Simulation of the Sour-Compression Unit (SCU) process for CO2 purification applied to flue gases coming from oxy-combustion cement industries

Accepted Manuscript

Simulation of the Sour-Compression Unit (SCU) process for CO2 purification applied to flue gases coming from oxy-combustion cement industries Sinda Laribi , Lionel Dubois , Marie-Eve Duprez , Guy De Weireld , Diane Thomas PII: DOI: Reference:

S0098-1354(18)30492-7 https://doi.org/10.1016/j.compchemeng.2018.11.010 CACE 6277

To appear in:

Computers and Chemical Engineering

Received date: Revised date: Accepted date:

18 May 2018 4 October 2018 11 November 2018

Please cite this article as: Sinda Laribi , Lionel Dubois , Marie-Eve Duprez , Guy De Weireld , Diane Thomas , Simulation of the Sour-Compression Unit (SCU) process for CO2 purification applied to flue gases coming from oxy-combustion cement industries, Computers and Chemical Engineering (2018), doi: https://doi.org/10.1016/j.compchemeng.2018.11.010

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Highlights  A CO2 de-SOx and de-NOx process was simulated for full oxy-fuel cement plants.  A comprehensive chemical mechanism has been elucidated and implemented in Aspen.  Results of a parametric study were presented for a single-column configuration.  Optimizations depending on purity targets were achieved by economic evaluations.

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Article submitted for publication to:

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Simulation of the Sour-Compression Unit (SCU) process for CO2 purification applied to flue gases coming from oxy-combustion cement industries Sinda Laribi, Lionel Dubois, Marie-Eve Duprez, Guy De Weireld and Diane Thomas1 Faculty of Engineering, University of Mons, 20 Place du Parc, 7000 Mons, Belgium

Abstract

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Keywords Full oxy-fuel combustion; CO2 purification ; Cement industry; De-SOx; De-NOx;

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The purpose of the present work was to investigate, for a gas issued from a full oxy-fuel combustion in the cement industry, a CO2 de-SOx and de-NOx process, called “Sour-Compression Unit” (SCU), thanks to simulations with Aspen PlusTM. An important stage necessary for the SCU modeling has been the construction of an accurate chemical mechanism. Two-column and single-column configurations have been evaluated and compared. A parametric study and a Design Of Experiments have been conducted on a single-column process to study the influence of the operating parameters on the SOx and NOx abatement ratios. As a demonstration of the effectiveness of the model, three SOx and NOx purity specifications (depending on the further applications of the CO2) were applied to the purified gas. The feature of the investigated model lies on optimizing the way to reach the purity target in order to decrease installation costs (CAPEX) and energy requirements (OPEX).

Nomenclature Ai ai aij APU ASU BAT 1

Abatement ratio of component i (%) Coefficient of the centered variable relative to factor i for SOx abatement Coefficient of the scaled variable relative to factors i and j for SOx abatement Acid Purification Unit Air Separation Unit Best Available Techniques

Corresponding author. Tel : +32 (0) 65/37.44.04; Fax : +32(0)65/37.44.07; E-mail address: [email protected]

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Coefficient of the centered variable relative to factor i for NOx abatement Coefficient of the scaled variable relative to factors i and j for NOx abatement Chilled Ammonia Process Capital Expenditure Carbon Capture Utilization or Storage Concentration of component i in gas or liquid phase (kmol/m³) CO2 Purification Unit Gibbs standard free energy (kJ/kmol) Design Of Experiments Direct Separation calcining technology Activation energy (kJ/kmol) European Cement Research Academy Enhanced Oil Recovery Hydroxylamine N,N-disulfonate Hydroxylamine N,N-disulfonic acid Hydroxylamine N-sulfonate Hydroxylamine N-sulfonic acid Column height (m) Relative to the inlet of the column Reaction kinetic constant Reaction equilibrium constant Low Emissions Intensity Lime and Cement Nitrososulfonate Nitrososulfonic acid Operating Expenditure Relative to the outlet of the column Mass liquid flow rate (t/h) Molar flow rate of component i (kmol/h) Ideal gas constant (m³ Pa/(mol K)) Reaction rate (kmol/(m³ s)) Central value of the original variable for factor i Original variable for factor i Sour Compression Unit Calculation step for factor i Temperature (K) Temperature Swing Adsorption Volatile Organic Compounds Centered (or coded) variable for factor i CO2 content in the gas phase (vol.%)

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bi bij CAP CAPEX CCUS Ci CPU ∆G° DOE DS EA ECRA EOR HADS HADSH HAMS HAMSH Hcol in Kcin Kequ LEILAC NSS NSSH OPEX out QL Qmol,i R r R0,i Ri SCU Stepi T TSA VOC Xi yCO2

1. Introduction Reduction of greenhouse gases emissions and especially CO2 emissions represents one of the major challenging environmental issues nowadays. Anthropic carbon dioxide emissions reach 32 GtCO2/year (International Energy Agency, 2017). CO2 emitted from the cement industry represents 5% of the global annual emissions and 30% of the annual industrial emissions (International Energy Agency, 2013). In cement processes, more than 60% of CO2 emission arise from the decarbonation reaction in the making of clinker (Syndicat français de l’industrie cimentière, 2011) and cannot be avoided. 3

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In addition to CO2, the combustion of fuels and the chemical reactions occurring in the manufacturing process can form SO2, NOx, HCl, heavy metals, VOC and dust in the exhaust gases of cement plants. Techniques with different levels of maturity are currently investigated for CO2 capture and purification. The application of Carbon Capture Utilization or Storage (CCUS) to power plants flue gases (CO2 content, yCO2, in the flue gas between 5% and 15%) has already been considered in many studies and applied at industrial scale (i.e. Petra Nova Carbon Capture Plant, Boundary Dam CCS plant (Stéphenne, 2014)). While the application of CO2 capture is more advanced for power plants, the field remains quite innovative and challenging for the cement industry. The first pilot-scale tests of post-combustion CO2 capture technologies applied to the cement industry were carried out at the Norcem Brevik Cement plant in Norway (Bjerge and Brevik, 2014). Additionally the CEMCAP CO2 Capture project that was set up by ECRA (Carrasco-Maldonado et al., 2016) investigates four CO2 capture technologies: oxy-fuel capture, Chilled Ammonia Process (CAP), membrane-assisted CO2 liquefaction and calcium looping. CEMCAP has advanced retrofittable CO2 capture technologies for the cement industry to TRL6 (Jordal et al., 2017). Recently, the LEILAC (Low Emissions Intensity Lime and Cement) project was launched at the Heidelbergcement’s Lixhe cement plant in Belgium. The main challenge of the LEILAC pilot unit is to assess a direct separation (DS) calcining technology. This process, provided by Calix (Hills et al., 2017), consists in keeping the process gases separate from the flue gas stream which allows to capture over 95% of the plant CO 2 emissions with reduced energy consumption and capital penalty. An innovative CO2 treatment technology investigated here is applied to flue gases deriving from cement production processes considering oxy-fuel combustion (Carrasco-Maldonado et al., 2016), namely conducting the combustion of the fuel using pure oxygen. The outcoming flue gas from this oxy-fuel cement plant contains high CO2 levels (between 70 and 90%), and a gas purification process is applied in this case to obtain a sufficiently purified CO2 to be valorized. Different CO2 purification technologies exist, at different levels of development. The Air Liquide CO2 Cryogenic Purification Unit taking part of the Callide Oxy-fuel power plant project (Australia) demonstrated a total removal of all toxic gaseous emissions comprising SOx, NOx, particulates and trace elements from the flue gas stream (Spero et al., 2014) . LINDE’s concept applied to oxy-fuel power plants is an alkaline wash unit (high pressure and low temperature ) applied to a real flue gas coming from an oxy-boiler installed in the CO2-pilot plant at Schwarze Pumpe (Santos, 2015). The CO2 Purification Unit simulated by Praxair (ECRA, 2012) and applied to cement flue gases consists of activated carbon beds as adsorbents/catalysts for the removal of SO2 and NO to respectively produce H2SO4 and HNO3. The CO2 purification process using water scrubbing under pressurized conditions and tested at pilot scale by Air Products at Schwarze Pumpe, showed quite interesting de-SOx and de-NOx performances for oxy-fuel power plants (Torrente-Murciano et al., 2011) (White et al., 2013a). In the present work, this treatment, consisting in a CO2 purification of the oxy-fuel cement plant flue gases from its SOx and NOx components, as illustrated in Figure 1, is specifically considered.

