novel process concept for cryogenic co2 capture

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Novel process concept for cryogenic CO2 capture Tuinier, M.J. DOI: 10.6100/IR719418 Published: 01/01/2011 Document Version Publisher’s PDF, also known as Version of Record (includes final page, issue and volume numbers) Please check the document version of this publication: • A submitted manuscript is the author's version of the article upon submission and before peer-review. There can be important differences between the submitted version and the official published version of record. People interested in the research are advised to contact the author for the final version of the publication, or visit the DOI to the publisher's website. • The final author version and the galley proof are versions of the publication after peer review. • The final published version features the final layout of the paper including the volume, issue and page numbers. Link to publication Citation for published version (APA): Tuinier, M. J. (2011). Novel process concept for cryogenic CO2 capture Eindhoven: Technische Universiteit Eindhoven DOI: 10.6100/IR719418 General rights Copyright and moral rights for the publications made accessible in the public portal are retained by the authors and/or other copyright owners and it is a condition of accessing publications that users recognise and abide by the legal requirements associated with these rights. • Users may download and print one copy of any publication from the public portal for the purpose of private study or research. • You may not further distribute the material or use it for any profit-making activity or commercial gain • You may freely distribute the URL identifying the publication in the public portal ? Take down policy If you believe that this document breaches copyright please contact us providing details, and we will remove access to the work immediately and investigate your claim. Download date: 09. Apr. 2018

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Page 1: Novel process concept for cryogenic CO2 capture

Novel process concept for cryogenic CO2 capture

Tuinier, M.J.

DOI:10.6100/IR719418

Published: 01/01/2011

Document VersionPublisher’s PDF, also known as Version of Record (includes final page, issue and volume numbers)

Please check the document version of this publication:

• A submitted manuscript is the author's version of the article upon submission and before peer-review. There can be important differencesbetween the submitted version and the official published version of record. People interested in the research are advised to contact theauthor for the final version of the publication, or visit the DOI to the publisher's website.• The final author version and the galley proof are versions of the publication after peer review.• The final published version features the final layout of the paper including the volume, issue and page numbers.

Link to publication

Citation for published version (APA):Tuinier, M. J. (2011). Novel process concept for cryogenic CO2 capture Eindhoven: Technische UniversiteitEindhoven DOI: 10.6100/IR719418

General rightsCopyright and moral rights for the publications made accessible in the public portal are retained by the authors and/or other copyright ownersand it is a condition of accessing publications that users recognise and abide by the legal requirements associated with these rights.

• Users may download and print one copy of any publication from the public portal for the purpose of private study or research. • You may not further distribute the material or use it for any profit-making activity or commercial gain • You may freely distribute the URL identifying the publication in the public portal ?

Take down policyIf you believe that this document breaches copyright please contact us providing details, and we will remove access to the work immediatelyand investigate your claim.

Download date: 09. Apr. 2018

Page 2: Novel process concept for cryogenic CO2 capture

Novel Process Concept for Cryogenic CO2 Capture

Page 3: Novel process concept for cryogenic CO2 capture

Samenstelling promotiecommissie:

prof.dr. J. Meuldijk, voorzitter Technische Universiteit Eindhovenprof.dr.ir. M. van Sint Annaland, promotor Technische Universiteit EindhovenProf.Dr.-Ing. A. Seidel-Morgenstern Otto-von-Guericke-Universitat

Magdeburgprof.dr. G.J. Kramer Shell / Universiteit Leidenprof.dr.ir. J.A.M. Kuipers Technische Universiteit Eindhovenprof.dr.ir. T.H. van der Meer Universiteit Twentedr.ir. D.W.F. Brilman Universiteit Twente

The research reported in this thesis was sponsored by Shell Global Solu-tions International.

c© M.J. Tuinier, Eindhoven, The Netherlands, 2011No part of this work may be reproduced in any form by print, photocopyor any other means without written permission from the author.

Publisher: Ipskamp Drukkers B.V., P.O. Box 333, 7500 AH, Enschede,The Netherlands.

A catalogue record is available from the Eindhoven University of Techno-logy Library.

ISBN: 978-90-386-2900-1

Page 4: Novel process concept for cryogenic CO2 capture

Novel Process Concept for Cryogenic CO2 Capture

PROEFSCHRIFT

ter verkrijging van de graad van doctor aan deTechnische Universiteit Eindhoven, op gezag van de

rector magnificus, prof.dr.ir. C.J. van Duijn, voor eencommissie aangewezen door het College voor

Promoties in het openbaar te verdedigenop donderdag 24 november 2011 om 16.00 uur

door

Martin Jan Tuinier

geboren te Wijhe

Page 5: Novel process concept for cryogenic CO2 capture

Dit proefschrift is goedgekeurd door de promotor:

prof.dr.ir. M. van Sint Annaland

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vi

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Contents

Summary 1

1 Introduction 51.1 Climate change . . . . . . . . . . . . . . . . . . . . . . . . . . . 51.2 Carbon capture and storage . . . . . . . . . . . . . . . . . . . 81.3 This thesis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15

2 Cryogenic packed bed process concept 192.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . 202.2 The process concept . . . . . . . . . . . . . . . . . . . . . . . . 202.3 Detailed numerical model . . . . . . . . . . . . . . . . . . . . . 242.4 Simplified model: Sharp front approach . . . . . . . . . . . . 312.5 Process analysis . . . . . . . . . . . . . . . . . . . . . . . . . . 372.6 Discussion and conclusions . . . . . . . . . . . . . . . . . . . 42

3 Experimental demonstration of the concept 453.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . 463.2 Experimental setup and procedure . . . . . . . . . . . . . . . 463.3 Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 483.4 Simulations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 523.5 Discussion and conclusions . . . . . . . . . . . . . . . . . . . 56

4 Mass deposition rates of carbon dioxide 594.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . 604.2 Experimental . . . . . . . . . . . . . . . . . . . . . . . . . . . . 614.3 Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 674.4 Development of a frost growth model . . . . . . . . . . . . . . 734.5 Discussion and conclusions . . . . . . . . . . . . . . . . . . . 87

vii

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viii Contents

5 Experimental demonstration in a pilot scale setup 935.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . 945.2 Experimental . . . . . . . . . . . . . . . . . . . . . . . . . . . . 945.3 Experimental results . . . . . . . . . . . . . . . . . . . . . . . 985.4 Simulations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1015.5 Discussion and conclusions . . . . . . . . . . . . . . . . . . . 107

6 Techno-economic evaluation 1116.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1126.2 Process evaluation . . . . . . . . . . . . . . . . . . . . . . . . . 1136.3 Comparison with absorption and membrane technology . . . 1246.4 Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 131

7 Biogas purification 1337.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1347.2 Adsorption . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1357.3 Cryogenic packed bed concept . . . . . . . . . . . . . . . . . . 1367.4 Adsorption versus cryogenic packed bed concept . . . . . . . 1427.5 Hydrogen sulfide removal . . . . . . . . . . . . . . . . . . . . . 1467.6 Discussion and conclusions . . . . . . . . . . . . . . . . . . . 146

8 Epilogue and outlook 1498.1 Important aspects for future development . . . . . . . . . . . 1508.2 Future of the proposed concept . . . . . . . . . . . . . . . . . 152

Bibliography 154

List of publications 163

Curriculum Vitae 165

Dankwoord 167

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Summary

Carbon capture and storage (CCS) is generally considered as one of thenecessary methods to mitigate anthropogenic CO2 emissions to combatclimate change. The costs of CCS can for a large extent be attributedto the capture process. Several post-combustion CO2 capture processeshave been developed, such as scrubbing, membrane processes and pres-sure swing adsorption. Amine scrubbing is currently the state of theart technology, in which CO2 is being removed by contacting the fluegas with a solvent in an absorber. Regeneration is carried out by heat-ing the loaded solvent in a stripping column. The main disadvantagesof this process are the energy costs related to the regeneration step andsolvent losses due to degradation. A promising novel option is to freezeout (desublimate) CO2 from flue gases using cryogenically cooled sur-faces. High cooling costs could be minimized by exploiting the cold dutyavailable at Liquefied Natural Gas (LNG) regasification sites. No stan-dard process equipment is available to deal with separations based ondesublimation. Therefore, a novel process concept has been developedand investigated in this dissertation, based on the periodic operation ofcryogenically cooled packed beds.

When feeding a flue gas to a previously cryogenically refrigeratedpacked bed, CO2 will freeze onto the packing surface, while permanentgases such as N2 pass through the bed unaltered. The amount of CO2

depositing onto the packing reaches an equilibrium value, because theamount of cold energy stored in the packing is limited. Therefore, plug-ging of the bed is intrinsically circumvented. A front of desublimatingCO2 will move through the bed, until breakthrough is observed. At thatpoint, the bed is switched to a recovery step, in which all previously de-posited CO2 will be removed by recycling gaseous CO2 through the bed.The energy required for the sublimation of CO2 can be provided to the

1

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2 Summary

bed by feeding the flue gas at elevated temperatures during the capturestep. A process cycle is finally finished with a cooling step, in which thebed is again refrigerated to its initial temperature. The proposed processconcept has several advantages: simple and low cost equipment can beused, large pressure drops can be avoided, deep CO2 removal is possibleand CO2, H2O and other impurities can be separated simultaneously.

The development in time of axial temperature, gas concentrationand mass deposition profiles in the packed beds during the differentprocess steps can be well described using a one-dimensional, pseudo-homogeneous, axially dispersed plug flow model, in which the mass andenergy balances are solved simultaneously using an advanced numericalscheme. When assuming that no axial dispersion and mass depositionrate limitations occur during the process, the fronts which are developedare very well defined (sharp). Based on this assumption, the process canbe well described with a simplified model (the ‘sharp front’ approach).With basic mass and energy conversation laws, the axial temperature,gas composition, and mass deposition profiles and front velocities can becalculated very fast using this model, making it a perfect tool for designand evaluation of the process. In the limit of negligible dispersion in thedetailed numerical model, the solution converges to the profiles predictedby the sharp front approach. Outcomes of the sharp front approach showthat the specific cooling duty required to capture a certain amount of CO2

increases for lower CO2 concentrations in the feed gas and for higher ini-tial bed temperatures.

A small scale experimental setup has been designed and constructedto measure axial temperature profiles and CO2 concentrations at theoutlet during the capture step. Experiments have been carried out forN2/CO2 and N2/CO2/H2O mixtures. The results showed that a goodseparation between the single components is possible. The experimentalfindings have been compared to results calculated using the numericalmodel. The front velocities and therefore CO2 breakthrough times and thetemperature profiles for different CO2 and H2O concentrations in the fluegas and initial bed temperatures are very well predicted by the developedmodel.

Expressions for mass deposition rates of CO2 are required to accu-rately describe the process. No information is available in literature,therefore a dedicated experimental setup has been designed and con-structed to measure mass deposition rates for different (gas and sur-face) temperatures, pressures and compositions. It is shown that the

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Summary 3

rate of desublimation is influenced by the thickness of the CO2 layer de-posited onto the cooled surface, indicating the importance of heat trans-fer through the frost layer. Furthermore, it is found that the presence ofN2 in the gas phase has a large effect on the desublimation rates, whichindicates the presence of mass transfer limitations. A model has been de-veloped to describe the observed behavior. The frost growth process hasbeen described as a moving boundary problem, in which both mass andheat transfer are taken into account. Based on the experimental results,expressions have been derived to describe the density and heat conduc-tivity of the frost layer. Using these expressions, the model is well able topredict the experimental results. It is shown that under the conditionsas prevailing in the packed beds, mass deposition rates are mainly deter-mined by mass transfer from the gas bulk phase to the packing surface.

The small scale experimental setup has been used to study the cap-ture step. In order to demonstrate the entire process cycle includingthe cooling and recovery step, a larger pilot setup has been constructed,containing three beds operated continuously. Test runs of more than10 hours showed that it is indeed possible to continuously capture CO2

with the proposed concept. Radial temperature differences were observedin the beds, which could be attributed to the influence of the steel wallvia simulations with the numerical model after including an additionalenergy balance for the wall.

In a techno-economic evaluation the influence of several process pa-rameters has been investigated; lower initial bed temperatures and higherCO2 concentrations in the feed result in more efficient use of the bed vol-ume. The pressure drop over the system plays an important role in theprocess economics, due to the high flow rates required in the process.The cryogenic concept has been compared to two competing technolo-gies: amine scrubbing and membrane separation. The results show thatthe preferred technology highly depends on assumptions related to theavailability of utilities. The novel cryogenic capture process can competewith other technologies, provided that cold duty is available at low cost.

An alternative promising application for the proposed technology isbiogas purification. By operating the capture step at elevated pressures,it is possible to remove CO2 during the regeneration step very efficientlyby reducing the pressure. The process performance has been comparedwith vacuum pressure swing adsorption technology. The required bedsfor treating a gas mixture containing 45 vol.% CO2 and 55 vol.% CH4 are

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4 Summary

eight times smaller for the cryogenic packed bed concept and the energyconsumption is 22% lower.

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1Introduction

1.1 Climate change

There is a growing worldwide awareness of the fact that the earth’s sur-face temperatures are changing globally. Although the climate of ourplanet has been altering continuously during its history, the currentchanges are taking place with unprecedented pace and are expected tohave dramatic consequences for human kind (IPCC, 2007). In the firstplace due to rising sea levels, which is threatening land at lower levels,but also of the expected negative impact on agriculture and fresh watersupply. Since the industrial revolution in the late 19th century, fossilfuels started to play an important role in our energy supply for trans-portation, heating and electricity. The combustion of fossil fuels resultsin large amounts of CO2, which are emitted into the atmosphere. Theincrease in CO2 concentrations coincides with the increase in global tem-peratures. There is a general consensus among most scientists that therise in CO2 concentrations is responsible for the observed increase intemperatures. In 1988 the Intergovernmental Panel on Climate Change(IPCC) was established in order to evaluate the risks of climate changecaused by human activity. Based on their observations, it is expected

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6 Introduction

that global temperatures will keep on rising in the next century, as illus-trated in Fig. 1.1. In order to prevent or at least minimize further temper-ature increases, it is necessary to reduce anthropogenic CO2 emissions.Not only for health and safety reasons, but also for economic reasons.In a study on the economical effects of climate change by Stern (2007)it is stated that the costs of mitigating climate change can be limited toaround 1% of global GDP each year. Doing nothing (‘business as usual’)and facing the consequences of climate change will be equivalent to los-ing 5% to possibly 20% of global GDP each year. Therefore, immediateaction to reduce CO2 emissions is essential.

This can be achieved in several ways, in the first place by improvingefficiencies. Developments in the automobile industry for example areleading to more and more economic engines, consuming less fuel. Theefficiency of power plants is also increasing, and chemical industry isable to save energy by for example heat integration. These developmentswill contribute to CO2 emission reductions. However, it is not expectedthat these efficiency improvements will be sufficient to bring down ouremissions to acceptable levels, mainly because of increased energy de-mands by developing countries such as China and India. Therefore moremeasures are required. A key measure is to switch our energy supplyto renewable energy sources such as biomass, solar and wind energy.Not only to reduce our CO2 emissions, but also to bring down our de-pendency on scarce fossil fuels. However, at this point power supplyby renewable energy sources is still under development and is not yetcompetitive with conventional power generation based on fossil fuels. Athird possible route to emission reductions is nuclear power, but safetyissues and nuclear waste disposal are causing moral and political con-cerns. Due to the above mentioned reasons, it is expected that fossilfuels will continue to play a significant role in our energy supply for thenext decades (U.S. Department of Energy, 2010), as shown in Fig. 1.2.It is therefore considered to be necessary to introduce Carbon Captureand Storage (CCS) to mitigate anthropogenic CO2 emissions as a mid-term solution until a full transition to energy supply by renewables canbe realized.

