor, which may also be considered for cooling in the hot
rolling mill to offset some of the water requirement.
CO2 capture does have additional costs, which become evident when examining
the increased power import and cooling duty of the PCL-DRI cases relative to the base
cases. The net power imports of the PCL-DRI processes are more than double the power
requirements for their respective base cases, despite the offset of power requirement
achieved through the use of an expansion turbine on the nitrogen stream produced from
the air reactor. Power consumption of PCL-DRI—M is greater than PCL-DRI—E due to
the greater compression requirement for the top gas used as fuel in the fuel reactor. The
Midrex top gas is near atmospheric pressure when it exits the shaft furnace, whereas the
Energiron top gas is already near the operating pressure of the PCL units.
The two major draws for power in the PCL-DRI process are (1) the air compressors
supplying air to the air reactor and (2) the CO2 compression train that brings the captured
CO2 to pipeline pressure. To understand the impact of these two main power consumers,
the power consumption for PCL-DRI is presented both including and excluding CO2
conditioning and compression. A partial capture case (Base—E + comp) also shows the
impact of adding a smaller CO2 compression unit to the CO2 stream that is separated
from the CO2 absorber that is inherently a part of Energiron-type DRI processes. CO2
compression represents 23 and 34% of the total power import for PCL-DRI—M and PCL-
DRI—E respectively, while it adds 46% for partial CO2 capture from Base—E. With the
push toward electrification and toward reducing the carbon intensity of the power grid,
the increased power import of the PCL-DRI processes does not have to be considered as
a negative impact; it is merely a shift in the type of energy that the process consumes. A
detailed techno-economic assessment (TEA) is required to assess how this shift may affect
operating costs of the DRI facility, as well as how the cost of carbon capture compares to
competing CO2 capture technologies. Further optimization can be performed to minimize
power import, if desired.
Considering the cooling duty of the PCL-DRI processes relative to their respective
base cases, the cooling duty excluding CO2 conditioning and compression is quite sim-
ilar. The addition of conditioning in a DCC, drying, and intercooling between the CO2
compressor stages results in a 123% increase in the total process cooling duty for PCL-
DRI—M and a 5% increase for PCL-DRI—E. The relatively small change for PCL-DRI—E
results from the much higher cooling duty already required for Energiron-type processes,
both in the condenser of the CO2 absorber and in the cooling of the syngas exiting the
reformer for water knock-out, prior to re-heating the reducing gas for entry to the shaft
furnace. Even after heat recovery from the flue gas, the stream is still relatively hot
(528 ◦ C for PCL-DRI—M and 130 ◦ C for PCL-DRI—E) and contains a significant amount
of waste heat that must be cooled before compression can occur.
Chemical looping has often been described as a carbon capture technology with little
to no energy penalty [38]. This is observed in the comparison of thermal efficiencies of
PCL-DRI—M to Base—M, in which there is no loss in efficiency with integration of PCL. For
PCL-DRI—E and Base—E + Comp, there is a small drop in efficiency that can be attributed to
the increase in power and cooling duty for the CO2 compression units, though this effect is
much smaller than would be the case if amine capture were applied to all flue gas streams.

4. Discussion
The Base—M and Base—E cases were developed to have a reasonable point of compar-
ison for PCL-DRI to current industrial standards; however, each of these models was built
based on a single operating point and may not be representative of all Midrex or Energiron-
III processes. If more industrial data were available at alternative operating points, or if a
Energies 2022, 15, 5219 14 of 17

detailed shaft furnace model could be developed and integrated with the current models,
future work could focus on broadening the understanding of PCL-DRI process performance
under a range of common operating conditions. To evaluate the accuracy of the developed
models, they were compared to emissions data in literature. The Midrex process typically
produces 0.65 t CO2 /t DRI, while Energiron III produces 0.56 t CO2 /t DRI [7]. In this work,
the idealized Base—M configuration emits 0.39 t CO2 /t DRI, however, if one considers the
combustion of the excess top gas as fuel in the reformer to make up for heat losses, as
would be done in a non-idealized plant, this value rises to 0.52 t CO2 /t DRI. The Base—E
configuration likewise undershoots the expected value, at 0.45 t CO2 /t DRI, but again, the
difference would likely be made up if heat losses were considered.