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Figure 1: Block diagram of an oxy-fuel cement plant including the CO2 de-SOx/de-NOx process

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The CO2 Purification Unit (CPU) includes three different processes and purifies the CO2 for further utilization applications. The first unit, called Sour-Compression Unit (SCU) comes just after the dedust step. The gas is cleaned from its SOx and NOx components by means of a reactive and pressurized absorption in water and then flows to the Dehydration Unit. In this second unit, a Temperature Swing Adsorption (TSA) dual-bed process captures water at high pressure (30 bar) onto a solid adsorbent which can be silica gel or zeolite. Typically, regenerative desiccant dryers supplying a dew point of -40°C to -70°C are required. In the last unit, the gas coming from the dehydration unit is cooled and condensates are removed in a cryogenic unit consisting of a first flash at 30 bar. The vapor stream is then cooled and condensates are removed again in a second flash with a lowest temperature of -55°C to avoid the formation of dry ice at this pressure. The operating conditions depend on the CO2 purity objectives for the final product, or on the required CO2 recovery of the unit. This work is focused on the simulation in Aspen PlusTM and optimization of a de-SOx and de-NOx Sour-Compression Unit (SCU) carried out in four successive relevant steps (see Table 1), depending on the following applications (CO2 chemical conversion, CO2 storage or Enhanced Oil Recovery (EOR) requiring additional post-treatments as shown in Figure 1). Table 1: Steps and tools used for the optimization of the Sour Compression Unit process

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Objectives Construction of a comprehensive chemical mechanism Modelling of a two-columns process (conventional SCU) Parametric study Design of experiments (DOE) for optimization of a single-column configuration Economic evaluation and performances comparison of the twocolumn and the single-column configurations Use of DOE results Design and economic analysis for the single-column configuration reaching defined purity targets

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Steps 1

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Tools Literature data Aspen PlusTM Aspen PlusTM JMPTM Aspen PlusTM Aspen PlusTM

(i) First of all, a two-column process composed by two absorption columns (two compression levels) was studied. The first step was the construction of an accurate, comprehensive and realistic (compared to other sources in the literature) chemical mechanism. (ii) Moreover, for the single-column design, composed by one absorption column (one compression level), a parametric study was conducted and complemented by a Design Of Experiments (DOE). Denitrification and desulfurization performances are adapted to the required purity of the outcoming gas stream . It is principally dependent on the following CO2 utilization. When considering pipeline 5

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transportations, water, NOx and SOx contents have to be limited for pipeline specifications in order to avoid corrosion in the pipes (Project DYNAMIS, 2007), requirements for CO2 purity specifications for storage (Kather and Kownatzki, 2011) and CO2 purity specifications for Enhanced Oil Recovery (Pipitone and Bolland, 2009) are also fixed. It has to be noted that for the dehydration unit using a TSA (second step of the CPU), stricter levels of purities on SOx and NOx are required at the outlet of the SCU to avoid poisoning the solid adsorbents involved in the TSA. (iii) The two configurations of the SCU (two-column and single-column design are compared in terms of performances and costs). (i˅) The selected purity specifications have different levels of stringency in terms of SOx and NOx final concentrations allowing to test the model in different conditions. In order to identify, for these purity targets, the most favorable parameters combinations allowing to reach these specifications, the Design Of Experiments conducted with JMP® software has been completed by an economic analysis for the single-column design.

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In the context of the application of the SCU process in an oxy-fuel cement plant, the innovative aspects of the present work can therefore be divided into two categories: on the one hand a complete reaction mechanism related to the combined absorption of SOx and NOx in water under pressurized conditions, and including the interaction reactions involving these species, was implemented in Aspen PlusTM software, and on the other hand the current two-column SCU process was optimized in an one-column process including a completed and detailed parametric study. 2. Modeling of the Sour-Compression Unit 2.1. De-SOx and de-NOx scrubbing process

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Figure 2 shows that in this purification unit the raw CO2 is firstly compressed to 15 bar in a two-stage compressor; it enters the bottom of the first absorber of the Sour-Compression Unit and flows counter-currently to the aqueous scrubbing liquid (approximately 98% water and 2% acids).

Figure 2: General scheme of the de-SOx and de-NOx process

A splitter is used to recycle a part of the liquid flow to the top of the absorber and a water make-up is also provided. The washed gas then leaves the column at the top and is further compressed to 30 bar before entering at the bottom of the second absorber to flow counter-currently to the aqueous scrubbing liquid and remove remaining impurities. Another splitter is used to recycle a part of the liquid flow to the top of the second absorber and a water make-up is also provided in this case. During these absorption steps enhanced by the pressurized conditions, successively, SOx and NOx species in the raw CO2 are transferred and react to form acids. 6

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This purification process demonstrated efficient elimination of SOx and NOx components from power plants deriving flue gases, (White et al., 2013a). Based on this process, a Sour-Compression Unit (SCU) was designed and implemented using Aspen Plus TM V8.8, in order to evaluate the de-SOX/deNOX performances on oxy-fuel cement plant flue gases.

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The compositions of an outcoming flue gas from a cement industry applying an oxy-fuel combustion and the corresponding treated gas flow rate were used as input compositions of the SCU (Table 2). These data come from simulations conducted by the European Cement Research Academy (ECRA) of an oxy-fuel Best Available Techniques (BAT) cement kiln based on a 3000 tclinker/day production: Table 2: Compositions of an outcoming gas from an oxy-fuel cement industry (simulation – courtesy of ECRA) Molar basis 3.27 % 11.11 % 1.34 % 1.00 % 0.04 % 83.13 % 861 ppm 96 ppm 156 ppm

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Components O2 N2 Ar H2O CO CO2 NO NO2 SO2

Gas volume flow (wet) (Nm³/h)

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It is important to note that the second major species (after CO2) is nitrogen. Even if the cement plant works in oxy-fuel conditions, the percentage of 11.1 % is due to air intrusions in the cement kiln and to the use of nitrogen for cement manufacturing specifications.

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Design and operating parameters have been roughly estimated by means of Air Products pilot plant parameters (White et al., 2013b). The inlet compositions (Table 2) were tested in two types of configurations: a base case called “twocolumn process”, that was presented previously and an optimized case composed by only one absorber (one compression stage) called “single-column process”. The latter configuration is likely to lower installation and operating costs and at the same time keep the aimed performances of the process. In order to represent accurately this reactive absorption of SOx and NOx components in pressurized flue gas systems, an adapted chemical mechanism has been elucidated taking into account the influence of the two key elements, i.e. SOx/NOx interactions and pH.