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1.1 Climate change 7

Figure 1.1: Projections of surface temperatures for the period 2020-2029 and2090-2099 relative to temperatures in 1980-1999, based on three scenarios(IPCC, 2007).

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8 Introduction

1990 2000 2010 2020 20300

100

200

300En

ergy

use

[qua

drill

ion

Btu

]

Year

Liquids Renewables Coal Nuclear Natural Gas

Figure 1.2: World marketed energy use by fuel type from 1990 to 2035(U.S. Department of Energy, 2010).

1.2 Carbon capture and storage

The goal of CCS is to remove CO2 from flue gases and to store it forthe long term. This process is schematically represented in Fig. 1.3.Before discussing the possible capture processes, first attention is paidto storage options.

1.2.1 Storage

It is proposed in literature (Leitner, 1995) to reuse captured CO2 as a rawmaterial for chemical syntheses. However, the CO2 molecule is thermo-dynamically very stable and relatively unreactive. Therefore severe andcostly process conditions are required. Another option is to reuse CO2 ingreenhouses, which is currently being applied by Shell at their refineryin Pernis, The Netherlands (Shell, 2011). Furthermore it is proposed toreuse CO2 as a feedstock for growing algae (Brennan and Owende, 2010).Although above mentioned options could play a role in some cases, thetotal amount of CO2 available is in general exceeding the demand by far.

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1.2 Carbon capture and storage 9

Therefore, it is also necessary to store CO2 instead of reusing it. Severaloptions are available: mineralization of CO2 (Seifritz, 1990) and storagein geological formations or in oceans. The problem of mineralization isthat reaction rates are low, limiting large scale application. The effects ofocean storage on marine life is still uncertain and therefore most atten-tion is paid to geological storage. CO2 could be stored in depleted oil orgas reservoirs or saline formations (IPCC, 2005). CO2 injection could alsobe used for Enhanced Oil Recovery (EOR) or Enhanced Coal Bed MethaneRecovery (ECBM). Experience with CO2 injection in for example the Sleip-ner project in Norway, indicate that storing CO2 in geological formationsis a feasible option to mitigate CO2 emissions (IPCC, 2005).

1.2.2 Capture

Fossil fuels are normally combusted using air. The flue gas is thereforecomposed of a large amount of N2 and 5-20 vol.% CO2. Furthermore itcontains H2O and impurities such as sulphur and nitrous oxides, de-pending on the feedstock and process. Compressing and storing the en-tire flue gas, including N2 will be too costly. Therefore it is necessaryto obtain CO2 in purified form first, before it can be stored in geologicalformations. About 75% of the costs involved in CCS are associated withthe capture step (Ebner and Ritter (2009)) and therefore many researchprojects focus on development or optimization of capture technologies.

CO2 capture technologies are often classified into oxyfuel, pre- andpost-combustion processes, which are schematically represented in Fig.1.4. In oxyfuel processes, fossil fuels are combusted using pure oxy-gen, circumventing dilution of CO2 with N2. Disadvantage is that anenergy intensive air separation unit is required to obtain pure O2, al-though this could be avoided using chemical looping combustion (seee.g. Ishida and Jin (1994), Noorman et al. (2007)).

In pre-combustion processes fossil fuels are first converted into H2

and CO2 (via (autothermal) reforming or partial oxidation and water-gas-shift), CO2 is subsequently captured and H2 is fed to the combustionchamber or fuel cell. Advantage is that the separation of CO2 and H2

can be carried out at a high pressure, resulting in a high driving forcefor separation. Disadvantage of pre-combustion is that these processesare applicable to mainly new plants and can therefore not be applied toalready operating facilities.

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10 Introduction

Figure 1.3: Schematic overview of CCS process (IPCC, 2005).

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1.2 Carbon capture and storage 11

Figure 1.4: Pathways to CO2 capture, a. conventional combustion process with-out capture, b. post-combustion, c. pre-combustion, d. oxy-fuel.

Post-combustion processes are based on capturing CO2 from fluegases from conventional air fired combustion processes. Disadvantageis that CO2 is dilute and at low pressures, reducing the driving force forseparation. However, this technology can be retrofitted to already operat-ing power plants and industries. For this reason post-combustion is con-sidered the most realistic technology on the short term, even though theefficiency of the alternatives could be higher (Kvamsdal et al., 2007). Sev-eral post-combustion technologies are under development, an overview isgiven below.

Scrubbing

The state-of-the-art post-combustion CO2 capture technology is scrub-bing. This technology is based on feeding the flue gas to an absorberto selectively absorb CO2. In a desorber column the solvent is strippedfrom CO2 by changing temperature and/or pressure. Amines, such asMono Ethanol Amine (MEA), are the most commonly used chemical sol-vents. Since 1930 amine scrubbing has already been used to remove CO2

from natural gas and hydrogen (Rochelle, 2009). Drawbacks of amine

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12 Introduction

scrubbing are high energy demands, mainly due to the stripping stepat elevated temperature, intolerance of solvents to impurities (SOx, NOx,O2) and equipment corrosion. Many novel amine-based solvents are un-der development with higher CO2 solubility, faster absorption kineticsand at the same time a better tolerance for impurities, (see e.g. Notz et al.(2007)). Also chemical solvents based on amino acids (Aronu et al., 2010),chilled ammonia (Darde et al., 2010) or carbonates (Figueroa et al., 2008)are under consideration. Instead of chemical solvents, physical solvents,e.g. Selexol (dimethyl ether of polyethylene glycol) or Rectisol (chilledmethanol) could be used for CO2 removal. However, physical solvents areconsidered inferior to chemical solvents because a high partial pressureof CO2 and low temperature is required, both of which are not the casein flue gas treatment (Notz et al., 2011).

Membranes

Three types of membrane systems are under consideration for post-combustion CO2 capture: gas separation membranes, membrane absorp-tion and facilitated transport membranes. Polymeric (Powell and Qiao,2006) and ceramic (Bredesen et al., 2004) membranes could both beused as gas separation membranes. Ceramic membranes require hightemperatures in general and will therefore have more potential for pre-combustion capture. The intolerance of polymeric membranes to im-purities but also H2O is generally considered as a limitation to appli-cation, although in recent research novel polymeric membranes havebeen developed which are able to separate CO2 and H2O simultaneously(Reijerkerk et al., 2011).

In a membrane absorption process, membranes are installed in a reg-ular scrubbing process, to act as a contact area between the flue gas andsolvent. Therefore impurities in the gas phase will not directly contactthe solvent, reducing solvent losses. Furthermore a higher contact areaper unit volume can be created, although the additional mass transferresistance associated with the introduction of the membrane might re-duce or possibly even cancel the advantage of the increased contact area(Notz et al., 2011).

In Facilitated Transport Membranes (FTM) a carrier medium selec-tively interacts with a specific molecule. A CO2 selective FTM wouldhave much lower processing costs and improve the equilibrium drivenprocesses. However, currently FTMs have stability problems, which are

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1.2 Carbon capture and storage 13

mainly caused by evaporation of the carrier medium (Ebner and Ritter,2009).

The challenge in developing membrane systems is that two importantcriteria are inversely related: a high permeability and a high selectivity.CO2 separation from for example natural gas has already been success-fully applied commercially, due to a high partial pressure of CO2 in thefeed and therefore high driving force for permeation. However, CO2 sepa-ration from flue gases is more difficult due to the lower partial pressuresof CO2 (Ebner and Ritter, 2009). The low driving force therefore requirescompression of the flue gas or requires large membrane areas, resultingin increased operational and capital costs respectively. Despite this rea-son, some authors claim that membrane processes could compete wellwith absorption processes (Favre, 2007).

Adsorption

Separation of CO2 from gas mixtures by adsorption is based on differ-ences in interaction with the adsorbent surface. Molecular sieves or acti-vated carbons are common used as adsorbents (IPCC, 2005). The processis normally carried out in several packed beds operated in parallel. Re-generation is done by either Pressure Swing Adsorption (PSA) (Ko et al.,2005), Temperature Swing Adsorption (TSA) (Merel et al., 2006) or Elec-trical Swing Adsorption (ESA) (Grande and Rodrigues, 2008). Adsorptiontechnology is already being used for CO2/H2 separation. The energy costscould be lower than scrubbing for post-combustion CO2 capture. How-ever, the disadvantages are the low CO2 selectivity, low loading capacityfor CO2 (resulting in large beds), low adsorption rates, relatively largepressure drops over the fixed beds and the high energy demand for re-generation (Notz et al., 2011). Research focuses on improved adsorbents,such as amine-functionalized zeolites (Su et al., 2010), or Metallic Or-ganic Frameworks (MOFs) (Britt et al., 2009). Another adsorption tech-nology under investigation is calcium looping, in which the flue gas iscontacted with CaO in a fluidized bed forming CaCO3. In a connectedsecond fluidized bed regeneration is carried out at elevated temperatures(Manovic et al., 2009).

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14 Introduction

Cryogenics

Cryogenics are another option for separating CO2 from gas mixtures. Theadvantages are that no chemical ab- or adsorbents or large pressure dif-ferences are needed and that high purity products can be obtained. How-ever, cryogenic CO2 capture is not included in most (economic) compar-ison studies, as it has been considered as an unrealistic candidate forpost-combustion CO2 capture. In the first place due to expected highcooling costs, but also because it has been considered as a gas-liquidseparation (Aaron and Tsouris, 2005; Ebner and Ritter, 2009). At atmo-spheric pressures CO2 will go directly from its gas phase to its solid phase(desublimation). In order to be able to carry out the CO2 removal from fluegases as a gas-liquid separation, it is necessary to compress the gas topressures above the triple point of CO2, which is at 5.2 bar and -56.6◦Cfor pure CO2, as shown in the phase diagram of pure CO2 in Fig. 1.5.Compressing flue gases to high pressures is too energy intensive.

Expensive refrigeration can possibly be avoided when exploiting thecold duty available at Liquefied Natural Gas (LNG) regasification sites.Currently, LNG is being regasified using seawater or by using water bathswhich are heated by burning a fuel gas (Ertl et al., 2006). The globalLNG market is strongly growing (John and Robertson, 2008), thereforeintegration of LNG regasification and a cryogenic CO2 process could bebeneficial.

Clodic and Younes (2002, 2005) have developed a cryogenic CO2 cap-ture process, in which CO2 is desublimated as a solid onto surfaces ofheat exchangers which are cooled by evaporating a refrigerants blend.With calculations and experimental tests they showed that their processcould compete with other post-combustion CO2 capture processes. Themain disadvantage of their system is that the water content in the feedstream to the cooling units should be minimal in order to prevent plug-ging by ice or an unacceptably high rise in pressure drop during op-eration. Therefore, several costly steps are required to remove all watertraces from the flue gas. In addition the increasing layer of solid CO2 ontoheat exchanger surfaces during the capture cycle will adversely affect theheat transfer, reducing the process efficiency. Moreover, the costly heatexchangers have to be switched to regeneration cycles operated at a dif-ferent temperature, which should be carried out with great care to avoidexcessive mechanical stresses.

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1.3 This thesis 15

10,000

1,000

100

10

1

supercritical fluid

200

T [K]

P [

bar

]

250 300 350 400

critical point

Figure 1.5: Phase diagram of pure CO2.

1.3 This thesis

To avoid above mentioned problems, an alternative cryogenic CO2 pro-cess concept based on dynamically operated packed has been developedat the University of Twente and Eindhoven University of Technology in co-operation with Shell Global Solutions International. This process conceptwill be described and analyzed in detail in this dissertation.

When feeding a flue gas to a previously refrigerated packed bed, CO2

will freeze onto the packing surface, while permanent gasses such asN2 pass through the bed unaltered. A front of desublimating CO2 willmove through the beds, until breakthrough is observed. At that point,a bed is switched to a recovery step, to remove all previously depositedCO2. A process cycle is finally finished with a cooling step, in which thebed is again refrigerated to its initial temperature. The proposed processconcept has several advantages: simple and low cost equipment, no largepressure drops (intrinsic circumvention of plugging) and the possibilityto separate CO2, H2O and other impurities simultaneously.

Page 25: Novel process concept for cryogenic CO2 capture

16 Introduction

Chapter 2 gives a detailed explanation of this novel process concept.The dynamic behavior of the process will be described by a simplifiedmodel, based on thermodynamic equilibrium and an advanced numericalmodel. The evolution of temperature and mass deposition profiles will beshown.

In order to demonstrate the capture step of the proposed process con-cept, a small scale experimental setup was designed and constructed tomeasure axial temperature profiles and CO2 concentrations at the outletduring the capture step. Chapter 3 will give the results of the exper-iments for a wide range of conditions using different compositions andinitial bed temperatures. These experimental results are compared tosimulation results.

Efficient operation of the proposed concept requires fast CO2 desubli-mation rates. Limited information is available in literature on these rates.Chapter 4 shows the design of a dedicated setup to measure desubli-mation rates. Obtained experimental results for different temperatures,pressures and compositions will be shown and a frost growth model isdeveloped to describe the observed behavior.

In Chapter 5 it is presented how the complete process cycle includ-ing all three steps can be demonstrated. The design of a continuouslyoperated setup will be discussed and results will be shown.

Chapter 6 will give a techno-economic evaluation of the process con-cept. The influence of initial bed temperatures, CO2 concentrations andpressure drops will be investigated. Furthermore the cryogenic conceptwill be compared to two competing technologies: amine absorption andmembrane separation.

The proposed process could be applied for alternative gas separations.Chapter 7 shows how biogas purification can be carried out using thecryogenic packed bed concept. It is presented how the process can be op-erated efficiently and compares the outcomes to a competing technology:Vacuum Pressure Swing Adsorption (VPSA).

Finally an outlook is given in Chapter 8, in which the status of thetechnology and the required future research and development are dis-cussed.

Page 26: Novel process concept for cryogenic CO2 capture

1.3 This thesis 17

Acknowledgment

The author would like to thank Shell Global Solutions International fortheir financial support and involvement in the project.

Page 27: Novel process concept for cryogenic CO2 capture

18 Introduction

Page 28: Novel process concept for cryogenic CO2 capture

2Cryogenic packed bed process concept

Abstract

This chapter elucidates the novel process concept based on cryogenicallycooled dynamically operated packed beds. In order to capture CO2 froma flue gas, three process steps are required: a capture, recovery andcooling step. The dynamic behavior of the axial temperature, gas phaseconcentrations and mass deposition profiles during these steps have beeninvestigated with two different models: An advanced numerical modelincluding effects such as axial dispersion in the beds, and a simplifiedmodel in which developed fronts are assumed to be perfectly well defined,referred to as the ’sharp front approach’. Using the advanced model, ithas been demonstrated that the simplified model can well capture themost salient process characteristics even quantitatively, in particular forcases with low axial dispersion. The sharp front approach is an excellenttool to evaluate the influence of process parameters such as the inletcomposition and initial bed temperatures on the process performance.The outcomes of this evaluation are presented in this chapter. Lowerinitial bed temperatures result in more CO2 mass deposition per unit ofbed volume, while the specific cooling duty (required cooling duty per

Page 29: Novel process concept for cryogenic CO2 capture

20 Cryogenic packed bed process concept

unit of mass of CO2 captured) reaches a constant value at low initial bedtemperatures. Lower CO2 concentrations in the flue gas result in lessmass deposition per unit of bed volume and increased required coolingduty.1

2.1 Introduction

This chapter starts with a qualitative description of the different processsteps involved: the capture, recovery and cooling step. During thesedifferent steps axial temperature, gas phase concentration and mass de-position profiles develop in time. An advanced numerical model based ona pseudo-homogeneous one-dimensional plug flow model with superim-posed axial dispersion is developed, which is able to describe the dynamicbehavior of the process. Subsequently, a simplified, but computationallymuch easier and faster model will be presented, in which fronts are as-sumed to be perfectly well defined (sharp front approach). These twomodels will be compared, and the chapter concludes with giving an eval-uation of the influence of several process parameters, such as the initialbed temperature and inlet composition.