The performance of the PCL-DRI processes are competitive with traditional DRI
processes, with the added benefit of full CO2 capture. There was also no optimization of the
top gas recycle loop for PCL-DRI; it is possible that superior performance could be achieved
if the overall process pressure and reformer feed conditions could be modified to better suit
PCL. For example, the shaft furnace in PCL-DRI—E could be operated such that the top gas
is at a pressure sufficient to supply the fuel reactor directly, without requiring an additional
fuel gas compressor. For the PCL-DRI—M configuration, options could be explored to
reduce the natural gas feed to the reformer through the use of the additional top gas that
would be available due to the reduced fuel requirement in the fuel reactor. Optimization
activities such as these would require a detailed kinetic model of the shaft furnace in order
to predict performance with changing reducing gas compositions and pressures.
Despite its many advantages, the challenges and drawbacks associated with PCL-
DRI must also be considered. As has been stated in Section 3.2, the duty of the main air
compressor(s) significantly increase the power consumption of the process. Due to their
size, previous work has shown that the rotating equipment represents the largest portion
of the capital costs of the plant, equal to or more than the cost of all other equipment
combined [25]. Scale-up may also be a challenge, given the very high duty and solids
circulation rates that would be required to meet the needs of a typical iron and steel facility.
In this work, heat recovery from the air reactor approaches 200 MW, depending on the case,
and the required solids circulation rate is up to 1370 t/h (oxidized). It is anticipated that
two or more PCL units would be required to meet this demand, allowing scale-up to take a
modular approach. Choice of reactor type will likely be a key driver in the economics and
feasibility of such a process scale. Future work will examine reactor sizing and economics
for a novel design with no external solids transfer between the air and fuel reactors.
PCL-DRI is not the only way to eliminate CO2 emissions from a DRI facility. Post
combustion capture units, such as amine absorbers, can be employed, or alternative feed-
stocks such as low-carbon hydrogen or bio-methane can be used [39–42]. PCL has already
been shown to be more efficient than conventional carbon capture technologies for other
applications [17,24,43]. Midrex and Energiron, as well as several other green steelmaking
projects, have configurations under development that eliminate the reformer from the DRI
process and allow direct use of H2 as the reducing gas in the shaft furnace (purchased
over-the-fence or produced on site); however, this alters the heat balance in the shaft furnace
and requires additional heat input to make up the difference [6,39]. There is the additional
caveat of the alteration of the composition of the DRI product when only hydrogen is
used as the reductant. To aid in the efficiency of the downstream EAF, it is preferred to
have 1.5–3 wt % carbon in the DRI. This carbon is then combusted with oxygen in the EAF
to help melt the iron; thus, those CO2 emissions would have to be managed downstream [6].
In traditional DRI processes, the carbon in the DRI is incorporated in the shaft furnace via
carbon deposition from the thermal decomposition of methane and the inverse Boudouard
reaction [8]. Switching to DRI-H2 means that carbon cannot be deposited onto the DRI, and
will increase the energy consumption of the EAF. PCL-DRI does not have this drawback,
and will allow the EAF to operate without modification or loss of efficiency. Furthermore,
PLC-DRI does not depend on the existence of a local low-carbon hydrogen producer or
hydrogen distribution network, allowing net-zero DRI to be produced before the infras-
Energies 2022, 15, 5219 15 of 17

tructure for a hydrogen economy becomes available, as well as in remote regions where
low-carbon hydrogen might be inaccessible.