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2.2. Construction and implementation of the SCU chemical mechanism

The SCU simulated performances are extremely dependent on the chemical mechanisms selected and equilibrium and kinetic data used (Laribi et al., 2017). The kinetic parameters of SOx/NOx interaction reactions were measured directly or evaluated on Air Products Vattenfall oxy-fuel CO2 Compression and Purification Pilot Plant at Schwarze Pumpe (White et al., 2013a) (Santos, 2015). These absorption tests showed that (i) SOx/NOx interaction reactions in liquid phase have a key influence on the simultaneous SOx and NOx absorption by water in pressurized flue gas systems and (ii) the pH level has a strong influence on the reaction pathways that occur and the type of products that are formed in the liquid phase (White et al., 2013).

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ACCEPTED MANUSCRIPT The chemistry and pH influence was studied by Normann and co-workers for pressurized flue gas systems deriving from an oxy-fuel power plant (Normann et al., 2013). They revealed similar conclusions as White et al. (White et al., 2013a). The selection of the reaction pathways related to the operating pH was based on Ajdari and coworkers studies (Ajdari et al., 2015). These revealed different possible reaction pathways that are dependent on the pH level and consequently the types of complex molecular and ionics sulfurnitrogen compounds (Table 3) that are formed in the liquid phase.

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In fact, according to pH operating conditions of 1
S-N complexes molecular species Nitrososulfonic acid ONSO3H alias NSSH

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S-N complexes ionic species Nitrososulfonate ONSO3- alias NSS Hydroxylamine N,N-disulfonate HON(SO3)2 2− alias HADS

Hydroxylamine N,N-disulfonic acid HON(SO3H)2 alias HADSH

Hydroxylamine N-sulfonate HONHSO3– alias HAMS

Hydroxylamine N-sulfonic acid HONHSO3H alias HAMSH

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Based on these studies, the elucidated complete chemical mechanism representing the SOx and NOx simultaneous absorption into water for pressurized systems in both gas and liquid phases, used in the model of the Sour-Compression Unit is detailed in Figure 3. It takes into account pH influence on the importance of the interaction reactions pathways that may occur and the S-N deriving complexes. The reaction mechanism without NOX/SOX interactions is also presented to show the effect on SOx/NOx absorption and the number of interaction reactions added: (i) the mechanism without considering interactions between the two species is composed by 15 reactions and (ii) the mechanism considering the SOx/NOx interactions occurring in pH range 1
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As the number of reactions is high (27), their implementation in Aspen PlusTM software was performed following a specific approach: in a first time the gas-phase reactions were implemented beginning with the interaction NO2+SO2, then the N-S complexes dissociation equilibria were added followed by the other kinetic reactions. Several reaction blocks were also created in the simulation environment in order to facilitate the convergence by progressiviely simulating with a higher number of reactions.

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Figure 3: Complete SCU Chemical Mechanisms involved in the pressurized SOx and NOx absorption process, without (light black) and with SOx/NOx interactions (light black + bold black) (reactions selected for 1
2.3. Kinetic and equilibrium data

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The complete reaction mechanism includes 27 reactions occurring in liquid and gas phases, reversible and irreversible reactions. The characteristics of the S-N complex species (Table 3) derived from the interactions between the SOx and NOx must be implemented and their properties estimated in order to have a comprehensive, accurate model and whose reactions in liquid phase are representative of a pH between 1 and 4 (Ajdari et al., 2015). Table 4 summarizes the selected mechanisms including the 27 reactions and their implementation in Aspen PlusTM for the process simulation. If reactions are available in Aspen PlusTM, these data are used, but otherwise, reactions are implemented using literature data or calculated from Gibbs free energies. In this case, the equilibrium constant of the reaction (Keq) is calculated from the Gibbs standard free energy ΔGo using the relation between ΔGo and Keq (∆G - Ln e ) where R is the ideal gas constant and T is the temperature.

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Concerning the kinetic constant (Kcin), a literature review was performed: depending on the data available, temperature dependent laws were implemented when possible, otherwise the values provided at 298 K were considered as the process temperature range is between 295 K and 305 K.

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Table 4: Characteristics of the reactions taken into account in the SCU chemical mechanisms.

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ACCEPTED MANUSCRIPT 2.4. Base case: the two-column process 2.4.1. Flowsheet and characteristics

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The flowsheet of an industrial two-column process design and corresponding to the base case implemented in Aspen PlusTM is presented in Figure 4. Design (columns sections, suitable column packing) and operating parameters (gas and liquid flow rates) have been adapted from the Air Products pilot installation (White et al., 2013b), checking adequate hydrodynamic conditions (actual gas velocity lower than 70% of the velocity leading to the flooding). The parameters used in the model for the design of the de-SOx and de-NOx gas/liquid contactors are specified in Table 5. The scaling-up to industrial scale was based on the same linear fluid velocities, giving the adapted sections and leading to the same degrees of approach to flooding. The wet gas flow of 70 000 Nm³/h from the BAT oxy-fuel cement plant (Table 2) at 50 °C and 1.063 bar must be conditioned in order to bring it to the temperature and pressure of absorption, i.e. 30 °C and 15 bar respectively. A conditioning unit is therefore necessary, composed successively by an isentropic compressor which brings the pressure of 1.063 bar to 4 bar, by a cooler that lowers the gas temperature to 30 °C and finally by a flash unit eliminating condensates due to cooling.

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Simulations of the Sour-Compression Unit were performed using rate-based calculations and Electrolyte NRTL model was selected for thermodynamic properties calculations as different electrolytes are considered (Song and Chen, 2009). The Redlich–Kwong-Soave equation of state with Boston-Mathias alpha function (Mathias, 1983) has been chosen for predicting gas phase properties and modelling gas-liquid phases equilibria, allowing accurate calculations at high pressures (Markočič and Knez, 2016).

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Figure 4: Detailed flowsheet of the SCU two-column process simulated in Aspen Plus including the temperature and conditioning step downstream of the reactive absorption process

In this case, the total gas flow rate used for the simulations corresponds to the outcoming gas of ECRA oxy-fuel combustion simulations (Table 2). Regarding the columns themselves, “ adFrac” model was selected. This model is able to handle both equilibrium reactions as well as kinetically limited reactions. For interested readers, more information on this model and more globally on the methods implemented in Aspen Plus TM software (e.g. mass and energy balances) are provided in (Taylor and Krishna, 2000).

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Liquid holdup per stage**

First absorber

Second absorber

16 Bottom Top 12 0.75 IMTP 25 mm 4 TM Calculated by Aspen Plus with (Stichlmair et al., 1989)

16 Bottom Top 12 0.75 IMTP 25 mm 4 TM Calculated by Aspen Plus with (Stichlmair et al., 1989)

OPERATING PARAMETERS: Calculation type Top stage pressure (bar) Total liquid flow rate (t/h) Recirculated flow rate (t/h) Recycle ratio Fresh water supply flow rate (t/h) T inlet column gas (°C) T inlet column liquid (°C)***

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Rate-Based 15 198 145 0.73 53 30 22

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DESIGN SPECIFICATIONS: Number of calculation stages Flue gas feed stage Water feed stage Column packed height (m) Packed height per stage (m)* Column packing type Column diameter (m)

Rate-Based 30 198 126.72 0.64 71.28 30 22

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*The most adequate packed height per calculation stage (m) was identified by means of several simulations varying the number of calculation stages for the same total packing height: a value of 0.75 of packed height per calculation stage allows obtaining an accurate result. **The liquid holdup per stage corresponds to the liquid quantity in a packing integration stage. When the holdup is specified by the user, the basis determines the units of the pre-exponential factor of the reaction kinetics. In such case, for rate-based calculations, the holdup values communicated by the user are used for initialization calculations. In the present case, the liquid holdup at each stage is calculated using the (Stichlmair et al., 1989) correlation which is valid for all packing types. ***A temperature of 18°C was taken for the fresh water. The fresh water supply at 18°C is mixed with the recirculated liquid heated by the gas (T inlet gas = 30°C), thereby the temperature of the liquid at the inlet of the column is 22°C.