2.2 The process concept

In this section the process concept is described in detail, where the fluegas is represented as a mixture of N2, CO2 and H2O to simplify the de-scription. Continuous separation of these components can be obtainedwhen three packed beds are operated in parallel in three different steps:a capture, recovery and cooling step. These three steps are discussedconsecutively below focusing on the evolution of axial temperature andmass deposition profiles (see Fig. 2.1).

2.2.1 Capture step

When a gas mixture consisting of N2, CO2 and H2O is being fed at arelatively high temperature Tc,in to an initially cryogenically refrigeratedpacked bed (at T0), an effective separation between these components can

1This chapter is based on the papers: Tuinier et al., Chem. Eng. Sci. , 65(1), 114-119,2010 and Tuinier et al., Int. J. Greenhouse Gas Control, 5(4), 694-701, 2011.

Page 30: Novel process concept for cryogenic CO2 capture

2.2 The process concept 21

be accomplished, due to differences in dew and sublimation points. Thegas mixture will cool and the packing material will heat, until H2O startsto condense at the packing surface. A certain amount of H2O per volumeof packing material (indicated as mH

2O in Fig. 2.1a) will condense, until

a local equilibrium is reached (at a temperature TH2O). Actually a very

small part of the H2O at the front will be frozen to ice, but simulationshave revealed that this is a very small part of the H2O and has negligibleinfluence on the resulting axial temperature and mass deposition pro-files. The cold energy stored in the packing will be consumed and a frontof condensing H2O will move through the bed towards the outlet of thebed. At the same time, previously condensed H2O will evaporate due tothe incoming relatively hot gas mixture. Therefore, two fronts of evaporat-ing and condensing water will move through the bed, with a faster movingcondensing front. After all water being condensed, the gas mixture willbe cooled further until CO2 starts to change phase. At atmospheric pres-sure CO2 will desublimate directly from gas to solid, and therefore solidCO2 is deposited onto the packing surface. Similar as for H2O, two CO2

fronts will move through the bed: an evaporation and a desublimationfront. Again an equilibrium is reached, a certain amount of CO2 (mCO

2)

is deposited at the packing surface at a temperature of TCO2. Note that in

this way an effective separation between CO2 and H2O is accomplished.N2 will not undergo any phase change (as long as T0 is not chosen toolow) and will therefore move through the bed unaffected. When the CO2

desublimation front reaches the end of the bed, CO2 may break throughand the bed should be switched to a recovery step just before that.

Page 31: Novel process concept for cryogenic CO2 capture

22

Cry

ogen

icp

ack

ed

bed

pro

cess

con

cep

t

2

2

T0

TCO

TH O

t1

t2

Temperature

Axial position

Tc,in

t1 H

2O t

2 H

2O

t1 CO

2 t

2 CO

2

Mas

s dep

ositi

on

Axial position

mH O2

2m

CO

(a) Capture step

Tr,in

t0

t1

Temperature

Axial position

Tc,in

TH O

2

TCO2

TCO *2

t0 H

2O

t0 CO

2

t1 CO

2

t1 CO

2

Mas

s dep

ositi

on

Axial position

mH O2

mCO2

mCO *2

(b) Recovery step

t0

t1

t2

Temperature

Axial position

TH O

Tc,in

Tr,in

T0

2

t0 H

2O

Mas

s dep

ositi

on

Axial position

mH O2

(c) Cooling step

Figure 2.1: Schematic axial temperature and corresponding mass deposition profiles for the capture (a),recovery (b) and cooling (c) step respectively.

Page 32: Novel process concept for cryogenic CO2 capture

2.2 The process concept 23

2.2.2 CO2 recovery step

The first zone of the bed has been heated to Tc,in during the capture step.This heat is used in the recovery step to evaporate the condensed H2Oand desublimated CO2. A gas flow consisting of pure CO2 is fed to thebed. When feeding a pure CO2 gas flow at a temperature Tr,in to thepacked bed, the gas will be heated up to Tc,in and all fronts will movethrough the bed, as illustrated in Fig. 2.1b. However, during the initialperiod of the recovery step the ingoing CO2 will deposit onto the packing.Due to the increase in CO2 partial pressure compared to the capturestep, more CO2 is able to desublimate at the packing surface (from mCO

2

to m∗

CO2

), and the bed temperature will slightly increase to T ∗

CO2

. Pure CO2

is obtained at the outlet of the bed after this new equilibrium is reached.Part of the outgoing CO2 should be compressed for transportation andsequestration, while the other part can be used to recycle to the inlet ofthe bed at a temperature Tr,in, which is slightly higher than T ∗

CO2

due to

the heat production associated with the compression in the recycle blowerand some unavoidable heat leaks. When all CO2 has been recovered, thebed is switched to a step in which H2O is removed and the bed is cooledsimultaneously.

Alternatively, the deposited CO2 could be recovered as a liquid, avoid-ing expensive compression costs required for transportation and storage.This could be accomplished by closing the valves connected to the bedand by introducing heat into the bed. CO2 evaporation occurs and pres-sure builds up until the system reaches the triple point of pure CO2 andliquid CO2 will be formed. The drawbacks of this process alternative arethat pressure vessels are required and that heat should be introducedinto the bed for example by means of internal tubes. Both measures willresult in a significant increase in capital costs. Furthermore not all liquidCO2 might be recovered from the packing, due to the static liquid hold upin the bed. This process option is not further explored in this work.

2.2.3 H2O recovery and cooling step

In the last step, the bed is cooled down using a gas flow refrigerated be-fore to temperature T0. The cleaned flue gas can be used for this purpose.Cooling can be performed using a cryogenic refrigerator or by evaporatingLNG. H2O is evaporated and removed from the bed during the first periodof the cooling step. The N2/H2O mixture can be released to the atmo-

Page 33: Novel process concept for cryogenic CO2 capture

24 Cryogenic packed bed process concept

sphere and when all H2O is recovered, the outgoing flow can be recycledto the inlet of the bed, via a cooler. Temperature and mass depositionprofiles are shown in Fig. 2.1c. It should be noted that it is not requiredto cool the entire bed to T0. The last zone can be kept at Tr,in, as duringthe capture step this last part will be cooled down by the cleaned flue gas.

2.3 Detailed numerical model

2.3.1 Model description

The prevailing heat and mass transfer processes in the periodically oper-ated packed beds have been investigated with a pseudo-homogeneousone-dimensional plug flow model with superimposed axial dispersion.The modeling was based on the following main assumptions:

• It is assumed that heat losses to the environment are small (i.e.adiabatic operation) and additionally that a uniform velocity pro-file exists in the absence of radial temperature and concentrationgradients allowing the consideration of the axial temperature andconcentration profiles only;

• Possible heat transfer limitations between the solid packing and thebulk of the gas phase are accounted for via effective axial heat dis-persion (pseudo-homogeneous model);

• The rate of mass deposition and sublimation of CO2 was assumedto be proportional to the local deviation from the phase equilibrium,taking a reasonably short equilibration time constant (g) of 1·10-6

s/m, which was assumed independent of temperature. The rate ofsublimation of previously deposited CO2 was assumed to approacha first order dependency on the mass deposition when this massdeposition approached zero.

The mass and energy conservation equations have been listed in Table2.1. The constitutive equations for the transport parameters and themass deposition rate have been summarized in Table 2.2 and 2.3, re-spectively. The gas phase (mixture) properties have been computed ac-cording to Reid et al. (1987), using the pure component data supplied byDaubert and Danner (1985). Uniform initial temperature profiles weretaken without any mass deposited onto the solid packing, where the gas

Page 34: Novel process concept for cryogenic CO2 capture

2.3 Detailed numerical model 25

phase in the bed was initially N2. Furthermore, the usual Danckwerts-type boundary conditions were applied at the inlet and outlet of the beds.

The system of strongly non-linear, coupled partial differential equa-tions was solved using a very efficient finite volume discretization tech-nique, using a second order SDIRK (Singly Diagonally Implicit Runge-Kutta) scheme for the accumulation terms, an explicit 5th order WENO(Weighted Essentially Non-Oscillatory) scheme for the convection terms(with implicit first order upwind treatment using the deferred correc-tion method), second order standard implicit central discretization forthe dispersion terms and the standard Newton-Raphson technique forthe linearly-implicit treatment of the source terms. Moreover, automatictime step adaptation and local grid refinement procedures have been im-plemented, making effective use of the WENO smoothness indicators andinterpolation polynomials (Smit et al., 2005). The steep temperature andmass deposition gradients in combination with the strongly non-linearsublimation kinetics require a very efficient and stable numerical imple-mentation using higher order implicit schemes.

2.3.2 Simulation results

Simulations have been carried out for all three process steps. The usedbed properties and process conditions are listed in Table 2.4 and 2.5 re-spectively. A stainless steel monolithic structure was chosen as packingmaterial, because axial dispersion and pressure drop are minimal for thistype of packing material while the volumetric heat capacity is relativelyhigh. The axial temperature and mass deposition profiles during the cap-ture step are shown in Fig. 2.2a and 2.2d respectively. At the choseninitial bed temperature of -140◦C, more than 99% of CO2 is recovered.After 600 seconds the CO2 desublimation front reaches the end of thebed and the capture cycle should be stopped. The conditions at 600 sec-onds are used as initial conditions for the simulation of the recovery step.Fig. 2.2b and 2.2e show that during the recovery step extra CO2 will bedeposited onto the packing surface, as explained in section 2.2.2, andthat all deposited CO2 is removed after again 600 seconds. Now the dataat the end of the recovery step are used as initial conditions for the cool-ing step. A refrigerated N2 flow is being fed to the bed and profiles willdevelop as illustrated in Fig. 2.2c and 2.2f. Note that not the entire bedis cooled down, the last zone is cooled down during the capture step. The

Page 35: Novel process concept for cryogenic CO2 capture

26 Cryogenic packed bed process concept

Table 2.1: Model equations for the 1-D pseudohomogeneous model.

Component mass balances for the gas phase:

εgρg∂ωi,g

∂t= −ρgvg

∂ωi,g

∂z + ∂∂z

(

ρgDeff∂ωi,g

∂z

)

−m′′

i as + ωi,g

nc∑

i=1

m′′

i as

Component mass balance for the solid phase:

∂mi

∂t= m′′

i as

Total continuity equation for the gas phase:

∂ (εgρg)

∂t= −

∂ (ρgvg)

∂z−

nc∑

i=1

m′′

i as

Energy balance (gas and solid phase):

(εgρgCp,g + ρs(1− εg)Cp,s)∂T

∂t= −ρgvgCp,g

∂T∂z + ∂

∂z

(

λeff∂T∂z

)

nc∑

i=1

m′′

i as∆Hi

Pressure drop over packing:

∂P

∂z= −4

f

dh

1

2ρgv

2g with: f =

14.9

Re

1 + 0.0445RedhL

Page 36: Novel process concept for cryogenic CO2 capture

2.3 Detailed numerical model 27

Table 2.2: Heat and mass transfer coefficients for a monolith packing.

Effective axial heat dispersion (Vortmeyer and Schaefer, 1974):

λeff = (1− εg)λs +

(

ρgvgCp,g

εg

)21

αg,sas

Gas to solid heat transfer coefficient (Hawthorn, 1974):

αg,s =λg

dh2.978

(

1 + 0.095RePrdhLc

)0.45

with: Re =ρgvgdhηgεg

Axial mass dispersion (Cybulski and Moulijn, 1994):

Dax = Deff ,i +v2gd

2h,c

192Deff ,iwith: Deff ,i =

1nc∑

j=1

yj,gDi,j

Table 2.3: Mass deposition rate.

Mass deposition rate:

m′′

i =

{

g (yi,sP − P σi ) if yi,sP ≥ P σ

i

g (yi,sP − P σi )

mi

mi+0.1 if yi,sP < P σi

Gas-solid equilibrium:

P σCO2

(T ) = exp

(

10.257−3082.7

T+ 4.08 lnT − 2.2658 · 10−2T

)

∆HsubCO2

= 5.682 · 105J · kg−1

Page 37: Novel process concept for cryogenic CO2 capture

28 Cryogenic packed bed process concept

Table 2.4: Bed properties used in the numerical study.

Length bed [m] 8Diameter bed [m] 3Packing type Steel monolithSolids density [kg ·m−3] 7750Channel diameter [m] 6.76 · 10−4

Wall thickness [m] 1.34 · 10−4

Porosity [-] 0.7Surface area [m2 ·m−3] 4124

Heat capacity [J · kg−1 ·K−1]5

i=0

ciTi

with:c0 −203.75c1 6.4335c2 −2.4320 · 10−2

c3 4.6266 · 10−5

c4 −4.2721 · 10−8

c5 1.5296 · 10−11

results show that CO2 and H2O capture can be integrated in one singlebed. However, the temperature profile during the recovery step showsthat the heat stored in the first zone during the capture step is only justsufficient to remove H2O from the bed. The hot zone is moved throughthe bed, but due to axial heat dispersion this hot zone will be spread outover the bed, which is well visible in Fig. 2.2b. When feeding the gasmixture at realistic flue gas temperatures (which are generally lower than250◦C) during the capture step, not sufficient heat is stored in the pack-ing to evaporate previously condensed water again. A possibility wouldbe to introduce extra heat into the bed in the initial period of the recoverystep. However, more practical is to carry out the H2O capture step in aseparate smaller bed, which can be cooled down to temperatures muchhigher than the initial bed temperature of the CO2 capture bed.

Page 38: Novel process concept for cryogenic CO2 capture

2.3 Detailed numerical model 29

Table 2.5: Conditions used in the numerical study.

Capture Recovery CoolingTin [◦C] 250 −70 −140Φin [kg · s-1] 20 100 70yCO

2,in [-] 0.1 1.0 0.0

yH2O,in [-] 0.01 0.0 0.0

Page 39: Novel process concept for cryogenic CO2 capture

30

Cry

ogen

icp

ack

ed

bed

pro

cess

con

cep

t

0 2 4 6 8-150

-100

-50

0

50

100

150

200

250

300

0 s 300 s 600 s

Tem

pera

ture

[°C

]

Axial position [m]

(a)

0 2 4 6 8-150

-100

-50

0

50

100

150

200

250

300

0 s 300 s 600 s

Tem

pera

ture

[°C

]

Axial position [m]

(b)

0 2 4 6 8-150

-100

-50

0

50

100

150

200

250

300

0 s 300 s 600 s

Tem

pera

ture

[°C

]

Axial position [m]

(c)

0 2 4 6 80

20

40

60

80

100

H2O CO2 300 s H2O CO2 600 s

Mas

s dep

ositi

on [k

g·m

-3]

Axial position [m]

(d)

0 2 4 6 80

20

40

60

80

100

H2O CO2 0 s H2O CO2 300 s H2O 600 s

Mas

s dep

ositi

on [k

g·m

-3]

Axial position [m]

(e)

0 2 4 6 80

20

40

60

80

100

H2O 0 s

Mas

s dep

ositi

on [k

g·m

-3]

Axial position [m]

(f)

Figure 2.2: Simulated axial temperature (a - c) and mass deposition (d - f) profiles for the capture, recoveryand cooling step. Bed properties and operating conditions can be found in Table 2.4 and 2.5 respectively.