This work considers the use of ilmenite ore as the oxygen carrier for PCL. This could
be an attractive option for steelmakers, as there would be the potential for iron and titanium
recovery from the spent oxygen carrier, thus valorizing an otherwise waste-product at its
point of use. Iron ore can also be used directly as an oxygen carrier [23], which would
simplify the supply chain for the iron and steel facility, since oxygen carrier makeup could
be supplied from the iron ore already brought in to feed the shaft furnace. Many other
oxygen carriers bear consideration, both natural and synthetic, and typically contain some
combination of Cu, Mn, Fe, Co, and Ni species [24,44,45]. The choice of oxygen carrier
will impact the heat balance of the PCL units as well as the flue gas purity, based on the
reactivity of the oxygen carrier with the fuel. In the current work with ilmenite ore, the
fuel reactor is net adiabatic (incoming sensible heat and exothermic reactions with CO
and H2 counterbalance the endothermic reactions with CH4 ); however, with some oxygen
carriers, heat is released in both the air and fuel reactors. This would alter the location of
some of the heat recovery units in the PCL units, though at this stage of the analysis, it
would not have a significant impact on the overall process performance. Ilmenite ore has
a relatively low reactivity with CH4 compared to CO and H2 [22,46], and for this reason,
the flue gas is expected to contain a fraction of unconverted methane (in this work, we
assume 98.5% fuel conversion). Other oxygen carriers may have differing reactivities with
the fuel, differing selectivities for products (resulting in differing yields of CO2 ) or have
oxygen uncoupling properties, all of which may impact flue gas treatment requirements to
achieve purity specifications for transportation and storage.

5. Conclusions and Future Work

Process simulation of two novel DRI configurations show that PCL-DRI is a promising
option for net-zero DRI production. The PCL-DRI process has many added benefits over
traditional DRI configurations in addition to inherent CO2 capture, including reduced fuel
consumption, production of potentially valuable by-products, and, little or no reduction in
thermal efficiency. Unlike other methods of eliminating emissions from iron production,
there is no required change in the chemistry of the reducing gas or DRI product, allowing
for operational flexibility and simplicity of integration with downstream processes as
steelmakers transition toward net-zero steel production.
Future work will consider a full TEA-LCA for the PCL-DRI process, as well as investi-
gation of alternative oxygen carriers. Additional configurations can be assessed to optimize
the operating pressure of the PCL units for DRI and reduce the net power import of the
process.

Supplementary Materials: The following supporting information can be downloaded at: https:
//www.mdpi.com/article/10.3390/en15145219/s1, Table S1: Base—M stream properties, Table S2:
Base—E stream properties, Table S3: PCL-DRI—M stream properties, Table S4: PCL-DRI—E stream
properties, Figure S1: CO2 compression and drying process flow diagram.
Author Contributions: Conceptualization, N.B., R.S. and R.H.; methodology, N.B.; formal analysis,
N.B.; investigation, N.B.; data curation, N.B.; writing—original draft preparation, N.B.; writing—
review and editing, R.S. and R.H.; visualization, N.B. and R.H.; supervision, R.S. and R.H.; project
administration, R.H.; funding acquisition, R.H. All authors have read and agreed to the published
version of the manuscript.
Funding: This research was funded by the Program for Energy Research and Development (PERD)
at Natural Resources Canada, Government of Canada (CEO-19-115).
Institutional Review Board Statement: Not applicable.
Informed Consent Statement: Not applicable.
Data Availability Statement: Not applicable.
Energies 2022, 15, 5219 16 of 17

Acknowledgments: The authors would like to thank Hatch for their expert feedback on the base case
and PCL-DRI models.
Conflicts of Interest: The authors declare no conflict of interest.

References
1. Natural Resources Canada. Iron Ore Facts. Available online: https://www.nrcan.gc.ca/our-natural-resources/minerals-mining/
minerals-metals-facts/iron-ore-facts/20517 (accessed on 13 May 2020).
2. Smil, V. Still the Iron Age; Butterworth Heinemann (Elsevier): Oxford, UK, 2016; ISBN 978-0-12-804233-5.
3. Canadian Steel Producers Assocation. Canada’s Steel Industry: A Sustainable Choice. 2020, pp. 1–19. Available online:
https://canadiansteel.ca/files/resources/CSPA_Climate-Call-to-Action-EN.pdf (accessed on 14 April 2022).