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2.4.2. Results of the implementation of the base case in Aspen PlusTM

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The implementation of the chemical mechanism with its corresponding kinetic/equilibrium parameters using the respective design specifications (Table 5) and the inlet flue gas composition given by ECRA (Table 2) for the SCU gas-liquid absorbers leads to the following gas/liquid compositions (Table 6) and de-SOx /de-NOx performances (Table 7).

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The corresponding removal or abatement performances of the SCU relative to SOx and NOx species are calculated from the gas phase compositions: ( )

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where Ai is the abatement ratio, index i corresponds to SOx or NOx compound and Qmol,i-in/out is the molar flow rate of i compound at the inlet or outlet of the column.

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Table 6: Main results of the two-column process implementation GAS

SOx/NOx interaction complexes

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Temperature (°C) Pressure (bar)

Outlet second absorber 2.40 E-06 1.70 E-06 1.45 E-10 trace 7.48 E-09 6.52 E-08 trace 1.09 E-07 4.58 E-07 5.55 E-09 1.87 E-08 9.53 E-05 trace 1.19 E-07 2.48 E-10 4.84 E-06 6.36 E-06 1.20 E-09 trace trace trace 1.24 E-07 trace 2.30 E-05

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NOx species

9.79 E-06 0 0 trace 0 0 trace 1.16 E-04 3.65 E-05 2.25 E-08 7.54 E-08 2.18 E-05 trace 1.23 E-05 1.64 E-10 0 0 0 0 0 trace trace trace 0 8.37 E-01 1.13 E-01 3.28 E-02 1.36 E-02 2.69 E-03 4.03 E-04

Outlet second absorber 1.42 E-07 0 0 trace 0 0 trace 2.16 E-06 6.00 E-07 trace trace 2.62 E-06 trace 1.25 E-05 trace 0 0 0 0 0 trace trace trace 0 8.36 E-01 1.15 E-01 3.34 E-02 1.39 E-02 1.56 E-03 4.10 E-04

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SOx species

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SO2 HSO3 -SO3 SO3 HSO4 -SO4 H2SO4 NO NO2 N2O3 N2O4 HNO2 HNO3 N2O HNO NO2 NO3 NSS HADS HAMS NSSH HADSH HAMSH + H3O CO2 N2 O2 Ar H2O CO

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Inlet second absorber

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Inlet first absorber 1.57 E-04 0 0 0 0 0 0 8.67 E-04 9.63 E-05 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 8.37 E-01 1.12 E-01 3.29 E-02 1.35 E-02 3.33 E-03 3.99 E-04 30 15

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NB: trace = mole fractions < 1.00 E-10 Table 7: De-SOx and de-NOx performances of the two-column process

Abatement ratios (%) SOx NO NO2 NO+NO2 NOx CO2

First absorber 93.83 86.79 62.42 84.35 80.84 0.84

Second absorber 98.58 98.16 98.39 98.22 90.59 2.01

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Total SCU 99.91 99.76 99.39 99.72 98.20 2.84

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The interaction reactions occurring in the liquid phase are led by the pH of the solution that will define the type of products (S-N complexes: HADS, HAMS and their corresponding acids) formed in the liquid phase. The obtained pH of the liquid phase (between 2.12 and 3.06) confirms the range of pH selected for the choice of the different pathways relative to the SCU chemical mechanism. In this case, the major species in the liquid phase besides the ions relative to the acids dissociations (HSO3- , NO2- , NO3-) are HNO2 and HADSH. Unlike the SOx species that are completely removed, the NOx species complete removal implicate more challenging conditions (limiting factor).

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It has to be noted that the major species (CO2, N2, O2, Ar) are not the reacting ones. Even if a small part of the CO2 is absorbed in water, its concentration remains quite constant, i.e. 83.7 %, characteristic of an oxy-fuel cement plant.

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The advantage of working with a two-column process lies in the global absorption efficiency (Table 7). High abatement ratios are obtained (99.98% for SOx species and 97.94% for NOx species), which can be considered as a validation of the conventional two-column process (White et al., 2013a). Studies of White et al., 2013 (White et al., 2013a) proved that the first absorber is the de-SOx column and the second one is the de-NOx column. Besides, White et al., 2013 (White et al., 2013a) also concluded that for NOx removal, up to the exit of the 15 bar column, around 80% of the mixture of NO and NO2 should be removed from the gas phase at this point, with the rest being eliminated in the 30 bar column which has also been observed in our results.

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Finally, emissions of N2O in the gas phase represent an important observation that has also been shown in results relative to the Sour-Compression pilot plant Unit of Air Products (White et al., 2013a) and has also been observed in the present work (Table 6). Results on Table 7 show that the two-column process is adapted to very strict purity specifications, e.g. for the second unit of the CPU, the dehydration unit (over 99% of SOx abatement ratio and over 98% of NOx abatement ratio). Challenging purity specifications could be reached with the twocolumn process; the purpose now is to test an eventual optimized process in a single-column configuration.

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2.5. A single-column process

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The strategy of the single-column process consisted in the investigation of the possible optimization of the base case. The purpose was to reach the fixed purity targets of the SCU and at the same time reduce the complexity of implementation, decrease both the installation and the operating costs. This methodology has been validated by Iloeje and co-workers (Iloeje et al., 2015). They developed a model based on works of (White et al., 2013a) and conducted a parametric analysis using a singlecolumn absorber for nitrogen and sulfur oxides removal (water scrubbing) from oxy-combustion flue gas applied to power plant cases. The chemical mechanism taken in the model of Iloeje and co-workers (Iloeje et al., 2015) consists of five reactions with only one SOx/NOx interaction reaction (reaction A.19 in Table 4) considered at equilibrium which will eliminate all the SOx and NOx components without necessity of considering the other interactions. Nevertheless, this reaction is usually considered as an irreversible one ((Jaffe and Klein, 1966); (Armitage and Cullis, 1971); (Penzhorn and Canosa, 1983) and (Ajdari et al., 2015)). This mechanism oversimplifies the chemistry and fails to capture the effect of pH on the SO2 absorption from the gas phase and the different products formed in the liquid phase (Ajdari et al., 2016). Despite the oversimplified chemical mechanism used, the methodology given by Iloeje and coworkers represents an interesting strategy: after simulation of a two-column process that achieves 14

ACCEPTED MANUSCRIPT the complete removal of SOx and NOx from the CO2 stream, the process was adapted from a twocolumn process to a single-column process. Iloeje et al. (Iloeje et al., 2015) demonstrated by means of pressure sensitivity studies that this new design can meet the same SOx/NOx removal targets as the two-column process in fewer column stages and half the feed water requirement by exploiting the pressure dependence of the rate determining NO oxidation reaction (reaction A.4 in Table 4). Accordingly, the optimization developed in this work is based on this strategy in order to decrease the installation costs and energy requirements of the global CO2 purification process: the twocolumn process has been adapted to a single-column process keeping the same complete reaction mechanism.