Page 40: Novel process concept for cryogenic CO2 capture

2.4 Simplified model: Sharp front approach 31

2.4 Simplified model: Sharp front approach

The process concept can be described by the advanced numerical modelas detailed in the previous section. However, when assuming that thefronts which are formed during the different process steps are perfectlywell defined (sharp), a simplified and relatively easy to solve and fastmodel can be developed, referred to as the ‘sharp front approach’. In thefirst place this approach provides a very useful tool to quickly investi-gate the influence of process parameters on process behavior. Further-more it can be used in conceptual design studies. This section describesthis sharp front approach and compares the outcomes with the moreadvanced numerical model. Finally, the influences of several process pa-rameters are studied.

2.4.1 Model description

Capture step

The model is derived for capturing a component i from a binary gas mix-ture consisting of components i and j (but could be easily extended tomulticomponent mixtures). When feeding this mixture to a refrigeratedbed, two fronts are formed: a ’frost’ and a ’defrost’ front as depicted inFig. 2.3a. The defrost front moves from zd,1 to zd,2 during a time period∆t. The mass of i evaporated is equal to the distance the front movedmultiplied with the bed cross sectional area A and the amount of massdeposited per unit of bed volume mi. Due to this evaporation of previouslydeposited i, the mass flow of i after the defrost front (Φ1ωi1 ) is equal to theinlet mass flow (Φ2ωi2 ) plus the amount of i evaporated. This results inthe following component mass balance:

Φ1ωi1 = Φ2ωi2 +Amivd (2.1)

In which the front velocity vd is defined as:

vd =(zd,2 − zd,1)

∆t(2.2)

Due to the evaporation of previously frosted i, the mass fraction afterthe first front (ωi1 ) will be higher than the inlet mass fraction (ωi2 ). Alsoan overall mass balance can be formulated. The total mass flow after

Page 41: Novel process concept for cryogenic CO2 capture

32 Cryogenic packed bed process concept

(a) (b)

Figure 2.3: Axial temperature, mass deposition and gas concentration profilesused in the derivation of the sharp front approach for the capture step (a) andrecovery step (b).

Page 42: Novel process concept for cryogenic CO2 capture

2.4 Simplified model: Sharp front approach 33

the defrost front (Φ1) is equal to the inlet flow (Φ2) plus the amount ofcomponent i evaporated:

Φ1 = Φ2 +Amivd (2.3)

At the frost front, i is deposited onto the packing surface, thereforethe outlet mass flow of component i is equal to the inlet flow minus theamount of i deposited in time ∆t, which can be written as:

Φ0ωi0 = Φ1ωi1 −Amivf (2.4)

The velocity of this frost front vf is described as:

vf =(zf,2 − zf,1)

∆t(2.5)

Again, also an overall mass balance can be formulated. The totalmass flow after the frost front (Φ0) is equal to Φ1 minus the amount ofcomponent i deposited onto the packing surface:

Φ0 = Φ1 −Amivf (2.6)

For both fronts also energy balances can be formulated. At the defrostfront heat is required to heat up the packing material from the satura-tion temperature (T1) to the inlet temperature (T2). Furthermore heat isconsumed due to the (endothermic) sublimation of component i. The re-quired energy is provided by the feed gas, which is cooled down from T2

to T1. The energy balance results in:

Avd [ρsCp,s (T2 − T1) +mi∆Hi] = Φ2 (T2 − T1) (ωi2Cp,i + ωj2Cp,j) (2.7)

At the frost front the exothermic desublimation is included in the en-ergy balance, resulting in:

Avf [ρsCp,s (T1 − T0)−mi∆Hi] = Φ0 (T1 − T0) (ωi0Cp,i + ωj0Cp,j) (2.8)

So, three balances have been derived for each front (component mass,overall mass and energy balances) giving a total of six balances. Thereare eight unknowns (T1, Φ0, Φ1, ωi0 , ωi1 , vd, vf and mi). The two massfractions of component i after the defrost front (ωi1 ) and in the outlet (ωi0 )are related to the temperature by the phase equilibrium:

Page 43: Novel process concept for cryogenic CO2 capture

34 Cryogenic packed bed process concept

ωi =P σi (T )

Ptot

Mi

M(2.9)

in which M is the average molar weight.Finally eight equations and eight unknowns are obtained. The gas and

solid heat capacities are dependent on the temperature and are normallydescribed by polynomial correlations and the vapor pressures as func-tion of the temperature are normally described by exponential relations.Therefore, a system of non-linear equations is obtained, which can besolved by standard root seeking methods such as the Newton-Raphsontechnique.

Recovery step

During the recovery step, the bed is fed with pure component i. Threefronts will develop, as illustrated in Fig. 2.3b. Initially, additional i willdeposit onto the packing surface, due to the higher pressure of i com-pared to the capture step. Therefore, a front will be formed and movestowards the outlet with a velocity vf,r. The new amount of mass depositedper unit of bed volume is now indicated as m∗

i . The component mass, theoverall mass balance and the energy balance for this frost front are listedbelow:

Φ1,rωi1,r = Φ∗

1,rω∗

i1,r −A (m∗

i −mi) vf,r (2.10)

Φ1,r = Φ∗

1,r −A (m∗

i −mi) vf,r (2.11)

Avf,r [ρsCp,s (T∗

1 − T1)− (m∗

i −mi)∆Hi]

= Φ1,r (T∗

1 − T1) (ωi1,rCp,i + ωj1,rCp,j) (2.12)

Also a defrost front will move through the bed during the recoverystep, for which the next mass and energy balances can be formulated(Note that ω∗

i1,rand ωi2,r are both unity when the bed is recovered with

pure i):

Φ∗

1,rω∗

i1,r = Φ2,rωi2,r +Amivd,r (2.13)

Page 44: Novel process concept for cryogenic CO2 capture

2.4 Simplified model: Sharp front approach 35

Φ∗

1,r = Φ2,r +Amivd,r (2.14)

Avd,r [ρsCp,s (T2 − T ∗

1 ) +m∗

i∆Hi]

= Φ2,r (T2 − T ∗

1 ) (ωi2,rCp,i + ωj2,rCp,j) (2.15)

Similar to the capture step, a system of non-linear equations is ob-tained which can be solved using a root seeking technique. During therecovery step a third front is being formed: The front closest to the inlet(moving with velocity vr). This front is formed due to the difference ininlet temperature (T3) and the temperature of the bed at the initial zoneafter the capture step (T2). No phase change of component i takes placeat this front, therefore the front velocity can be described with:

vr =Φ3,r (ωi3,rCp,i + ωj3,rCp,j)

AρsCp,s(2.16)

When the heat capacities of the gas and solid phase are independentof the temperature, the front velocity is not a function of the temperature.

Cooling step

During the cooling step no phase change is involved. The bed is cooleddown with an inert gas, fed at temperature T0. A temperature front willmove through the bed, with a velocity:

vc =Φ0,c (ωi0,cCp,i + ωj0,cCp,j)

AρsCp,s(2.17)

All equations for the three steps for the sharp front approach havebeen summarized in Table 2.6.

2.4.2 Simulation results

The outcomes of the sharp front approach are presented in this section,and compared to the simulation results of the advanced numerical model.Fig. 2.4 shows axial temperature and mass profiles which are formed af-ter 400 seconds, when feeding a binary N2/CO2 mixture at an inlet tem-perature of 150◦C. Other conditions are equal to those listed in Table 2.4

Page 45: Novel process concept for cryogenic CO2 capture

36 Cryogenic packed bed process concept

Table 2.6: Equations for sharp front approach.*

Capture step:

Defrost front:Φ1ωi1 = Φ2ωi2 +AmivdΦ1 = Φ2 +AmivdAvd [ρsCp,s (T2 − T1) +mi∆Hi] = Φ2 (T2 − T1) (ωi2Cp,i + ωj2Cp,j)

Frost front:Φ0ωi0 = Φ1ωi1 −AmivfΦ0 = Φ1 −AmivfAvf [ρsCp,s (T1 − T0)−mi∆Hi] = Φ0 (T1 − T0) (ωi0Cp,i + ωj0Cp,j)

Recovery step:

Temperature front: vr =Φ3,r(ωi3,rCp,i+ωj3,rCp,j)

AρsCp,s

Defrost front:Φ∗

1,rω∗

i1,r= Φ2,rωi2,r +Amivd,r

Φ∗

1,r = Φ2,r +Amivd,rAvd,r

[

ρsCp,s

(

T2 − T ∗

1

)

+m∗

i∆Hi

]

= Φ2,r

(

T2 − T ∗

1

)

(ωi2,rCp,i + ωj2,rCp,j)

Frost front:Φ1,rωi1,r = Φ∗

1,rω∗

i1,r−A

(

m∗

i −mi

)

vf,rΦ1,r = Φ∗

1,r −A(

m∗

i −mi

)

vf,rAvf,r

[

ρsCp,s

(

T ∗

1− T1

)

(

m∗

i −mi

)

∆Hi

]

= Φ1,r

(

T ∗

1− T1

)

(ωi1,rCp,i + ωj1,rCp,j)

Cooling step:

Temperature front: vc =Φ0,c(ωi0,cCp,i+ωj0,cCp,j)

AρsCp,s

* Note that for a temperature dependent heat capacity, the term Cp (Ty − Tx)

should be replaced by:

∫ Ty

Tx

CpdT

Page 46: Novel process concept for cryogenic CO2 capture

2.5 Process analysis 37

and 2.5. It can be observed that the front positions, equilibrium temper-ature, as well as the amount of mass deposited per unit of bed volumematch very well between the two approaches. Although heat and massdispersion is included in the advanced model, fronts are reasonably sharpduring the capture step. Especially the frost front is well defined, whichcan be attributed to a ‘self sharpening’ effect of the exothermic desub-limation. In order to demonstrate that the outcomes of the advancedmodel approach the simplified model even closer, an additional simula-tion has been carried out using the detailed model in which the actualheat and mass dispersion coefficients have been decreased by a factor100. The excellent agreement between the two models can be discernedfrom Fig. 2.4. Also for other inlet compositions and initial bed tempera-tures the two models agree very well (results are not included here).

The recovery step has also been simulated using the two models. Asalready explained before, additional CO2 will deposit onto the packing inthe initial phase of the recovery step. This is described by both modelsand again matches well, as observed in the temperature and mass de-position profiles after 10 seconds in Fig. 2.5a and 2.5b respectively. Itis observed again that when assuming low axial heat and mass disper-sion, the solution of the advanced model is approaching the sharp frontapproach. Finally the temperature and mass deposition profiles havebeen computed for the recovery step after 400 seconds, as illustrated inFig. 2.5c and 2.5d. It can be observed that dispersion is playing a moreprominent role during the recovery step, which is related to higher flowrates in comparison to the capture step. For that reason, the sharp frontapproach is especially suited to describe the capture step.

2.5 Process analysis

This section aims at giving an overview of the influences of several processparameters on the process performance, using the sharp front approach.The influences of the initial bed temperature, inlet composition, inlet tem-perature and packing material are analyzed on the basis of two aspects:the amount of CO2 deposited per unit of bed volume and the requiredspecific cooling duty, which is defined as:

Q =Vbed (1− εg) ρsCp,s (Ts − T0)

(ΦCO2,in − ΦCO

2,out)tstep

[

J

kgCO2

]

(2.18)

Page 47: Novel process concept for cryogenic CO2 capture

38 Cryogenic packed bed process concept

0 2 4 6 8

-150

-100

-50

0

50

100

150

Tem

pera

ture

[°C

]

Axial position [m]

SF approach Numerical model Numerical model - low dispersion

(a)

0 2 4 6 8

0

20

40

60

80

100 SF approach Numerical model Numerical model - low dispersion

Mas

s dep

ositi

on [k

g·m

-3]

Axial position [m](b)

Figure 2.4: Simulated axial temperature (a) and mass deposition (b) profiles forthe capture step. The line indicated with ‘low dispersion’ shows the profile cal-culated with the advanced model, in which the axial mass and heat dispersioncoefficients have been decreased by a factor 100.

The numerator represents the amount of energy required to cool downthe bed after a recovery step from Ts to a temperature T0, which is theinitial bed temperature before starting a capture step. The temperatureTs is assumed to be -70◦C for all cases. Initially the bed temperature atthe zone where CO2 is deposited will increase to -78◦C during the recov-ery step, as pure CO2 is being fed to the system and is being depositedadditionally onto the packing surface. However, the outlet flow during therecovery step is recycled and will increase slightly in temperature and isassumed to be at -70◦C, which will therefore be the initial bed tempera-ture before a cooling step is started. In practice, not the entire bed will beat this temperature at the end of the recovery step. The zone located closeto the outlet of the bed will have a slightly higher temperature. On theother hand, as explained earlier, it is actually not necessary to cool downthe entire bed to T0 during the cooling step and therefore the numeratorgives a realistic value for the required cooling duty. The denominator in(2.18) gives the amount of CO2 captured during a capture step. ΦCO

2,in

and ΦCO2,out are the inlet and outlet mass flow rates of CO2 during a

capture step.

Page 48: Novel process concept for cryogenic CO2 capture

2.5 Process analysis 39

0 2 4 6 8

-150

-100

-50

0

50

100

150

Tem

pera

ture

[°C

]

Axial position [m]

SF approach Numerical model Numerical model - low dispersion

(a)

0 2 4 6 8

0

20

40

60

80

100

120 SF approach Numerical model Numerical model - low dispersion

Mas

s dep

ositi

on [k

g·m

-3]

Axial position [m](b)

0 2 4 6 8

-150

-100

-50

0

50

100

150

Tem

pera

ture

[°C

]

Axial position [m]

SF approach Numerical model Numerical model - low dispersion

(c)

0 2 4 6 8

0

20

40

60

80

100

120 SF approach Numerical model Numerical model - low dispersion

Mas

s dep

ositi

on [k

g·m

-3]

Axial position [m](d)

Figure 2.5: Simulated axial temperature and mass deposition profiles for therecovery step, after 10s (a), (b) and 400s (c), (d). The lines indicated with ‘lowdispersion’ show the profiles calculated with the advanced model, in which theaxial mass and heat dispersion coefficients have been decreased by a factor 100.

Page 49: Novel process concept for cryogenic CO2 capture

40 Cryogenic packed bed process concept

The amount of CO2 deposited per unit of bed volume is an importantindicator of the required capital costs, while the calculated specific cool-ing duty gives a good indication of the energy requirements for differentcases. Blowers and compressors are responsible for part of the powerconsumption. The techno-economic evaluation described in Chapter 6includes these requirements and discussed the economic feasibility ofthe newly developed process concept.

2.5.1 Initial bed temperature

The required specific cooling duty and CO2 mass deposition as functionof the initial bed temperature are illustrated in Fig. 2.6. Below -130◦Cthe initial bed temperature hardly affects the specific cooling duty, theadditional energy required to cool the bed to lower temperatures will atthe same time result in a correspondingly higher CO2 storing capacity ofthe bed and therefore causing a constant specific cooling duty. However,at initial bed temperatures above -120◦C, the specific cooling duty willincrease strongly. This is related to the decreasing amount of CO2 beingcaptured, which will decrease exponentially above -120◦C. For example:when feeding a N2/CO2 mixture containing 10 vol.% CO2 to a bed cooledat -120◦C, 90% of the fed CO2 is recovered. However, when feeding thesame mixture to a bed cooled at -110◦C, only 12% CO2 is recovered.Therefore, a relatively low amount of extra cooling will result in muchhigher CO2 recovery rates. This effect is directly related to the exponentialtemperature dependency of the equilibrium CO2 vapor pressure.

2.5.2 CO2 inlet concentration

When the CO2 fraction in the flue gas feed decreases, the specific coolingduty will increase, as shown in Fig. 2.6a. This can be explained by thefact that less CO2 is stored in the bed at lower inlet CO2 concentrations(see Fig. 2.6b), while the energy required to cool down the bed remainsthe same.