4. International Energy Agency. Net Zero by 2050—A Roadmap for the Global Energy Sector; International Energy Agency: Paris, France,
2021; p. 13.
5. World Steel Association. Overview of the Steelmaking Process. Available online: https://www.worldsteel.org/en/dam/jcr:
177c8e5c-e02a-4e08-9dc6-cce7372b41c2/Overview+of+the+Steelmaking+Process_poster.pdf (accessed on 7 May 2020).
6. Chevrier, V. MIDREX H2TM: Ultra Low CO2 Ironmaking in the Transition to the Hydrogen Economy. Steel Times Int. 2020, 44,
29–34.
7. Baig, S. Cost Effectiveness Analysis of HYL and Midrex DRI Technologies for the Iron and Steel-Making Industry. Masters’ Thesis,
Duke University, Durham, NC, USA, 2016.
8. Hamadeh, H.; Mirgaux, O.; Patisson, F. Detailed Modeling of the Direct Reduction of Iron Ore in a Shaft Furnace. Materials 2018,
11, 1865. [CrossRef]
9. Farhadi, F.; Motemed Hashemi, M.Y.; Bahrami Babaheidari, M. Modelling and Simulation of Syngas Unit in Large Scale Direct
Reduction Plant. Ironmak. Steelmak. 2003, 30, 18–24. [CrossRef]
10. Energiron. ENERGIRON HYL: DRI Technology by Tenova and Danieli. 2014. Available online: https://www.energiron.com/
(accessed on 8 May 2020).
11. Midrex Technologies, Inc. The MIDREX®Process. 2018. Available online: https://www.midrex.com/wp-content/uploads/
MIdrex_Process_Brochure_4-12-18.pdf (accessed on 1 April 2022).
12. Hamadeh, H. Modélisation Mathématique Détaillée du Procédé de Réduction Directe du Minerai de Fer. Ph.D. Thesis, Université
de Lorraine, Metz, France, 2017.
13. Marx, K.; Bertsch, O.; Pröll, T.; Hofbauer, H. Next Scale Chemical Looping Combustion: Process Integration and Part Load
Investigations for a 10 MW Demonstration Unit. Energy Proc. 2013, 37, 635–644. [CrossRef]
14. Yazdanpanah, M.M.; Forret, A.; Gauthier, T.; Delebarre, A. An Experimental Investigation of Loop-Seal Operation in an
Interconnected Circulating Fluidized Bed System. Powder Technol. 2013, 237, 266–275. [CrossRef]
15. Zhu, L.; He, Y.; Li, L.; Wu, P. Tech-Economic Assessment of Second-Generation CCS: Chemical Looping Combustion. Energy 2018,
144, 915–927. [CrossRef]
16. De Guido, G.; Compagnoni, M.; Pellegrini, L.A.; Rossetti, I. Mature versus Emerging Technologies for CO2 Capture in Power
Plants: Key Open Issues in Post-Combustion Amine Scrubbing and in Chemical Looping Combustion. Front. Chem. Sci. Eng.
2018, 12, 315–325. [CrossRef]
17. Symonds, R.T.; Hughes, R.W.; Lu, D.Y.; Navarri, P.; Ashrafi, O. Systems Analysis of Pressurized Chemical Looping Combustion
for SAGD Applications. Int. J. Greenh. Gas Control 2018, 73, 111–123. [CrossRef]
18. Zhang, S.; Saha, C.; Yang, Y.; Bhattacharya, S.; Xiao, R. Use of Fe2 O3 -Containing Industrial Wastes As the Oxygen Carrier for
Chemical-Looping Combustion of Coal: Effects of Pressure and Cycles. Energy Fuels 2011, 25, 4357–4366. [CrossRef]
19. Primetals Technologies Austria GmbH. 2 MTPY MIDREX®Hot Briquette Iron Plant; Primetals Technologies Ltd.: London, UK,
2019.