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2.5.1. Flowsheet and characteristics

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The overall single-column flowsheet investigated here is illustrated in Figure 5. It consists of a multistaged isentropic compression (including heat exchanger(s) between the compression stages) where the flue gas outcoming from an oxy-fuel cement industry is compressed to a pressure between 1 bar and 60 bar, depending on the pressure level in the single absorption column. The next step is a counter-current absorption in a water scrubber with simultaneous removal of SOx and NOx impurities from CO2 gas stream.

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Figure 5: Detailed flowsheet of the SCU optimized single-column process simulated in Aspen Plus

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The single-column process operating and dimensional parameters are presented in Table 8. The reference case was chosen by minimizing the operating and dimensional parameters while avoiding the flooding conditions. For this reason, the column packed height was set to 6 m, the total liquid flow rate to 198 t/h, the maximum recirculation rate was chosen (0.9) and the pressure was fixed to 15 bar. For the variations around these reference values the column packed height ranged from 3 m to 12 m, the pressure of the system between 1 bar and 60 bar, the total liquid flow rate between 132 t/h and 395 t/h and the recycle ratio between 0 and 1 (the other parameters remaining fixed in each case).

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Table 8: Design specifications and operating parameters of the SCU gas-liquid contactor

4-16 Bottom Top 3-12 0.75 IMTP 25 mm 4 TM Calculated by Aspen Plus with (Stichlmair et al., 1989) Packing Holdup Correlation

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Rate-Based 15 198 178 0.90 20 30 22

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OPERATING PARAMETERS : Calculation type Top stage pressure (bar) Total liquid flow rate (t/h) Recirculated flow rate (t/h) Recycle ratio Fresh water supply flow rate (t/h) T inlet column gas (°C) T inlet column liquid (°C)

8 Bottom Top 6 0.75 IMTP 25 mm 4 TM Calculated by Aspen Plus with (Stichlmair et al., 1989) Packing Holdup Correlation

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Liquid holdup per stage

Ranges of variation for sensitivity analysis

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DESIGN SPECIFICATIONS: Number of calculation stages Flue gas feed stage Water feed stage Column packed height (m) Packed height per calculation stage (m) Column packing type Column diameter (m)

Reference values

Rate-Based 1-60 132-395 0-198 0-1 0-198 30 22

2.5.2. Parametric study of the SCU

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Sensitivity analyses were conducted for the SCU in order to better understand its behavior when changing operating and dimensional parameters. The relevant purpose of this study is the identification of the most influent parameters that will allow further optimization of the CO2 purification process. Among the various operating parameters of the SCU, the diameter, packing type, inlet gas flow rate and composition were kept constant and the effects of varying the column packed height, the operating pressure, the total liquid flow rate and the recycle ratio of the column were studied separately (keeping the other parameters to reference values as specified in Table 8). Results of the sensitivity analyses are presented in Figure 6.

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Figure 6: Influence on the SOx and NOx abatement ratios of: (a) the total column packed height; (b) the operating pressure of the absorber; (c) the total liquid flow rate and (d) the recycle ratio.

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Even if the abatement ratios in SOx and NOx species increase obviously with the column packing height (especially for NO2), globally ANOx and ASOx are not significantly dependent on the packing height (Figure 6 (a)). However, the abatement ratios in SOx and NOx components are largely dependent on the operating pressure of the system (Figure 6 (b)). Indeed, both for SOx and NOx components, when the pressure is augmented, the abatement ratios increase; relative increases of 67% and 94% are shown for respectively SOx and NOx components, the evolution being calculated from 1 bar to 60 bar. Moreover, almost all the SOx components are removed at pressure of 5 bar with a stagnation of the abatement ratio at approximately 94% for higher pressures. An important conclusion arising from the analysis of Figure 6 (b) is that compared to the SOx components, the NOx species require a more significant increase of pressure in order to be removed, with a maximum of abatement ratio of approximately 89% reached at 60 bar. Indeed, it has to be noted that even when working with the highest operating levels for all the tested parameters, the NOx components will not be completely removed, and consequently, the limiting factor will be the abatement ratio of NOx components. Besides, for both SOx and NOx components, the removal performances of the purification unit are enhanced when working with higher liquid flow rates (QL) in the SCU (Figure 6 (c)). Nevertheless, there is a maximum abatement ratio for both SOx and NOx components, reached and stagnating at respectively 98% for QL≥200 t/h and 88% for QL≥250 t/h. 17

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Finally, increasing the recycle ratio of the absorbers has a negative impact on the removal of SOx and NOx species (Figure 6 (d)). This is linked to the fact that in order to maintain a constant total liquid flow rate, there is less fresh water supplied to the system when increasing the column bottom recycle. It has to be noted that, even a very low recycle ratio (i.e. 0.1) could not allow the complete removal of the NOx components which are therefore more problematic, the maximum NOx abatement ratio being 90%, corresponding to a remaining quantity of 96 ppm in the purified gas (contrary to SOx components which are totally removed). In the case of Iloeje et al. (Iloeje et al., 2015), almost all the NOx components were removed with an abatement ratio of 96.53% on NOx species corresponding to an outlet composition of 9.1 ppm (with 262 ppm as inlet composition). However, this phenomenon results from their model including a simplified reaction mechanism and the consideration of the interaction reaction A.19 as an equilibrium reaction (nonrealistic conditions). The SOx/NOx simultaneous absorption in pressurized flue gas systems into water has as consequence the formation of different acid forms in the liquid phase, as represented in Table 6. As a result, when the bottom recycle of the loaded solution is increased, the solution becomes logically even more acid (and consequently less efficient for SOx/NOx removal) as shown in Table 9.

Recycle ratio pH inlet absorber pH outlet absorber

0 3.29 2.61

0.1 2.88 2.28

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Table 9: Evolution of the pH with the bottom recycle ratio 0.2 2.66 2.16

0.3 2.60 2.12

0.4 2.43 2.11

0.5 2.28 2.10

0.7 2.22 2.08

0.8 2.21 2.08

0.9 2.13 2.04

1 2.05 1.91

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The main conclusions highlighted in these sensitivity analyses are: (i) the most influent factors are the column operating pressure and the recycle ratio; (ii) the NOx components removal requires more challenging conditions than the SOx species for all the tested parameters, thus limiting the operating points of the purification process.

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The effect of each parameter has been studied separately in this section. However, in order to find the operating points allowing to reach a specific purity target of SOx and NOx species and at the same time to optimize economically, energetically and environmentally the process, the simultaneous variation of the most influent parameters is necessary to identify the best compromise for the SCU.

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3. Optimization of the SCU process through a Design Of Experiments 3.1. Characteristics of the full factorial design

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A comparative study was conducted through a Design Of Experiments (DOE) in order to identify the working points of the system for each purity target fixed for NOx and SOx removal and at the same time optimize the SCU. The statistical application JMP®13 was used for the experimental design and regression analysis of the experimental data. The operating factors taken into account are the same ones used for the parametric study of the SCU, keeping the column diameter, packing type and gas flow rate constants as indicated in Table 10. Note that for four parameters selected, the two levels have been coded replacing the low level by -1 and the high level by 1, following the conversion (Goupy and Lee, 2007): (

-

)

(2)

Where is the centered and scaled variable (or coded variable or coded unit) corresponding to the levels -1 or +1 of the factor i (i ranging from 1 to 4); is the original variable corresponding to the 18

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(

is the central value (level 0) of the original variable (

) and the

(

(

-

)

) being calculated in function of

as indicated in

Table 10.