2.5.3 Inlet temperature

A higher gas inlet temperature will cause the defrost front to move faster,because more heat is available for desublimation. This causes the CO2

Page 50: Novel process concept for cryogenic CO2 capture

2.5 Process analysis 41

-160 -150 -140 -130 -120 -1100

1

2

3

4

5

6

7y

CO2,in: 0.05 0.10 0.15 0.2

Coo

ling

duty

[MJ·k

g-1C

O2]

Initial bed temperature [°C]

(a)

-160 -150 -140 -130 -120 -1100

10

20

30

40

50

60

70

80y

CO2,in: 0.05 0.10 0.15 0.2

Mas

s dep

ositi

on [k

g·m

-3 bed]

Initial bed temperature [°C]

(b)

Figure 2.6: Specific cooling duty (a) and mass deposition (b) as function of theinitial bed temperature for different inlet CO2 fractions.

concentration after the defrost front to increase at the same time. Alsomore CO2 will deposit per unit of packing volume and the frost frontmoves faster, and consequently the step time reduces. Therefore, thespecific cooling duty will increase when increasing the inlet temperature.However, the specific cooling duty reaches an asymptotic value when in-creasing the temperature. This is related to the fact that the heat re-quired to heat up the packing (and cool down the gas) relative to theheat involved in the evaporation of previously deposited CO2 is playinga dominant role at higher inlet temperatures. When for example feedinga mixture containing 10% CO2 to a bed refrigerated to -140◦C at an in-let temperature of 150◦C, the amount of heat required for sublimation isonly 8.2% of the amount of energy involved in cooling down the gas. Itshould be noted that the contribution of desublimation at the frost frontdoes play a significant role in the energy balance.

2.5.4 Bed properties

The packing properties (i.e. material or porosity) will not influence thespecific cooling duty. When the heat capacity of a packing material is forexample doubled (and the bed dimensions remain the same), the amountof cooling will also be doubled. However, the amount of CO2 depositedper unit of bed volume will change proportionally, and therefore the spe-

Page 51: Novel process concept for cryogenic CO2 capture

42 Cryogenic packed bed process concept

cific cooling costs will not be influenced. However, bed dimensions areinfluenced when changing the packing material and keeping step timesconstant and therefore capital costs are influenced by the packing choice.On the other hand, the packing choice will influence the pressure dropand axial dispersion: Effects which are not accounted for in this analysis,but which will be included in the techno-economic analysis in Chapter 6.

2.6 Discussion and conclusions

A novel cryogenic CO2 capture concept, based on dynamically operatedpacked beds has been presented in this chapter. An advanced numericalmodel has been derived, which is able to describe the development ofaxial temperature, mass deposition and gas concentration profiles in thepacked beds. In addition to the advanced numerical model, a simplifiedmodel based on a sharp front approach has been presented. This sharpfront approach offers the possibility to carry out quick calculations andevaluate process designs.

The proposed process concept shows several advantages compared tocryogenic separation based on conventional heat exchangers. In the firstplace, plugging is intrinsically avoided. Due to the limited amount of coldstored in the packing material, also a limited amount of CO2 is desub-limated. In Fig. 2.5b it can for example be observed that the amount ofCO2 during the recovery step is approximately 80 kg/m3, which is cor-responding to a volume fraction of approximately 0.06, depending on thedensity of solid CO2. The gas void fraction of the used packing was muchhigher (0.7) and plugging or an increase in pressure drop is thereforeavoided.

Another advantage of the proposed concept is the very high purity ofthe treated gas. An initial bed temperature of -140◦C will result in aCO2 fraction in the outlet of less than 0.1% during the capture step. Theproposed concept can therefore recover more than 99% of the CO2 whenfeeding a mixture containing 10 vol.% CO2, while for example scrubbingtechnology is only able to capture 90% of CO2 at acceptable absorbersizes.

Another possible advantage of the proposed concept is that other polu-tants such as NOx, SOx and H2O can be captured simultaneously, avoid-ing expensive pre-treatment. It was demonstrated in this chapter thatH2O and CO2 can be captured simultaneously. However, it was also ex-

Page 52: Novel process concept for cryogenic CO2 capture

2.6 Discussion and conclusions 43

plained that the allowed water content in the flue gas is limited, due tothe heat required for evaporation of the condensed H2O in the recoverystep. Finally, an additional advantage of the concept is that simple lowpressure vessels packed with low cost inert packing material can be used,reducing capital costs.

Notation

A cross sectional area, [m2]as specific solid surface area per unit bed volume, [m2/m3]ci constant used in polynomial for heat capacityCp heat capacity, [J/kg/K]D diffusion coefficient, [m2/s]Deff effective diffusion coefficient, [m2/s]dh hydraulic diameter, [m]f friction factor, [-]g mass deposition rate constant, [s/m]L monolith length, [m]Mi molecular weight of component i, [kg/mol]mi mass deposition of component i per unit bed volume, [kg/m3]m′′

i mass deposition rate per unit surface area for component i,[kg/m2/s]

nc number of components, [-]P pressure, [Pa]Pr Prandtl number (Cp,gηg/λg)Q Specific cooling duty, [J/kgCO2

]Re Reynolds number (ρgvgdh/ηgεg)t time, [s]T temperature, [K], [◦C]Vbed bed volume, [m3]v front velocity, [m/s]v superficial velocity, [m/s]y mole fraction, [-]z axial coordinate, [m]

Greek letters

αg,s heat transfer coefficient solids - gas bulk, [W/m2/K]∆Hi enthalpy change related to the phase change of component i,

Page 53: Novel process concept for cryogenic CO2 capture

44 Cryogenic packed bed process concept

[J/kg]εg bed void fraction, [-]η viscosity, [kg/m/s]λ thermal conductivity, [W/m/K]λeff effective conductivity, [W/m/K]ρ density, [kg/m3]Φ mass flow, [kg/s]ω mass fraction, [-]

Subscripts

0 initialc captured defrostf frostg gas phasei component iin inletr recoverys solid phase

Superscripts

σ equilibriumsub sublimation

Page 54: Novel process concept for cryogenic CO2 capture

3Experimental demonstration of the concept

Abstract

The fundamental principles of the novel process concept for cryogenicCO2 capture have been elucidated in Chapter 2. This chapter presentsan experimental demonstration of the concept. For this purpose, anexperimental setup has been constructed in which the evolution of theprevailing axial temperature profiles can be measured accurately duringthe capture step. Furthermore, the outlet CO2 concentration has beenmonitored during measurements. Experiments have been carried out forN2/CO2 mixtures with and without the addition of H2O. The influences ofthe initial bed temperature and inlet gas composition on the evolution ofthe axial temperature profiles have been studied. Finally, the experimen-tal profiles have been compared with profiles calculated using the detailednumerical model outlined in Chapter 2. When accounting for radiationeffects, the model can very well describe the experimental observations.1

1This chapter is based on the papers: Tuinier et al., Chem. Eng. Sci. , 65(1), 114-119,2010 and Tuinier et al., Int. J. Greenhouse Gas Control, 5(4), 694-701, 2011.

Page 55: Novel process concept for cryogenic CO2 capture

46 Experimental demonstration of the concept

3.1 Introduction

A novel process concept to cryogenically capture CO2 from flue gases waspresented in Chapter 2. This chapter gives an experimental validationof the capture step. The used experimental setup and procedure will bedescribed first. After that, experimental results for measurements withN2/CO2 and N2/CO2/H2O mixtures are presented. Finally, the obtainedresults are compared to outcomes of the advanced numerical model.

3.2 Experimental setup and procedure

A flowsheet and a picture of the experimental setup are shown in Fig. 3.1and 3.2 respectively. A glass tube (OD × ID × L = 40 × 35 × 300 mm)surrounded by a glass vacuum jacket (OD × ID = 60 × 55 mm) wasfilled with monodisperse blue spherical glass beads (dp = 4.04 mm, ρs =2547 kg/m3). The bed was cooled with a N2 gas flow which was refriger-ated in a coil positioned in a liquid nitrogen bath. After cooling, the feedwas switched to a N2/CO2 mixture with or without the addition of H2O.These mixtures were prepared by controlling the gas flow rates of N2 andCO2 with mass flow controllers (Bronkhorst El-flow). This mixture couldthen be fed to a saturator filled with demineralized water. The water con-tent of the feed flow could be varied by changing the temperature in thesaturator. The flow from the saturator could be further overheated byusing an electrically traced line. The temperatures in the bed were mea-sured along the bed length in the radial center with 11 thermocouples(Thermo-Electric K-type) at every 3 cm in axial direction. The pressure atthe inlet of the bed was monitored using an analogue pressure indicator.The CO2 content in the outlet stream was analyzed with an IR-analyzer(Sick-Maihak, s610, 0-3 vol.%). The front of sublimated CO2 was visuallyinspected with a camera. Axial temperature profiles have been measuredfor different initial bed temperatures and inlet mole CO2 and H2O frac-tions (see Table 3.1).

Page 56: Novel process concept for cryogenic CO2 capture

3.2

Exp

erim

en

talsetu

pa

nd

pro

ced

ure

47

Figure 3.1: Flowsheet of the vacuum insulated packed bed.

Page 57: Novel process concept for cryogenic CO2 capture

48 Experimental demonstration of the concept

Table 3.1: Experimental conditions

Experiment #1 #2 #3 #4 #5 #6Initial bed temperature [◦C] −140 −140 −140 −140 −120 −140Inlet CO2 mole fraction [-] 0.2 0.3 0.2 0.1 0.2 0.2Inlet H2O mole fraction [-] 0 0 0.022 0.020 0.021 0.046Total mass flow [10−4 kg/s] 2.55 2.68 2.35 2.23 2.35 2.39

3.3 Results

The bed was cooled using refrigerated N2 until the bed reached a station-ary temperature profile. Due to heat radiation into the system this initialprofile is slightly increasing (almost linearly) from the inlet. The tempera-ture difference between the inlet and outlet is typically about 20◦C, whichis relatively small compared to the temperature difference between the re-frigerated bed and the gas being fed during the capture cycle. It shouldbe noted that during experiments no pressure excursions were observed,and the pressure drop was very small.

3.3.1 N2/CO2 mixtures

When feeding N2/CO2 mixtures to the refrigerated bed, a moving front ofdeposited CO2 was observed visually as depicted in Fig. 3.4. The axialtemperature profiles at several time steps for experiment #1 are shownin Fig. 3.3a. After approximately 200 seconds the frost front reached theend of the bed and CO2 breakthrough was detected in the outlet stream,as shown in Fig. 3.5. The mixture is fed through the same inlet tubeas the refrigerated N2 during the cooling step. Therefore the tempera-ture of the packing at the inlet does not attain ambient temperaturesimmediately, but increased slowly as cam be seen from Fig. 3.3a. TheCO2 content in experiment #2 was increased to 30 vol.% CO2. The in-creased CO2 inlet concentration resulted in a higher saturation tempera-ture (-91.5◦C versus -94.5◦C) and therefore the packing storage capacityslightly increased. However, due to the higher molar CO2 feed flow rate,the front velocity of the frost front increased. Fig. 3.3b shows that CO2

breakthrough already occurred after about 150 seconds.

Page 58: Novel process concept for cryogenic CO2 capture

3.3 Results 49

Figure 3.2: Picture of the experimental setup.

Page 59: Novel process concept for cryogenic CO2 capture

50 Experimental demonstration of the concept

0.00 0.05 0.10 0.15 0.20 0.25 0.30-140-120-100-80-60-40-200

2040

50 s 100 s 150 s 200 s

Tem

pera

ture

[°C

]

Axial position [m](a)

0.00 0.05 0.10 0.15 0.20 0.25 0.30-140-120-100-80-60-40-200

2040

50 s 100 s 150 s

Tem

pera

ture

[°C

]

Axial position [m](b)

Figure 3.3: Experimental (markers) and simulated (lines) axial temperature pro-files for experiments #1 (a) and #2 (b).

Figure 3.4: CO2 ice formed at the packing surface during a capture step.

Page 60: Novel process concept for cryogenic CO2 capture

3.3 Results 51

0 50 100 150 200 250 300

0

5

10

15

20

25

30

Vol

.% C

O2

Time [s]

Experiment Simulation

Figure 3.5: Experimental (markers) and simulated (line) CO2 concentration inthe outlet during experiment #1.

3.3.2 N2/CO2/H2O mixtures

Fig. 3.7a shows axial temperature profiles measured during experi-ment #3. Again deposition of CO2 occurs and a CO2 front develops, simi-lar to experiment #1. The gas feed mixture also contains H2O during thisexperiment, resulting in H2O condensation at the packing surface, whichis shown in Fig. 3.6. In between the zones in which H2O is condensed andCO2 is desublimated, a small amount of H2O ice is also observed. Thisis caused by the decrease in the H2O concentration and the temperatureat the condensing front. At some point the mixtures reaches conditionsbelow the triple point of H2O and desublimation of H2O is observed. Theamount of desublimated H2O remains constant during a capture stepand no influence on the temperature profiles is observed. The effect ofH2O condensation on the axial temperatures profiles is small in the first200 seconds of the measurement. However, after CO2 breakthrough, thegas flow was not stopped in order to show further development of theH2O front. Fig. 3.7b shows that a water front is also moving through thebed and that an equilibrium is formed at a temperature of approximately28◦C. Lowering the CO2 inlet concentration from 20% to 10% in exper-iment #4 results in a lower equilibrium temperature for CO2 (-98◦C), asillustrated in Fig. 3.7c. The front of freezing CO2 will move slower in thiscase, breakthrough is observed after approximately 350 seconds. Thedevelopment of the H2O front is not influenced by the changed CO2 in-

Page 61: Novel process concept for cryogenic CO2 capture

52 Experimental demonstration of the concept

let concetration, as shown in Fig. 3.7d. Fig. 3.8a and 3.8b show axialtemperature profiles for experiment #5 with a higher initial bed tempera-ture. This higher bed temperature results in shorter cycle times, becauseless CO2 is stored per unit of bed volume. Moreover, the purity of thecleaned flue gas will be somewhat lower (98.8 % N2 compared to 99.9 %in the other experiments). Finally, the effect of an increased H2O inletconcentration can be observed in Fig. 3.8c and 3.8d. The evolution ofthe CO2 fronts is not much influenced, but a clear effect is visible forthe H2O front. The equilibrium temperature reaches a higher value ofapproximately 40◦C, compared to 28◦C in the other experiments.

3.4 Simulations

The experimental results have been compared with profiles calculated bythe advanced numerical model, which was described in Chapter 2. Thepacking material used for the experiments consisted of spherical parti-cles. Therefore the correlations listed in Table 2.2, which are valid fora structured monolith packing are not applicable here, and are replacedby the correlations summarized in Table 3.2. As no information on thesublimation rates is available in literature, the equilibrium time constant(g) was determined by comparing simulation results with the experimen-tal findings. Results are best described when using a constant of about1·10-6 s/m. Chapter 4 focuses on desublimation rates in more detail.

The initial temperature profile used in the simulations is taken fromthe experiments. As mentioned before, the initial bed temperature isnot totally uniform, but is increasing slightly from the inlet towards theoutlet. This is caused by the heat leak into the system. To accountfor this heat leak in more detail, the tube was first cooled down, thencooling was stopped and the temperature rise was measured as functionof the time in the radial center and close to the tube wall. It was foundthat the temperature difference between these two locations was minimaland that the temperature rise could be well described by an additionalradiative energy influx. Therefore, the following contribution was addedto the energy balance:

φrad =4

dtube

σ(

T 4h − T 4

c

)

1

ǫc+ Ac

Ah

(

1

ǫh− 1

)

[

J

m3s

]

(3.1)

Page 62: Novel process concept for cryogenic CO2 capture

3.4 Simulations 53

Figure 3.6: Desublimated and condensed components at the packing surfaceobserved during a capture step (feed enters from the bottom). Three zones areobserved (from bottom to top): liquid H2O, solid H2O and solid CO2.