20. Béchara, R.; Hamadeh, H.; Mirgaux, O.; Patisson, F. Optimization of the Iron Ore Direct Reduction Process through Multiscale
Process Modeling. Materials 2018, 11, 1094. [CrossRef]
21. Atsushi, M.; Uemura, H.; Sakaguchi, T. MIDREX Processes. Kobelco Technol. Rev. 2010, 29, 50–57.
22. Tan, Y.; Ridha, F.N.; Duchesne, M.A.; Lu, D.Y.; Hughes, R.W. Reduction Kinetics of Ilmenite Ore as an Oxygen Carrier for
Pressurized Chemical Looping Combustion of Methane. Energy Fuels 2017, 31, 7598–7605. [CrossRef]
23. Yu, Z.; Yang, Y.; Yang, S.; Zhang, Q.; Zhao, J.; Fang, Y.; Hao, X.; Guan, G. Iron-Based Oxygen Carriers in Chemical Looping
Conversions: A Review. Carbon Resour. Convers. 2019, 2, 23–34. [CrossRef]
24. Czakiert, T.; Krzywanski, J.; Zylka, A.; Nowak, W. Chemical Looping Combustion: A Brief Overview. Energies 2022, 15, 1563.
[CrossRef]
25. Cabello, A.; Hughes, R.W.; Symonds, R.T.; Champagne, S.; Lu, D.Y.; Mostafavi, E.; Mahinpey, N. Economic Analysis of Pressurized
Chemical Looping Combustion for Steam Assisted Gravity Drainage Applications. Int. J. Greenh. Gas Control 2019, 90, 102786.
[CrossRef]
26. Bui, M.; Adjiman, C.S.; Bardow, A.; Anthony, E.J.; Boston, A.; Brown, S.; Fennell, P.S.; Fuss, S.; Galindo, A.; Hackett, L.A.; et al.
Carbon Capture and Storage (CCS): The Way Forward. Energy Environ. Sci. 2018, 11, 1062–1176. [CrossRef]
Energies 2022, 15, 5219 17 of 17

27. IPCC. IPCC Special Report on Carbon Dioxide Capture and Storage; Intergovernmental Panel on Climate Change: Cambridge, UK,
2005; pp. 180–193.
28. Enhanced Energy Inc.; North West Redwater Partnership. ACTL Knowledge Sharing Report: Calendar Year 2016; North West
Redwater Partnership: Redwater, AB, Canada; Enhance Energy Inc.: Calgary, AB, Canada, 2017; pp. 70–73.
29. Lyngfelt, A.; Leckner, B. A 1000 MW Th Boiler for Chemical-Looping Combustion of Solid Fuels—Discussion of Design and
Costs. Appl. Energy 2015, 157, 475–487. [CrossRef]
30. Xu, G.; Liang, F.; Yang, Y.; Hu, Y.; Zhang, K.; Liu, W. An Improved CO2 Separation and Purification System Based on Cryogenic
Separation and Distillation Theory. Energies 2014, 7, 3484–3502. [CrossRef]
31. Zugliano, A.; Primavera, A.; Pignattone, D.; Martinis, A. Online Modelling of ENERGIRON Direct Reduction Shaft Furnaces.
IFAC Proc. Vol. 2013, 46, 346–351. [CrossRef]
32. Bale, C.W.; Bélisle, E.; Chartrand, P.; Decterov, S.A.; Eriksson, G.; Gheribi, A.E.; Hack, K.; Jung, I.-H.; Kang, Y.-B.; Melançon, J.;
et al. FactSage Thermochemical Software and Databases, 2010–2016. Calphad 2016, 54, 35–53. [CrossRef]
33. Pröll, T.; Lyngfelt, A. Steam Methane Reforming in Fluidized-Bed Heat Exchangers—A Case for Chemical-Looping Combustion.
In Proceedings of the 24th Fluidized Bed Conversion Conference (FBC24), Gothenberg, Sweden; 8–11 May 2022; FBC24: Gothenberg,
Sweden, 2022; pp. 1–10.