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For this purpose, and using the parameters presented in Table 10, a full factorial design consisting of four factors, each one with two "levels" (24) was chosen. The variation ranges relative to the factors were fixed for the DOE taking into account the different evolutions revealed by the sensitivity analysis conducted before, then the central point and the step were deduced. There is an interaction effect when a response is conditioned by the coupling of two factorial variables, the product of two variables interaction is also called the cross product. These interactions may occur between two different factors, or between the same factors corresponding to a second-degree model. A full factorial design may allow studying the effect of each factor on the two response variables namely the abatement ratios of SOx (ASOx) and NOx (ANOx) components, as well as the effects of interactions between the factors by means of their different combinations. Therefore, the responses ASOx and ANOx are calculated using the following relations:

)

(

)

∑

∑

∑

∑

∑

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(

∑

(3)

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Where ( ) is the abatement ratio of SOx or NOx components taking the central value, , of each parameter (see Table 11 and Table 12); are the different coefficients of the centered and scaled variables relative to the factors i and j for SOx abatement and NOx abatement respectively. Table 10: Characteristics of the SCU DOE

( )

Corresponding levels ( ) Central value (level 0)

)

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(

Recycle ratio

30

198

395

6

12

0.4

0.9

-1

+1

-1

+1

-1

+1

-1

+1

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Range of variation

Responses

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Pressure (bar) X1

Factors Total liquid flow rate Column packed height (t/h) (m) X2 X3

X4

22.5

296

9

0.65

7.5

99

3

0.25

ASOx (%) ANOx (%)

For a 24 two-level full factorial design, 16 simulations have been conducted in Aspen PlusTM. A second-order model was fitted to the Aspen PlusTM simulation results by multiple regression analysis for each response variable studied. The quality of the fitted model was statistically verified by the magnitude of the coefficient of determination and its statistical significance was evaluated by the F-test analysis of variance (ANOVA). The coefficients of the response surface were evaluated using the Student t-test. Note that a Student t-test is a commonly used statistic test that can be applied, for example, to determine if two data sets are significantly different from each other. More information on the Student t-test and on the Student’s t-distribution can be found in (Dreier and Kotz, 2002). 19

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3.2. Results of the simulations

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The simulations results of the full factorial design of the SCU are characterized in Table 11. The values of the responses relative to the central point given by Aspen PlusTM (simulation number 17 in Table 11), i.e. 98.37 and 90.75 for ASOx (central point) and ANOx (central point) respectively allowed evaluating the accuracy of the model by comparing them with the responses predicted by the JMP® model, i.e. 96.70 and 88.37 for ASOx (central point) and ANOx (central point) respectively. The evolution of the operating parameters and the dimensional parameter represented in Figure 6 refers, as expected, to a non-linear response of the abatement ratios of SOx and NOx, particularly in the case of pressure that presents an evolution of the second degree reflecting a quadratic model. For this reason, eight additional points were required for the full factorial design: these points are called axial points (points from 18 to 25 in Table 11) and the corresponding model becomes a central composite design. Table 11: Simulation results of the central composite design Points relative to the full factorial design X4

ANOx (%)

X2

X3

1

15

198

2

30

198

6

0.4

99.07

90.88

6

0.4

99.90

96.84

3

15

395

6

0.4

99.51

92.96

4

30

5

15

395

6

0.4

99.90

98.07

198

12

0.4

99.27

95.64

6 7

30

198

12

0.4

100.00

97.20

15

395

12

0.4

99.68

8

30

395

12

0.4

100.00

0.9

94.34

72.89

0.9

94.60

84.26

0.9

95.16

84.27

0.9

96.61

85.96

X3

X4

ASOx (%)

ANOx (%)

18

33.075

296

9

0.65

98.66

92.99

19

11.925

296

9

0.65

97.17

88.48

20

22.5

435

9

0.65

99.38

96.18

21

22.5

157

9

0.65

96.19

85.13

22

22.5

296

13.23

0.65

97.58

90.24

22.5

296

4.77

0.65

93.38

86.12

24

22.5

296

9

1

92.08

70.20

98.49

25

22.5

296

9

0.2975

100.00

98.87

198

6

30

198

6

11

15

395

6

12

30

395

6

13

15

198

12

0.9

95.73

73.66

14

30

198

12

0.9

96.71

91.52

15

15

395

12

0.9

96.76

90.88

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X2

23

9

17

X1

96.57

10

16

Simulation N°

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X1

Complementary axial points of the central composite design

ASOx (%)

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30

395

12

0.9

98.27

91.53

22.5

296

9

0.65

98.37

90.75

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3.3. Modeling characteristics in JMP® software

The analysis of the results in JMP® allows determining the most influent factor(s) for each abatement ratio (Table 12). p-value is the statistic parameter used to evaluate the importance of the effect on the calculated responses; the smaller p-value is, the more significant is the effect of the factor. The results of these parameters interactions and their estimated coefficients (effect) in the equations relative to SOx abatement ratio and NOx abatement ratio are presented in Table 12. The absolute value of the coefficients is proportional to the effect of each interaction.

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Table 12: Estimated coefficients and effects of the different factors for ASOx and ANOx

X4 X1 X2 X3 X1² X1*X2 X2² X2*X4 X3*X4 X4² X1*X4 X3² X1*X3 X2*X3

p-value

0.00001 0.01448 0.01611 0.02715 0.07198 0.09436 0.09465 0.12387 0.20672 0.26302 0.26545 0.34102 0.83026 0.91220

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A at Center point Recycle ratio Pressure Total liquid flow rate Column packed height Pressure² Pressure*Total liquid flow rate Total liquid flow rate² Total liquid flow rate*Recycle ratio Column packed height*Recycle ratio Recycle ratio² Pressure*Recycle ratio Column packed height² Pressure*Column packed height Total liquid flow rate*Column packed height

Coefficients estimations ai or aij bi or bij 96.70 88.37 -2.02 -6.55 0.43 2.56 0.54 2.51 0.66 1.70 0.82 1.49 0.05 -1.79 0.75 1.45 0.29 1.63 0.39 0.77 -0.13 -1.63 0.12 1.14 -0.41 0.20 0.04 -0.21 -0.02 0.11

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Globally, the lower p-value is, the greater the influence of the factor (Bataille et al., 2002) (Greenland et al., 2016). Consequently, we can deduct from Table 12 that the recycle ratio is the most influent parameter on the SOx and NOx removal efficiencies, the operating pressure and total liquid flow rate being also influent parameters but with more restricted effects. The same conclusion could be drawn for ANOx. Concerning the SOx species, the most influent parameters are the recycle ratio which has a negative impact on the performances and the column packed height. The effects of the interactions between the factors are not considered to be influent in this case. The observed values of the central point are determined by the simulations on Aspen PlusTM but are also predicted by modeling on JMP®. The respective relative differences between these responses are representative of the accuracy of the JMP® model. On the one hand, the difference between the observed value and the predicted value of the abatement ratio is 1.70 % for the SOx compounds and 2.62 % for the NOx compounds, corresponding to a quite precise model. On the other hand, the accuracy of the model to predict the results is characterized in Figure 7, showing the difference between the observed values (obtained by the simulation in Aspen PlusTM) and the ones predicted by the JMP® model.