Page 63: Novel process concept for cryogenic CO2 capture

54 Experimental demonstration of the concept

0.00 0.05 0.10 0.15 0.20 0.25 0.30-140-120-100-80-60-40-200

2040

100 s 200 s

Tem

pera

ture

[°C

]

Axial position [m](a)

0.00 0.05 0.10 0.15 0.20 0.25 0.30-100-80-60-40-200

20406080

100

400 s 1000 s

Tem

pera

ture

[°C

]

Axial position [m](b)

0.00 0.05 0.10 0.15 0.20 0.25 0.30-140-120-100-80-60-40-200

2040

100 s 200 s

Tem

pera

ture

[°C

]

Axial position [m](c)

0.00 0.05 0.10 0.15 0.20 0.25 0.30-100-80-60-40-200

20406080

100

400 s 1000 s

Tem

pera

ture

[°C

]

Axial position [m](d)

Figure 3.7: Experimental (markers) and simulated (lines) axial temperature pro-files for experiments #3 (a) (b) and #4 (c)(d).

Page 64: Novel process concept for cryogenic CO2 capture

3.4 Simulations 55

0.00 0.05 0.10 0.15 0.20 0.25 0.30-140-120-100-80-60-40-200

2040

100 s 200 s

Tem

pera

ture

[°C

]

Axial position [m](a)

0.00 0.05 0.10 0.15 0.20 0.25 0.30-100-80-60-40-200

20406080

100

400 s 1000 s

Tem

pera

ture

[°C

]

Axial position [m](b)

0.00 0.05 0.10 0.15 0.20 0.25 0.30-140-120-100-80-60-40-200

2040

100 s 200 s

Tem

pera

ture

[°C

]

Axial position [m](c)

0.00 0.05 0.10 0.15 0.20 0.25 0.30-100-80-60-40-200

20406080

100

400 s 1000 s

Tem

pera

ture

[°C

]

Axial position [m](d)

Figure 3.8: Experimental (markers) and simulated (lines) axial temperature pro-files for experiments #5 (a) (b) and #6 (c) (d).

Page 65: Novel process concept for cryogenic CO2 capture

56 Experimental demonstration of the concept

In which σ is the Stefan-Boltzmann constant, ǫ the integral emissivityand the subscripts h and c stand for the hot and cold side. In the usedexperimental setup, both the hot side (jacket) and cold side (tube) arecomposed of glass, and therefore an equal integral emissivity of 0.9 isassumed.

Furthermore, it was found that the inlet temperature of the gas wasonly rising slowly in time after switching to the capture step, due to coldstored in the inlet of the tube, insulation material, piping etc. Therefore,the inlet temperature (and corresponding composition) used in simula-tions was based on the measured inlet temperature. Finally, it should benoted that the heat capacity of the glass wall was included in simulations(in the accumulation term).

The resulting calculated temperature profiles have been plotted in thesame figures as in which the experimental profiles were shown. It can beconcluded that the developed model is very well capable to describe allexperimental findings.

3.5 Discussion and conclusions

The novel process concept for cryogenic CO2 capture has been demon-strated experimentally in this chapter. The developed numerical model isvery well capable to describe the experiments for N2/CO2 mixtures, withand without the addition of H2O. Both the axial temperature profiles,as the outlet concentration can be very well predicted by the numeri-cal model, when assuming a desublimation equilibration time constantg equal to 1·10-6 s/m. A detailed study on the desublimation rates ispresented in Chapter 4.

The influences of initial bed temperature and inlet gas compositionwere shown. The effect of these parameters on the process economicswill be discussed in Chapter 6. This chapter only focused on the capturestep. An experimental setup in which all three process steps (capture,recovery and cooling) can be operated in parallel has been constructed aswell. A description of this setup and the results are discussed in Chapter5.

Page 66: Novel process concept for cryogenic CO2 capture

3.5 Discussion and conclusions 57

Table 3.2: Heat and mass transfer coefficients.

Effective axial heat dispersion in a transient packed bed(Vortmeyer and Berninger, 1982):

λeff = λbed,0 +RePrλg

Peax+

Re2 Pr2 λg

6(1− εg)Nu

In which Peax is calculated according to Gunn and Misbah (1993):

Peax = 2p1−p p = 0.17 + 0.33 exp

−24

Re

λbed,0 is calculated according to Zehner and Schlunder (1970).

Gas-to-particle heat transfer coefficient (Gunn, 1978):

Nu =(

7− 10εg + 5ε2g)

(

1 + 0.7Re0.2 Pr1/3)

+(

1.33− 2.4εg + 1.2ε2g)

Re0.7 Pr1/3

Axial mass dispersion (Edwards and Richardson, 1968):

Deff

vgdp=

0.73

Re Sc+

0.5

εg

(

1 +9.7εgRe Sc

)

Page 67: Novel process concept for cryogenic CO2 capture

58 Experimental demonstration of the concept

Notation

A surface of tube wall, [m2]Cp heat capacity, [J/kg/K]Deff effective diffusion coefficient, [m2/s]dp particle diameter, [m]dtube tube diameter, [m]g mass deposition rate constant, [s/m]ID inner diameter, [mm]L bed length, [mm]Nu Nusselt number (αgsdp/λg)OD outer diameter, [mm]p parameter in axial heat dispersion coefficientPeax Peclet number for axial heat dispersion (ρgvgdpCp,g/λax)Pr Prandtl number (Cp,gηg/λg)Re Reynolds number (ρgvgdp/ηg)Sc Schmidt number (ηg/ρg/D)T temperature, [K]v superficial velocity, [m/s]

Greek letters

εg bed void fraction, [-]ǫ integral emissivity, [-]η viscosity, [kg/m/s]λ thermal conductivity, [W/m/K]λbed,0 effective bed conductivity at no flow conditions, [W/m/K]λeff effective conductivity, [W/m/K]ρ density, [kg/m3]σ Stefan-Boltzmann constant, [J/s/m2/K4]φrad radiation heat influx, [J/m3/s]

Subscripts

c cold sideg gas phaseh hot sides solid phase

Page 68: Novel process concept for cryogenic CO2 capture

4Mass deposition rates of carbon dioxide onto a

cryogenically cooled surface

Abstract

The rates of CO2 mass deposition onto a cryogenically cooled surface areimportant for CO2 removal processes based on cryogenics. A slow massdeposition rate would lead to early breakthrough of CO2 in the conceptas proposed in this thesis. Fast mass deposition is therefore required.However, no detailed experimental data for kinetic rate expressions areavailable in literature. Therefore, a dedicated experimental setup to mea-sure CO2 mass deposition rates under well defined conditions was de-signed. Experiments have been carried out for both pure CO2 as well asCO2/N2 mixtures. The experiments showed that heat transfer throughthe frost layer will slow down the mass deposition process. Furthermore,it was found that the addition of N2 to the gas phase has a large influenceon the mass deposition rates, due to the introduction of a mass trans-fer resistance towards the frost surface. To describe the experimentallyobserved behavior, a detailed frost growth model was developed, basedon mass and energy balances. Expressions for the frost density as func-tion of the frost temperature, and for the effective frost conductivity as

Page 69: Novel process concept for cryogenic CO2 capture

60 Mass deposition rates of carbon dioxide

function of the frost density were derived and implemented in the model.When accounting for drift fluxes, the model is well able to describe theobserved behavior. For packed bed conditions, the formed frost layer isthin and mass deposition rates are mainly determined by external masstransfer. However, for process concepts in which heat exchanging sur-faces are continuously cooled, the frost layer will likley influence frostgrowth rates and the developed model is a useful tool for design studies.

4.1 Introduction

In this thesis a novel process concept has been developed for the re-moval of CO2 from flue gases by feeding the mixture to a cryogenicallyrefrigerated packed bed. CO2 will freeze onto the packing surface, andan effective separation between CO2 and N2 is obtained. Information onmass deposition rates is important for the design of this process. Whenmass deposition would be relatively slow, the front of freezing CO2 woulddisperse and early CO2 breakthrough would occur, resulting in an ineffi-cient use of the cold stored in the bed. For other CO2 removal processesin which the heat exchanging surface is being cooled continuously, massdeposition rates are also important. It was found (Chang et al., 2009;Chang and Smith, 1990) that solid CO2 will build up locally in a heatexchanger, eventually leading to a large pressure drop or even pluggingwhen not switched to a regeneration step in time. Knowledge on desub-limation rates and the effect of layer thickness on the desublimation isrequired to predict the location where CO2 will accumulate and the timewhen regeneration is necessary.

To the best of our knowledge the number of studies on the fundamen-tals of CO2 desublimation is rather limited. Shchelkunov et al. (1986)studied CO2 desublimation on a plate which is positioned in a flow par-allel to the plate, resulting in mass transfer perpendicular to the flow di-rection. Frost was measured for only one CO2 partial pressure (4.35 kPa)and was not uniform along the plate. Ogunbameru et al. (1973) studiedCO2 frost formation on a flat plate, which was refrigerated with liquid N2,but again for only one concentration (2 vol.%).

This chapter describes an extensive study on CO2 desublimation ratesto obtain more insight in the role of kinetics and mass and heat trans-fer for different CO2 (partial) pressures, surface and gas phase tempera-tures and flow conditions. The chapter is organized as follows: first the

Page 70: Novel process concept for cryogenic CO2 capture

4.2 Experimental 61

experimental setup and procedure are elucidated, followed by a descrip-tion of the results of experiments for pure CO2 and mixtures of N2 andCO2. Then, a frost growth model is derived to describe the experimentalfindings. Finally the conclusions and its significance for cryogenic CO2

removal equipment is given.

4.2 Experimental

This section gives a description of the dedicated experimental setup andthe experimental procedure. Also the interpretation and processing of theobtained experimental data is discussed.

4.2.1 Setup

The design of the experimental setup was based on a setup used formeasurements of phase equilibria of CO2 containing gas mixtures asdescribed by Le and Trebble (2007). A schematic representation of theexperimental setup is given in Fig. 4.1. The setup consists of a stirredcell with glass walls (inner diameter = 85 mm, height = 110 mm), con-taining a round cooled surface (d = 20 mm) at the bottom, at which CO2

is desublimating during a measurement. The axial propeller type stirrerhas a diameter of 50 mm, and has a maximum rotation speed of 1730RPM. The stirred cell is positioned in a large custom made dewar vesselwith viewing stripes (KGW-isotherm). This dewar is filled with specialcooling liquid (3M Novec 7200) and contains a large coil which can befed with liquid N2 for cooling. The flow of liquid N2 is controlled by amagnetic valve in order to control the temperature of the (stirred) liquidand hence of the gas phase in the stirred cell. The dewar contains two(K-type) thermocouples to monitor the temperature of the cooling liquid.The stirred cell is furthermore connected to a vacuum pump, a premixvessel and a CO2 feed line. The cell is equipped with an accurate vacuumgauge (Inficon CDG025D) and contains two K-type thermocouples in thegas phase at two different axial positions. The temperature of the cooledsurface is controlled by a refrigerated gaseous N2 flow. This N2 flow ispassed through a coil submerged in a dewar filled with liquid N2 andthe flow rate is controlled using a mass flow controller (Brooks Smart,60 NL/min) connected to a PID control loop. The refrigerated N2 is thencontacted intensively with the cooled copper made plate in the stirred

Page 71: Novel process concept for cryogenic CO2 capture

62 Mass deposition rates of carbon dioxide

cell. Inside the cooled plate four 0.5 mm calibrated T-type thermocou-ple have been installed at different locations, to monitor the temperaturein the cooled surface. The cooled surface is surrounded by PVC, whichminimizes heat transfer to other parts of the bottom of the cell due to itslow conductivity. Monitoring temperatures and pressures and control-ling valves and mass flow controllers is performed using NI LabVIEW. Apicture of the setup is shown in Fig. 4.2.

4.2.2 Procedure

An experiment is started by evacuating the stirred cell and cooling downthe cooling liquid in the dewar vessel surrounding the stirred cell to therequired set point temperature. Then the cold surface is brought to thedesired temperature. The premix vessel (equipped with a GE Sensing PTX1400 pressure gauge) is filled with pure CO2 or a mixture of N2 and CO2.Valve V-3 is opened to allow the gas to enter the cell until the user givenpressure set point is reached. At this point CO2 immediately starts todesublimate at the surface and the pressure in the cell decreases. Bycontrolling MFC-1 (using a PID control loop), fresh (pure) CO2 is fed tothe cell to maintain the set point pressure. The flow through the massflow controller is accurately monitored and is directly proportional to theamount of CO2 desublimating at the cold surface. Frost growth duringthe measurement is recorded using a Canon EOS digital camera. After ameasurement, cooling of the surface is stopped and the system is flushedwith N2.

Page 72: Novel process concept for cryogenic CO2 capture

4.2

Exp

erim

en

tal

63

Figure 4.1: Simplified process scheme of the experimental setup.

Page 73: Novel process concept for cryogenic CO2 capture

64 Mass deposition rates of carbon dioxide

Figure 4.2: Picture of the experimental setup (with the dewar vessel lowered).

To make sure that no temperature and concentration gradients arepresent in the gas phase, the mixing behavior of the stirred cell was ex-amined before starting experiments. This was done by feeding a N2/CO2

mixture to the cell and measuring the CO2 content in the outlet withan IR spectrometer. At some point, the CO2 feed was suddenly stoppedand the decrease in CO2 concentration in the outlet was measured intime. This was repeated for several rotation speeds, and the outcomeswere compared to the theoretical profile of an ideally stirred tank reactor.It was found that mixing approached ideal behavior at rotation speedsabove 130 RPM.

Experiments have been carried out under a wide range of conditions.To exclude any effects of dilution with N2, first measurements are per-formed with pure CO2 in the gas phase. The influence of the pressureof CO2 and the cold plate temperature have been studied. Subsequently,the effect of the amount of N2 present in the gas phase was studied, againfor different cold plate temperatures, but also for different gas phase tem-

Page 74: Novel process concept for cryogenic CO2 capture

4.2 Experimental 65

Table 4.1: Conditions for all experiments

Experiment PCO2

[mbar] PN2

[mbar] T0 [◦C] Tg [◦C] Ns [%]

P100 100 0 -130 -30 100P150 150 0 -130 -30 100P200 200 0 -130 -30 100P250 250 0 -130 -30 100T0-145 100 0 -145 -30 100T0-160 100 0 -160 -30 100

50T0-130 100 100 -130 -30 10050T0-140 100 100 -140 -30 10050T0-150 100 100 -150 -30 100

10T0-130 100 900 -130 -30 10010T0-140 100 900 -140 -30 10010T0-150 100 900 -150 -30 10010Tg-45 100 900 -130 -45 10010Tg-60 100 900 -130 -60 10010N75 100 900 -130 -30 7510N50 100 900 -130 -30 50

peratures and stirrer rotation speeds. The conditions for all experimentsare summarized in Table 4.1.