34. Pham Minh, D.; Siang, T.J.; Vo, D.-V.N.; Phan, T.S.; Ridart, C.; Nzihou, A.; Grouset, D. Chapter 4—Hydrogen production from
biogas reforming: An overview of steam reforming, dry reforming, dual reforming, and tri-reforming of methane. In Hydrogen
Supply Chains; Azzaro-Pantel, C., Ed.; Academic Press: Cambridge, MA, USA, 2018; pp. 111–166. ISBN 978-0-12-811197-0.
35. García, L. Hydrogen production by steam reforming of natural gas and other nonrenewable feedstocks. In Compendium of
Hydrogen Energy; Elsevier: Amsterdam, The Netherlands, 2015; pp. 83–107. ISBN 978-1-78242-361-4.
36. Fahim, M.A.; Alsahhaf, T.A.; Elkilani, A. Chapter 11—Hydrogen production. In Fundamentals of Petroleum Refining; Fahim, M.A.,
Alsahhaf, T.A., Elkilani, A., Eds.; Elsevier: Amsterdam, The Netherlands, 2010; pp. 285–302. ISBN 978-0-444-52785-1.
37. Walling, F.B.; Otts, L.E. Water Requirements of the Iron and Steel Industry; Water Requirements of Selected Industries; United States
Department of the Interior: Washington, DC, USA, 1967; pp. 341–394.
38. Khan, M.N.; Chiesa, P.; Cloete, S.; Amini, S. Integration of Chemical Looping Combustion for Cost-Effective CO2 Capture from
State-of-the-Art Natural Gas Combined Cycles. Energy Convers. Manag. X 2020, 7, 100044. [CrossRef]
39. Wang, R.R.; Zhao, Y.Q.; Babich, A.; Senk, D.; Fan, X.Y. Hydrogen Direct Reduction (H-DR) in Steel Industry—An Overview of
Challenges and Opportunities. J. Clean. Prod. 2021, 329, 129797. [CrossRef]
40. Quader, M.A.; Ahmed, S.; Ghazilla, R.A.R.; Ahmed, S.; Dahari, M. A Comprehensive Review on Energy Efficient CO2 Break-
through Technologies for Sustainable Green Iron and Steel Manufacturing. Renew. Sustain. Energy Rev. 2015, 50, 594–614.
[CrossRef]
41. Bhaskar, A.; Assadi, M.; Nikpey Somehsaraei, H. Decarbonization of the Iron and Steel Industry with Direct Reduction of Iron
Ore with Green Hydrogen. Energies 2020, 13, 758. [CrossRef]
42. Guo, D.; Li, Y.; Cui, B.; Chen, Z.; Luo, S.; Xiao, B.; Zhu, H.; Hu, M. Direct Reduction of Iron Ore/Biomass Composite Pellets Using
Simulated Biomass-Derived Syngas: Experimental Analysis and Kinetic Modelling. Chem. Eng. J. 2017, 327, 822–830. [CrossRef]
43. Mantripragada, H.C.; Rubin, E.S. Chemical Looping for Pre-Combustion and Post-Combustion CO2 Capture. Energy Proc. 2017,
114, 6403–6410. [CrossRef]
44. Siriwardane, R.; Riley, J.; Benincosa, W.; Bayham, S.; Bobek, M.; Straub, D.; Weber, J. Development of CuFeMnAlO4 +δ Oxygen
Carrier with High Attrition Resistance and 50-KWth Methane/Air Chemical Looping Combustion Tests. Appl. Energy 2021, 286,
116507. [CrossRef]
45. Riley, J.; Siriwardane, R.; Tian, H.; Benincosa, W.; Poston, J. Particle Scale Modeling of CuFeAlO4 during Reduction with CO in
Chemical Looping Applications. Appl. Energy 2019, 251, 113178. [CrossRef]
46. Bidwe, A.R.; Mayer, F.; Hawthorne, C.; Charitos, A.; Schuster, A.; Scheffknecht, G. Use of Ilmenite as an Oxygen Carrier in
Chemical Looping Combustion-Batch and Continuous Dual Fluidized Bed Investigation. Energy Proc. 2011, 4, 433–440. [CrossRef]