Figure 7: Leverage plot of the observed values in function of the predicted ones by JMP® software of: (a) ASOx and (b) ANOx

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ACCEPTED MANUSCRIPT The visualization of the confidence intervals (red zones in Figure 7) can be applied to the F-test by using a leverage plot (Figure 7). The plots display the values of ASOx and ANOx calculated by Aspen PlusTM in function of the predicted values of ASOx and ANOx respectively, a regression line (red line), and 95% confidence intervals. The predicted values by JMP® software can be considered as accurate to represent the system since the regression coefficient has a value of 0.9 and the graphic shows only few points for ASOx and ANOx outside of the student’s distribution (95% confidence interval). 3.4. Optimization for three purity targets of the SCU

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3.4.1. Areas of responses obtained with JMP®

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The required purity of the gas stream outcoming from the SCU is principally dependent on the following CO2 transport/storage/utilization steps. The model obtained allows predicting the area in which a specific purity target is reached in function of the selected factors, within the ranges defined in Table 10. As an illustration of the model effectiveness, three specifications with increasing levels of SOx and NOx purity requirements, detailed hereafter, were considered: target 1, target 2 and target 3. The same concentrations of SOx and NOx components (in ppm) remaining in the CO2 after treatment will be taken for each target and applied to cement plant flue gases in this work. For target 1, the purity specifications presented in Table 13 (Kather and Kownatzki, 2011) concern purified gases from power plants working with coal fired oxy-fuel process. The applied purification process for storage purposes (COORAL German research project) consists in a distillation technology allowing a final purity of the separated CO2 stream from coal fired oxy-fuel process over 99.9 vol.%.

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Concerning target 2, it corresponds to a purified flue gas for Enhanced Oil Recovery (EOR)/storage purposes (Pipitone and Bolland, 2009), deriving from a distillation process applied to CO2 impurities from oxy-fuel power plants working with pulverized coal (Table 13). The flue gas separation takes place through a distillation process applied to gas deriving from pulverized coal (operating temperature between 54 °C and 10 °C). Before EOR/storage, CO2 purity is over 99%.

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The specifications relative to target 3 are linked to the second block of the CO2 Purification Unit, namely the dehydration unit. These targets are the most challenging specifications adopted for the SCU, to avoid poisoning adsorbents (zeolites) intervening in the Temperature Swing Adsorption process and are presented in Table 13.

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SCU inlet compositions

SOx (ppmv) NOx (ppmv)

Table 13: CO2 purity targets tested for the SCU

Target 1 Outlet Abatement compositions ratios

Target 2 Outlet Abatement compositions ratios

Target 3 Outlet Abatement compositions ratios

156

70

55.13%

37

76.28%

10

93.59%

957

100

89.63%

33

96.55%

20

97.92%

Several configurations of the factors allow reaching the purity targets fixed in Table 13. Nevertheless, to limit environmental impact of acid water, the strategy used here is limiting the amount of purged acid effluents, therefore limiting the fresh water supply and increasing the recycle ratio (for a fixed total liquid flow rate). The strategy described leads to minimizing the total liquid flow rate (low level in Table 10), namely QL=198 t/h, keeping the column packed height, Hcol=12 m and studying the 22

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effect of the remaining two variables, namely recycle ratio and pressure, in the fixed ranges. Results of these three targets are given in Figure 8 and investigated by means of an economic analysis. The red and blue lines represent the iso-response curves of the SOx and NOx abatement ratios respectively. The pink zone is the area where the SOx purity requirements are reached, the blue zone corresponds to the area where the NOx purity requirements are obtained. The overlap of both delimited zones, purple zone, is the area where the global purity target is attained for both species.

Figure 8: Area representing the purity specifications of: (a) target 1; (b) target 2 and (c) target 3 for QL =198 t/h and Hcol = 12 m in function of P and recycle ratio (blue lines: ANOx and red lines: ASOx)

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The graphs in Figure 8 (a), (b) and (c) showed that the three purity targets could be reached with the single-column configuration. These graphs confirm the observation relative to NOx abatement ratio being the limiting factor since the area where the global purity target is reached corresponds to the one of NOx species. Besides, this area becomes smaller when the severity of the purity target increases, showing a logic response of the JMP® model. For each target, two extreme scenarios are shown in Figure 8 and Table 14 allowing to reach the purity specifications fixed in Table 13:  scenario of highest pressure level and highest recycle ratio;  scenario of lowest pressure level and lowest recycle ratio.

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Table 14: Characteristics of the extreme scenarios for each purity target for QL =198 t/h and Hcol = 12 m Scenario

1.a.

1.b.

2.a.

2.b.

3.a.

3.b.

Pressure (bar)

30

15

30

26

30

28

Recycle ratio

0.822

0.55

0.568

0.4

0.498

0.4

Fresh water supply (t/h)

35.6

89.1

85.5

118.8

99.4

118.8

Having identified for each purity target the operating conditions corresponding to each scenario, an economic analysis was performed in order to evaluate the capital (CAPEX) and operating (OPEX) costs for the selected scenarios. It has to be noted that taking into account the strategy adopted for optimization (fixing QL =198 t/h and Hcol = 12 m) for target 3, the two scenarios presented very similar operating conditions (and quite equivalent costs); hence only one scenario (3.a.) has been economically analyzed. Details of the economic evaluations conducted for the two-column and single-column configurations of the SCU are given in section 4. 23

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4. Economic evaluation of the Sour-Compression Unit

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In these analyses, the economic evaluation tool in Aspen PlusTM was used to calculate the Capital expenditure (CAPEX) and Operating expenditure (OPEX) to reach the considered targets. The flowsheets considered for the economic analysis are presented in Figure 4 for the two-column process and in Figure 5 for the single-column process. The analysis has been conducted considering the totality of the flue gas coming from the oxy-fuel cement plant (70 000 Nm³/h) implying a fixed CO2 purified flow rate for the SCU. The acid water treatment technique considered here has been taken into account in the economic analysis of Cullivan et al. (Cullivan and Cullivan, 2016) and consists in the use of ion exchange resins. These resins are polymers that have the property of exchanging specific ions with other ions contained in a solution passing through them. The synthetic resins are used for elements separation applications and water purification (water softeners and water deionizers). Acid Retardation (Acid Sorption) is a sorption process by which ion exchange resins adsorb strong acids (HCl 6 wt.% in the case of Cullivan et al.) while excluding other elements like water and metal salts of these acids. Acid adsorption takes place in a device called Acid Purification Unit or APU (Dejak and Munns, 2017) and has been proven to be a low-cost purification system (Dejak and Munns, 2017) commercialized by Eco-Tec (Eco-Tec, 2016). It can be applicable to a variety of acids including sulfuric/nitric/hydrofluoric mixtures used for stainless steels. It is important to note that the concentrated acids formed in this process are destined to further commercialization. Therefore, this allows to recover the capital costs required for the acid water treatment process (Dejak and Munns, 2017). The details relative to the utilities costs for the OPEX calculations are listed in Table 15.