4.2.3 Data processing

This section shows how the experimental data have been processed. Thisis demonstrated in detail for one single experiment: experiment 50T0-

140. The temperature of the cold plate, of the gas phase in the stirredcell and the temperature of the cooling liquid in the dewar vessel duringthis experiment are shown in Fig. 4.3a. The temperature of the cold platecould be very well maintained at approximately -140◦C. The gas phasetemperature is decreasing slowly during the measurement by approxi-mately two degrees. This can be attributed to the heat exchange with thecold surface at the bottom of the cell, which has a much lower temper-ature. The pressure in the cell, shown in Fig. 4.3b, could be controlledwell by compensating the depositing CO2 by feeding fresh CO2 into the

Page 75: Novel process concept for cryogenic CO2 capture

66 Mass deposition rates of carbon dioxide

cell. The output of the mass flow controller is initially high, and is de-creasing in time, as illustrated in Fig. 4.3c. Based on the output of themass flow controller, the mass accumulated at the cold plate can be cal-culated, which is shown in Fig. 4.3d. The increase in volume of the frostlayer was observed by the camera, the pictures of the growing ice layerare depicted in Fig. 4.4. These pictures clearly show that the ice layer isnot only growing in the vertical direction, but also in the horizontal di-rection and that the ice layer attains a curved surface at the edges. Thisbehavior complicates the derivation of the mass deposition rate per unitof surface area, as the surface area is changing in time. Furthermore,heat transfer within the layer will also take place in the radial directionat the edges of the ice layer. In order to facilitate the interpretation of theexperimental results, and later on to be able to develop a model for icelayer growth, it is desired to approach the growth as a one-dimensionalprocess. The pictures show that within a certain radius, the ice layermaintains an almost horizontal surface, even after 900 seconds of mea-surement time. It can be assumed that within this inner part of the icelayer, radial effects can be ignored. Therefore only the growth of the layerwithin the inner 1 cm radius is considered. However, the mass deposi-tion rate onto this inner part of the ice layer is not directly measured. Theaccumulated mass as plotted in Fig. 4.3d is formed at the entire surface,including the curved outer parts. The mass of CO2 depositing onto theinner 1 cm radius is being calculated as follows. First both the total vol-ume of the ice layer as well as the volume within the inner radius of 1cm have been calculated from the images. To calculate the volume fromthe (two-dimensional) image, the surface area of a pixel and the distanceto the central axis are determined using the image magnification factor.Then, this area is rotated around the central axis of the ice layer to ob-tain the volume of revolution of the pixel. This is done for the pixels atthe left side of the central axis as well as the right side and the result isaveraged, as the image is not perfectly symmetrical. This is carried outboth for the inner radius, as well as the entire frost layer. Thus, the ratioof the inner volume and the total volume is determined for every image.Fig. 4.5 shows the result of this procedure for experiment 50T0-140. Itcan be observed that especially in the beginning of the experiment thevolume ratio is decreasing, indicating that especially at the start of theexperiment the ice layer is growing significantly in the radial direction.Note that when there would be no ice layer formation in the radial direc-tion, the ratio would be 0.25 (due to the total radius of 2 cm compared

Page 76: Novel process concept for cryogenic CO2 capture

4.3 Results 67

to the inner 1 cm). When extrapolating to t = 0, it can indeed be ob-served that this ratio will approach approximately 0.25. To compute theamount of CO2 deposited in the inner part, the total mass accumulated(as plotted in Fig. 4.3d) is being multiplied with the obtained volume ra-tio. This is allowed, when assuming that the frost density at the innerpart of the ice layer is equal to the density of the entire layer. Fig. 4.5shows this amount of CO2 accumulated at the inner part of the ice layer.Finally, the mass deposition rate is calculated by fitting the accumulatedmass with a power law (also shown in Fig. 4.5), and differentiating theobtained equation. After data processing the following information hasbeen obtained for an experiment: the layer thickness, bulk density andmass deposition rate as function of the time. These results are presentedfor all experiments in the next section.

4.3 Results

The results for the experiments listed in Table 4.1 are presented anddiscussed in this section. The effect of the pressure of CO2 (withoutthe addition of N2) is shown in Fig. 4.6. It can be observed that for allmeasurements the initial mass deposition rate is very high, but is quicklydecreasing in time. The only property changing during an experiment isthe thickness of the ice layer (shown in Fig. 4.6b). It can therefore beconcluded that heat conduction through the ice layer towards the coldplate is playing an important role. Remarkable is the fact that there is nolarge effect of the CO2 pressure on the deposition process. Apparently,surface kinetics do not play a role at these conditions.

To further investigate the effect of heat conduction through the icelayer, experiments have been carried out with different plate tempera-tures. At lower temperatures, the heat flux through the ice layer be-comes higher, and therefore a higher mass deposition rate is measured,as shown in Fig. 4.7.

To study the effect of N2 on the deposition rates, measurements havebeen carried out by adding N2 to the gas, while keeping the CO2 partialpressure constant at 100 mbar. Fig. 4.8 shows results of measurementswith pure CO2, 50 vol.% (total pressure of 200 mbar) and 10 vol.% (totalpressure of 1000 mbar) mixtures. It can be very clearly observed thatmass deposition rates and the layer growth are slowed down by the pres-ence of N2 in the gas phase, which clearly indicates that mass transfer

Page 77: Novel process concept for cryogenic CO2 capture

68 Mass deposition rates of carbon dioxide

0 200 400 600 800

-48

-44

-40

-36

-32

-28

-24

-20 Cooling liquid dewar Gas phase cell Cold plate

Time [s]

Tem

pera

ture

[°C

]

-150

-140

-130

-120

-110

Temperature [°C

]

(a)

0 200 400 600 8000

50

100

150

200

250

Pres

sure

cel

l [m

bar]

Time [s]

(b)

0 200 400 600 800

0.00

0.05

0.10

0.15

0.20

0.25

CO

2 flow

into

cel

l [N

L·m

in-1

]

Time [s]

(c)

0 200 400 600 800

0.0

5.0x10-4

1.0x10-3

1.5x10-3

2.0x10-3

2.5x10-3

3.0x10-3

Mas

s CO

2 acc

umul

ated

[kg]

Time [s]

(d)

Figure 4.3: The measured temperatures (a), pressure (b), flow into the cell (c) andthe accumulated mass on the cold plate (d) during experiment 50T0-140.

Page 78: Novel process concept for cryogenic CO2 capture

4.3 Results 69

Figure 4.4: Photographs taken every minute of the CO2 layer for experiment50T0-140.

Page 79: Novel process concept for cryogenic CO2 capture

70 Mass deposition rates of carbon dioxide

0 200 400 600 8000.0

0.1

0.2

0.3

0.4

0.5

Volume ratio Mass accumulated inner part Power law fit

Time [s]

Vol

ume

ratio

[-]

0

2

4

6

8 Mass accum

ulated [kg·m-2]

Figure 4.5: The ratio of the ice layer volume formed at the inner 1 cm radius andthe entire layer volume plotted as a function of time for experiment 50T0-140.The accumulated mass per area for the inner 1 cm radius is plotted at the righty-axis.

Page 80: Novel process concept for cryogenic CO2 capture

4.3 Results 71

0 200 400 600 800

0.00

0.01

0.02

0.03

0.04

0.05 P

CO2 = 100 mbar

PCO2

= 150 mbar P

CO2 = 200 mbar

PCO2

= 250 mbar

Mas

s dep

ositi

on ra

te [k

g·m

-2·s-1

]

Time [s]

(a)

0 200 400 600 8000.000

0.002

0.004

0.006

0.008

0.010

0.012 P

CO2 = 100 mbar

PCO2

= 150 mbar P

CO2 = 200 mbar

PCO2

= 250 mbar

Laye

r thi

ckne

ss [m

]

Time [s]

(b)

0 200 400 600 800

0

400

800

1200

1600

PCO2

= 100 mbar P

CO2 = 150 mbar

PCO2

= 200 mbar P

CO2 = 250 mbar

Laye

r den

sity

[kg·

m-3

]

Time [s]

(c)

Figure 4.6: The experimental (markers) and simulated (lines) mass depositionrate (a), frost layer thickness (b) and layer density (c) as function of the time forthe pure CO2 experiments at different pressures (P100, P150, P200 and P250,see Table 4.1).

Page 81: Novel process concept for cryogenic CO2 capture

72 Mass deposition rates of carbon dioxide

0 200 400 600 800

0.00

0.01

0.02

0.03

0.04

0.05 T

0 = -130°C

T0 = -145°C

T0 = -160°C

Mas

s dep

ositi

on ra

te [k

g·m

-2·s-1

]

Time [s]

(a)

0 200 400 600 8000.000

0.002

0.004

0.006

0.008

0.010

0.012 T

0 = -130°C

T0 = -145°C

T0 = -160°C

Laye

r thi

ckne

ss [m

]

Time [s]

(b)

0 200 400 600 800

0

400

800

1200

1600

T0 = -130°C

T0 = -145°C

T0 = -160°C

Laye

r den

sity

[kg·

m-3

]

Time [s]

(c)

Figure 4.7: The experimental (markers) and simulated (lines) mass depositionrate (a), frost layer thickness (b) and layer density (c) as function of the time forthe pure CO2 experiments at different cold plate temperatures (P100, T0-145 andT0-160, see Table 4.1).

Page 82: Novel process concept for cryogenic CO2 capture

4.4 Development of a frost growth model 73

of CO2 from the gas phase to the forst layer plays a significant role. Itcan also be discerned that the density of the frost layer is influenced bythe presence of N2 in the gas phase. For the pure CO2 measurements,the frost density is quickly increasing to a plateau value, while for themixtures the density is slowly increasing in time. These effects will befurther discussed in section 4.4.

Concluding, mass transfer has a large influence on the deposition pro-cess. On the other hand, the mass deposition rate is still decreasing intime for the measurements with the CO2/N2 mixtures, which implies thatheat transfer through the solid layer still plays a role. To study this inmore detail, again the plate temperature has been varied, both for the50 vol.% (Fig. 4.9) as well as the 10 vol.% CO2/N2 (Fig. 4.10) mixtures.Again, it can be observed that there is a general trend of increasing massdeposition rates, when the cold plate temperature is decreasing. However,it should be noted that the differences are small.

The influence of the gas phase temperature on the mass depositionrates is shown in Fig. 4.11. A lower gas phase temperature results inslightly faster deposition rates, although the differences are close to therange of the experimental accuracy. Finally, the rotation speed of thestirrer has been changed. The mass transfer towards the frost surface isdecreasing at lower rotation speeds, resulting in lower deposition rates,as shown in Fig. 4.12.

4.4 Development of a frost growth model

In order to obtain a better understanding of the desublimation process ofCO2 onto a cold surface, a model will be developed in this section. Theoutcomes will be compared with the experimental results and finally thesignificance of the findings for the cryogenic packed bed concept for CO2

capture are discussed.

4.4.1 Model development

The experimental results revealed that both heat transfer through theformed solid ice layer and mass transfer of CO2 towards the ice surfaceplay an important role. Therefore, the frost growth is described as a mov-ing boundary problem, with both mass as well as heat transfer included.

Page 83: Novel process concept for cryogenic CO2 capture

74 Mass deposition rates of carbon dioxide

0 200 400 600 800

0.00

0.01

0.02

0.03

0.04

0.05 Pure CO

2

50 vol.% CO2

10 vol.% CO2

Mas

s dep

ositi

on ra

te [k

g·m

-2·s-1

]

Time [s]

(a)

0 200 400 600 8000.000

0.002

0.004

0.006

0.008

0.010 Pure CO

2

50 vol.% CO2

10 vol.% CO2

Laye

r thi

ckne

ss [m

]

Time [s]

(b)

0 200 400 600 800

0

400

800

1200

1600

Pure CO2

50 vol.% CO2

10 vol.% CO2

Laye

r den

sity

[kg·

m-3

]

Time [s]

(c)

Figure 4.8: The experimental (markers) and simulated (lines) mass depositionrate (a), frost layer thickness (b) and layer density (c) as function of the time forexperiments with different CO2 concentrations (T0-145, 50T0-140 and 10T0-140,see Table 4.1.)

Page 84: Novel process concept for cryogenic CO2 capture

4.4 Development of a frost growth model 75

0 200 400 600 800

0.000

0.005

0.010

0.015

0.020 T0 = -130°C

T0 = -140°C

T0 = -150°C

Mas

s dep

ositi

on ra

te [k

g·m

-2·s-1

]

Time [s]

(a)

0 200 400 600 8000.000

0.002

0.004

0.006

0.008

0.010 T

0 = -130°C

T0 = -140°C

T0 = -150°C

Laye

r thi

ckne

ss [m

]

Time [s]

(b)

0 200 400 600 800

0

400

800

1200

1600 T0 = -130°C

T0 = -140°C

T0 = -150°C

Laye

r den

sity

[kg·

m-3

]

Time [s]

(c)

Figure 4.9: The experimental (markers) and simulated (lines) mass depositionrate (a), frost layer thickness (b) and layer density (c) as function of the time forthe 50 vol.% CO2 experiments at different cold plate temperatures (50T0-130,50T0-140 and 50T0-150, see Table 4.1).

Page 85: Novel process concept for cryogenic CO2 capture

76 Mass deposition rates of carbon dioxide

0 200 400 600 8000.000

0.002

0.004

0.006

0.008

0.010 T

0 = -130°C

T0 = -140°C

T0 = -150°C

Mas

s dep

ositi

on ra

te [k

g·m

-2·s-1

]

Time [s]

(a)

0 200 400 600 800

0.000

0.001

0.002

0.003

0.004

0.005

0.006 T

0 = -130°C

T0 = -140°C

T0 = -150°C

Laye

r thi

ckne

ss [m

]

Time [s]

(b)

0 200 400 600 800

0

400

800

1200

1600 T

0 = -130°C

T0 = -140°C

T0 = -150°C

Laye

r den

sity

[kg·

m-3

]

Time [s]

(c)

Figure 4.10: The experimental (markers) and simulated (lines) mass depositionrate (a), frost layer thickness (b) and layer density (c) as function of the time forthe 10 vol.% CO2 experiments at different cold plate temperatures (10T0-130,10T0-140 and 10T0-150, see Table 4.1).

Page 86: Novel process concept for cryogenic CO2 capture

4.4 Development of a frost growth model 77

0 200 400 600 8000.000

0.002

0.004

0.006

0.008

0.010

Tg = -30°C

Tg = -45°C

Tg = -60°C

Mas

s dep

ositi

on ra

te [k

g·m

-2·s-1

]

Time [s]

(a)

0 200 400 600 800

0.000

0.001

0.002

0.003

0.004

0.005

0.006 T

g = -30°C

Tg = -45°C

Tg = -60°C

Laye

r thi

ckne

ss [m

]

Time [s]

(b)

0 200 400 600 800

0

400

800

1200

1600 Tg = -30°C

Tg = -45°C

Tg = -60°C

Laye

r den

sity

[kg·

m-3

]

Time [s]

(c)

Figure 4.11: The experimental (markers) and simulated (lines) mass depositionrate (a), frost layer thickness (b) and layer density (c) as function of the time forthe 10 vol.% CO2 experiments at different gas phase temperatures (10T0-130,10Tg-45 and 10Tg-60, see Table 4.1).

Page 87: Novel process concept for cryogenic CO2 capture

78 Mass deposition rates of carbon dioxide

0 200 400 600 8000.000

0.002

0.004

0.006

0.008

0.010

Ns = 100%

Ns = 75%

Ns = 50%

Mas

s dep

ositi

on ra

te [k

g·m

-2·s-1

]

Time [s]

(a)

0 200 400 600 800

0.000

0.001

0.002

0.003

0.004

0.005

0.006

Ns = 100%

Ns = 75%

Ns = 50%

Laye

r thi

ckne

ss [m

]

Time [s]

(b)

0 200 400 600 800

0

400

800

1200

1600 N

s = 100%

Ns = 75%

Ns = 50%

Laye

r den

sity

[kg·

m-3

]

Time [s]

(c)

Figure 4.12: The experimental (markers) and simulated (lines) mass depositionrate (a), frost layer thickness (b) and layer density (c) as function of the timefor the 10 vol.% CO2 experiments at different stirrer rotation speeds (10T0-130,10N75 and 10N50, see Table 4.1).

Page 88: Novel process concept for cryogenic CO2 capture

4.4 Development of a frost growth model 79

(a) (b)

Figure 4.13: Schematic representation of temperature (a) and CO2 concentration(b) profiles during the frost growth process.