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Table 15: Utilities costs taken for OPEX calculations Utilities

Costs

50 €/MWh (based on 2016 average electricity costs (European Commission, 2017). * 0.30 €/t fresh water (Li et al., 2016) 1.964 €/t water to treat (Cullivan and Cullivan, 2016) * 28.22 €/ton of steam (Sun et al., 2016)

Electricity

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Fresh water

Water treatment by acid sorption

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Steam

Total treated CO2 flow rate

114 ttreated CO2 /h

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*Currency conversion in September 2018: 1 US $ is e ual to 0.85 €.

As indicated in Table 15, released acid effluents treatment is considered here. In fact, the cost of acid neutralization and acid resin sorption being quite similar, the second method has been chosen since it allows replacing the fresh water supply utility by the amount of treated water. The equation 4 was used to calculate the OPEX (in €/ttreated CO2) for each utility presented in Table 15. -

-

(

(4)

)

Where OPEX-U is the OPEX relative to the utility U (i.e.: electricity, fresh water, water treatment, steam), in €/ttreated CO2 , Cost-U is the utility cost in €/MWh (electricity) or €/t (others) and Consumption-U is the process consumption of the utility U in MW or in t/h. 4.1. Comparison of the two-column and the single-column configurations for target 3 24

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The comparison of the installations costs relative to the two configurations of the SCU are presented in Figure 9.

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Figure 9: Comparison of the CAPEX relative to the two-column (hatched in blue) and single-column (full in orange) configurations for target 3

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The analysis of the equipment costs relative to the two-column and single-column configurations shows that the total CAPEX relative to the latter is logically lower than the one of the base case (twocolumn process). Within the e uipment’s direct costs, the compressors present the most significant proportion with over 80% of the total CAPEX.

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4.1.2. Analysis of Operating expense (OPEX)

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The comparison of the operating costs (representing an ongoing day-to-day basis cost) relative to the two configurations of the SCU is presented in Figure 10.

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Figure 10: Comparison of the OPEX relative to the two-column and single-columns configurations for target 3

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The analysis relative to the operating costs presented in Figure 10 shows that the compression (electricity) is preeminent for both configurations and the costs relative to the compressions are quite similar since in both cases there are three stages of compression (up to 30 bar). An important difference can be observed for water treatment utility, which is quite higher for the two-column process. Indeed, steam utility for the heat exchanger (Figure 4) is required for a two-column process before the second column while in the single-column configuration the heat exchanger is included in the multi-stage compressor. This operating costs analysis shows globally an OPEX of 7.45 €/ttreated CO2 for the two-column configuration and 6.98 €/ttreated CO2 for the single-column one. Consequently, for the most stringent purity target (target 3), the single-column configuration was shown to be more advantageous in terms of CAPEX and OPEX than the base case. 4.2. Economic analyses of the single-column configuration relative to targets 1 and 2

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An economic evaluation is required in order to optimize the CO2 purification process considered for a 70 000 Nm³/h oxy-fuel cement plant outcoming gas and for a treated CO2 flow rate of 114 ttreated CO2 /h. Therefore, the costs of the SCU were evaluated for the two purity targets selected (targets 1 and 2). For each target, the costs of two extreme scenarios (Figure 8) allowing to reach the required purity were compared.

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4.2.1. Analysis of the Capital expenditure (CAPEX)

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Within the CAPEX (piping, electricity, paint, insulation…) the equipment involves the most important contribution to the direct costs, thus Figure 11 (a) and (b) presents the repartition of the equipment direct costs and the total CAPEX for both scenarios of target 1 and target 2 respectively.

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Figure 11: Representation of the total CAPEX and the equipment total direct costs for: (a) Target 1 and (b) Target 2 (scenario a. hatched blue and scenario b. in orange)

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For both targets considered, the compression step is clearly the most cost demanding operation, requiring over 88% of the total direct costs and approximatively 50% of the total CAPEX. 4.2.2. Analysis of Operating expense (OPEX)

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The other significant costs to consider are quantified by the operating expenditure (OPEX). Results of the total OPEX considering treated water reuse including the repartition of the different utilities are characterized in Figure 12 for targets 1 and 2.

Figure 12: Details of the OPEX relative to target 1 and target 2 (scenarios a. and b. respectively)

The OPEX is quite similar for scenario 1.a. (P = 30 bar and recycle ratio = 0.822) and scenario 1.b. (P = 15 bar and recycle ratio = 0.55) due to the fact that even if the utility concerning the compression is lower for scenario 1.b., the one concerning fresh water supply and consequently acid water treatment is more important for scenario 1.b. than for scenario 1.a. due to the lower recycle ratio. Nevertheless, even if the recycle ratio and consequently the fresh water supply utility has an important impact on the total OPEX, the compression to 30 bar for scenario 1.a. and to 15 bar for 27

ACCEPTED MANUSCRIPT scenario 1.b. is clearly the most predominant, the electricity costs representing 89% and 74% of the total OPEX for respectively scenarios 1.a. and 1.b.. An important observation from Figure 12 is that even if the pressure is higher, the OPEX is globally less important for scenario 2.a. (P = 30 bar and recycle ratio = 0.568) than for scenario 2.b. (P = 26 bar and recycle ratio = 0.4). Consequently, minimizing the recycle ratio is important. However as shown in Figure 12 for target 1, recycle ratio does not have the most important impact on the OPEX which is clearly determined by the compression step in this case (the electricity costs representing 78% and 70% of the total OPEX for respectively scenario 2.a. and 2.b.).

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5. Conclusion

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The development of the oxy-fuel combustion in the cement plant is currently investigated with possible future retrofitting. Therefore, the study of the entire CO2 treatment chain applied to oxyfuel combustion, before implanting it at industrial scale is necessary for the cement industry. In this context, a CO2 de-SOx and de-NOx process was here simulated in Aspen PlusTM. The purpose of the present work was to optimize the configuration and operating conditions of the CO2 de-SOx/deNOx. This step is necessary to reach the purity specifications on SOx and NOx components required for different CO2 utilizations, together with minimizing the operating (OPEX) and installation (CAPEX) costs. The elucidation of the coupled effect of the pH and the main SOx/NOx interactions occurring in the liquid phase, allowed to represent properly the reactive absorption of SOx and NOx components and to evaluate the performances of the SCU. The de-SOx and de-NOx two-column SCU unit allowed to reach purity targets of the three CO2 applications defined in this work. These purity targets were characterized by different levels of stringency concerning NOx components. For these purity targets, an optimization of the operating parameters and conditions was achieved by means of a parametric analysis and a Design Of Experiments with a single-column configuration. This optimization was completed by economic evaluations of the CAPEX and OPEX in order to lower the economic and energetic consumptions of the process. The single-column process showed more reduced installation (CAPEX) and operating (OPEX) costs than the two-column process, namely a relative decrease of 10% and 6% respectively, confirming the optimization strategy adopted. For this advantageous SCU single-column configuration, the installation costs (CAPEX) ranged between 22 and 25 M€ and the operating costs (OPEX) between 5.9 and 7 €/ttreated CO2. CAPEX and OPEX studies showed that the compression step is the most cost demanding operation with over 88% of the total equipment costs and over 70% of the total operating requirements.

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The developed de-SOx and de-NOx modeling tool could also be considered for other CO2 applications requiring different CO2 purity targets. An example is the CO2 transport or the CO2 purification applied to flue gases derived from partial oxy-fuel combustions, since the CO2 composition in the flue gas is also quite high in these conditions. Acknowledgments The authors would like to acknowledge the European Cement Research Academy (ECRA) and HeidelbergCement Company for their technical and financial support accorded to the ECRA Academic Chair at the University of Mons. References 28

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