A schematic representation of the involved processes together with theused variables is given in Fig. 4.13. The following assumptions are made:

• The temperature profiles within the layer are developed instanta-neously (quasi-steady state approach), i.e. the frost growth is muchslower than the temporal change of the temperature profiles withinthe layer.

• The desublimated CO2 at the frost surface is in equilibrium with thegas phase.

• The measurements showed that the density is changing in time. Itis likely that the formed ice layer has a porous structure, whichbecomes denser in time. The porosity, density and therefore alsolayer heat conductivity might be a function of the location withinthe frost layer. However, this information could not be obtainedfrom our experiments, therefore, it is assumed that the layer has auniform density and heat conductivity.

• The mass and heat transfer coefficient for the transfer from the gasbulk towards the frost surface are coupled according to the Chilton-Colburn analogy.

CO2 is transfered from the bulk of the gas phase to the solid phaseand deposits at the surface. Due to the removal of CO2 from the gas

Page 89: Novel process concept for cryogenic CO2 capture

80 Mass deposition rates of carbon dioxide

phase, a net flow towards the solid is generated and Fick’s diffusion is nolonger valid. Drift fluxes have to be taken into account:

NCO2= xgNtot + ctotk

g (xg − xσ) (4.1)

There is no net flux of N2, therefore:

NCO2= Ntot (4.2)

resulting in:

NCO2= ctotk

g

(xg − xσ)

(1− xg)with: k•g = kgξ (4.3)

and:

ξ =Φ

exp(Φ)− 1Φ =

NCO2

ctotkg(4.4)

and therefore:

k•g =NCO

2

ctot

1

exp(

NCO2

ctotkg

)

− 1(4.5)

Substituting Eq. 4.5 in Eq. 4.3 results in the following relation for thenet CO2 mole flux to the ice layer, which is known as Stefan flow(Taylor and Krishna, 1993):

NCO2= ctotkg ln

(

1− xσ

1− xg

)

(4.6)

The heat transfer from the gas to solid phase by conduction and by ra-diation, together with the heat formed by desublimation are conductedthrough the solid layer towards the cold surface. Again the effects of driftfluxes for the heat conduction through the gas towards the ice surfaceare taken into account:

αg (ξh +Φh) (Tg − T σ) + σǫ(

T 4g − T σ4

)

+NCO2∆Hs =

(

λeff

δ

)

(T σ − T0) (4.7)

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4.4 Development of a frost growth model 81

in which:

Φh =NCO

2Cp,CO

2

αgξh =

Φh

exp(Φh)− 1(4.8)

The concentration of CO2 at the frost surface follows from combining (4.6)and (4.7). The layer thickness is calculated by:

d (ρsδ)

dt= NCO

2MCO

2with:δ(t = 0) = 0 (4.9)

which can be solved with a standard ODE solver. Constitutive equationsare summarized in Table 4.2.

The density ρs and conductivity λeff of solid CO2 are unknown. As-suming these variables to be constant is not correct, as it was concludedfrom the experiments that the density is increasing in time, but reachesa plateau value for the measurements with pure CO2 in the gas phaserelatively quickly. For pure CO2, the CO2 concentration at the frost sur-face is equal to the gas phase concentration, which means that it can beexpected that the surface temperature will be constant during a measure-ment. For the measurements with mixtures, the surface temperature isnot constant, but will increase in time. Therefore, a correlation for thelayer density as a function of the surface temperature is proposed. Wheninspecting the measurements for different pure CO2 pressures in Fig. 4.6,it can indeed be observed that the measured density of the layer is higherfor measurements with a higher CO2 pressure in the gas phase (corre-sponding to a higher interface temperature). To investigate this relationin more detail, the following procedure is followed. Based on the mea-sured mass deposition rates, it is calculated using equation (4.6) whatthe concentration at the surface should have been. This concentrationcan again be used to calculate the surface temperature, because of equi-librium. In this way the (measured) density can be plotted as function ofthe calculated surface temperature for a large number of measurements,as shown in Fig. 4.14a. Although the results are somewhat scattered,there is a clear trend that the density is increasing when the surface tem-perature is increasing and leveling off at higher temperatures. Althoughnot measured, it can be expected that for lower temperatures the densitywill also level off. This behavior can be described using the following fit:

ρs =a

1 + b exp (−cT σ)(4.10)

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82 Mass deposition rates of carbon dioxide

Table 4.2: Constitutive equations used in the frost growth model.

The mass transfer coefficient kg is calculated according to:

Sh = aReb Scc in which is assumed: a = 0.1, b = 0.75, c = 0.33

in which b and c are taken from Winkelman et al. (2002), and a fitted to experi-mental results.The binary diffusivity for N2/CO2 mixtures is calculated according to the Fuller-Schettler-Giddings correlation (Fuller et al., 1966):

DAB =0.0143T

1.75(

1

1000MA+ 1

1000MB

)0.5

Ptot

[

(∑

υ)1/3A + (

υ)1/3B

]2

in which∑

υ is the sum of atomic diffusion volumes, which is 26.9 for CO2 and17.9 for N2.

The CO2 mole fraction at the interface is the equilibrium value at the surfacetemperature Tσ (Daubert and Danner, 1985):

xσ =P σ(T σ)

Ptot

=exp

(

10.257− 3082.7Tσ + 4.08 lnT σ − 2.2658 · 10−2T σ

)

Ptot

The gas to solid heat transfer coefficient αg is coupled to the mass transfer coeffi-cient kg, according to the Chilton-Colburn analogy:

Nu

Sh=

Pr1/3

Sc1/3

The physical properties of the gas phase were computed at the average gas phasetemperature according to Reid et al. (1987), using the pure component data sup-plied by Daubert and Danner (1985).

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4.4 Development of a frost growth model 83

150 160 170 1800

400

800

1200

1600 10T0-150 T0-160 10T0-140 T0-145 10T0-130 P100 50T0-150 P150 50T0-140 P200 50T0-130 P250 Fit

Fros

t den

sity

[kg·

m-3

]

Surface temperature [K]

(a)

0 400 800 1200 16000.0

0.2

0.4

0.6

0.8

1.0

1.2 10T0-150 T0-160 10T0-140 T0-145 10T0-130 P100 50T0-150 P150 50T0-140 P200 50T0-130 P250 Fit

Fros

t con

duct

ivity

[W·m

-1·K

-1]

Frost density [kg·m-3](b)

Figure 4.14: The measured density plotted as function of the calculated surfacetemperature (a) and the conductivity of the ice layer as function of the density (b).

with: a = 1.4768·103, b = 1.122·1015 and c = 2.133·10-1

The effective conductivity of the frost layer is very likely to be related tothe bulk density. For higher densities, one would expect higher conduc-tivities. In order to verify this, the surface temperature found by solvingequation (4.6) using the measured mass deposition rates, is substitutedinto the energy balance (4.7). When also the measured layer thickness issubstituted into the equation, the effective conductivity of the frost layercan be computed. The calculated conductivity is plotted as function ofthe measured density, again for a large number of measurements (seeFig. 4.14b). Although the results are scattered again, there is a cleartrend visible. This relation is described using a linear fit:

λeff = a+ bρs (4.11)

with: a = 6.200·10-3 (which is the CO2 gas conductivity) andb = 6.523·10-4.

4.4.2 Simulation results

The model developed in the previous section is used to simulate the massdeposition rate, layer thickness and frost density as a function of thetime, for all the measurements listed in Table 4.1. The outcomes have

Page 93: Novel process concept for cryogenic CO2 capture

84 Mass deposition rates of carbon dioxide

been plotted in the same figures as where the experimental results havebeen plotted. The general conclusion is that the experimental and simu-lation results agree very reasonably. When looking at the measurementsfor pure CO2, it can be observed that the fast decreasing mass depositionrates are very well described. Also the initially fast increasing frost den-sity to a constant value is well predicted by the model. It was observedduring experiments that the different pressures did not have a large ef-fect on the deposition rates and layer growth, as shown in Fig. 4.6. Thedifference in surface temperatures for the different measurements is toosmall to have a significant effect on the heat fluxes through the solid. Thesimulation results show comparable results. The cold plate temperaturehad a larger effect (Fig. 4.7), which is again well predicted by the model.

The differences between the measurements with different CO2 concen-trations, as shown in Fig. 4.8, are very well described by the developedmodel. The different measurements for the 50 vol.% and 10 vol.% mix-tures are in general reasonably described. The mass deposition rates arewell described, although the discrepancies in the measured and the com-puted frost density and layer thickness are sometimes somewhat larger,which can be well attributed to the inaccuracies in the derived correla-tions for the layer density and conductivity, and the inaccuracy in themeasurements. However, the calculated trends are very well matchingthe trends found in experiments.

To show the importance of taking drift fluxes into account, the massdeposition rates have been calculated for experiment 50T0-140 and 10T0-

140 with and without including drift effects. It can be observed inFig. 4.15 that the effect is quite small for the 10 vol.% CO2 measure-ment, but has a substantial effect for the 50 vol.% CO2 measurement.Furthermore, it is noted that the contribution due to radiation in theenergy balance (4.7) is negligible.

4.4.3 Significance for the cryogenic packed bed concept

The conditions of the stirred cell are different from those encountered incryogenically refrigerated packed beds. The cold plate temperature in thestirred cell is kept at a constant temperature by continuous cooling, whilein the packed bed the temperature of the packing material will increasein time when a front of desublimating CO2 passes by. In order to showthe significance of the developed model for the mass deposition rates,

Page 94: Novel process concept for cryogenic CO2 capture

4.4 Development of a frost growth model 85

0 200 400 600 800

0.000

0.005

0.010

0.015

0.020 50T0-140 10T0-140 50 Incl. drift 10 Incl. drift 50 No drift 10 No drift

Mas

s dep

ositi

on ra

te [k

g·m

-2·s-1

]

Time [s]

Figure 4.15: The experimental mass deposition rates for experiments 50T0-140

and 10T0-140 (markers) and the the simulated profiles (lines), with and withouttaking drift fluxes into account.

the next situation will be evaluated: A refrigerated spherical particle, ini-tially at Tp,0 is suddenly positioned in a gas phase of temperature Tg andCO2 mole fraction of xCO

2,g, which is the equilibrium concentration at

the temperature Tg. This is similar to the situation in the packed bed atthe frost front, where saturated gas will be contacted with cold packing.Due to this step change in temperature and composition, CO2 will startto desublimate at the particle surface, and the particle will heat until thesystem is in equilibrium, i.e. the particle temperature Tp is equal to thegas phase temperature. The temperature increase of the particle can bedescribed by:

ρpVpCp,pdTp

dt= (Φ′′

h +NCO2MCO

2∆Hs)Ap (4.12)

where it is assumed that the heat capacity of the desublimated CO2 ismuch smaller than the heat capacity of the particle. The particle is be-ing heated by desublimation of CO2 and by heat transfer from the gasphase to the ice layer (Φ′′

h). As given in (4.7), this amount is transferred

Page 95: Novel process concept for cryogenic CO2 capture

86 Mass deposition rates of carbon dioxide

Table 4.3: Properties and conditions used for a case study in which CO2 desubli-mates onto a spherical particle.

Particle diameter [m] 0.004Particle material [-] glassSolids density [kg ·m−3] 2546Heat capacity [J · kg−1K−1] −2.8 · 10−3T 2 + 3.48T − 47.9Initial particle temperature [◦C] −140Gas phase temperature [◦C] −103.3Gas phase mole fraction CO2 [-] 0.1Gas mass flow [kg ·m−2 · s−1] 0.249

through the solid CO2 layer towards the particle, and therefore (4.12) canbe written as:

dTp

dt=

λeff

δ(T σ − Tp)

6

dp

1

ρpCp,p(4.13)

This heat balance coupled with the frost growth model are solved nu-merically using a standard ODE solver. The frost growth at the sphericalparticle is computed using the conditions and properties as given in Ta-ble 4.3, which are similar to those used in the experiments described inChapter 3. The mass and heat transfer coefficients are calculated for theconditions as occurring in the packed bed, according to Gunn (1978).

The computed temperature of the particle as function of the time isshown in Fig. 4.16. It can be observed that equilibrium is already reachedafter approximately 5 seconds. The layer of solid CO2 formed at the par-ticle surface at that moment is only 3.24·10-5 m thick (corresponding to57.1 kg/m3

packing). Therefore, the heat transfer through the solid layerplays a negligible role and the frost surface temperature is approximatelyequal to the particle temperature. To illustrate this in more detail, thetemperature increase of the particle has been calculated without takinglayer formation into account. The concentration of CO2 at the frost sur-face is calculated using the particle temperature. Fig. 4.16 indeed showsthat the computed temperature rise of the particle is hardly influenced.Furthermore, it should be noted that heat transfer from the gas to solidphase plays a minor role, meaning that the mass deposition rate for theconditions used in this case are mainly determined by mass transfer ofCO2 from the gas to the solid phase.

Page 96: Novel process concept for cryogenic CO2 capture

4.5 Discussion and conclusions 87

It was shown in Chapter 3 that the axial temperature profiles withinthe packed beds could be well described when assuming the followingmass deposition rate equation:

NCO2MCO

2= g (xgPtot − P σ) (4.14)

with g = 1·10-6 s/m. The increase in temperature of the particle is alsocalculated using this simplified mass deposition rate for different val-ues of g. Fig. 4.16 shows that indeed a value slightly above 1·10-6 s/mmatches the temperature profile calculated using the frost growth modelbest. It was shown in Fig. 4.15 that for low concentrations of CO2 driftfluxes have negligible effects on the mass deposition rate. Therefore themass deposition rate could be written as:

NCO2= kgctot (xg − xσ) (4.15)

Combination of (4.14) and (4.15) and using the ideal gas law to couplethe concentration and the pressure, yields:

g =kgMCO

2

RT(4.16)

The values of kg and the average temperature in the film layer aroundthe particle T are slightly changing in time. When taking the averagevalues (kg = 4.45· 10-2 m/s and T = -105.5◦C) the value of g amounts 1.4·10-6 s/m, which is close to the assumed value of g.

Finally, it is noted that the heat transfer within the particle itself isrelatively fast, but could play a role for glass particles. Fig. 4.17 showsthat the temperature in a spherical particle is well developed for a Fouriernumber ( at

R2 ) of 0.4, which corresponds to 2.2 seconds for a glass particlewith a diameter of 4 mm. Note that for steel particles or monoliths thiseffect plays no role.

4.5 Discussion and conclusions

This chapter has presented a study on the the mass deposition rates ofCO2 onto cold surfaces. A dedicated experimental setup was designedto measure CO2 deposition rates under well defined conditions. The re-sults showed that heat transfer through the formed frost layer plays animportant role in the deposition rates. Furthermore, it was shown that

Page 97: Novel process concept for cryogenic CO2 capture

88 Mass deposition rates of carbon dioxide

0 4 8 12 16 20

-140

-120

-100

-80 g = 1·10-5 Frost growth model g = 1·10-6 No layer g = 1·10-7

Parti

cle

tem

pera

ture

[°C

]

Time [s]

Figure 4.16: Temperature increase calculated for a cold particle positioned in aN2/CO2 mixture at t = 0, using different expressions for the mass deposition rate.

Figure 4.17: Unsteady temperature profiles in a spherical particle(Carslaw and Jaeger, 1952).

Page 98: Novel process concept for cryogenic CO2 capture

4.5 Discussion and conclusions 89

diluting CO2 with N2 has a large effect on the mass deposition rate, due tothe introduction of mass transfer limitations from the gas bulk towardsthe frost surface. To describe the results, a frost growth model was de-veloped. Based on the experimental results, fitted correlations for thebulk density of the frost layer as function of the frost surface tempera-ture and a correlation for the effective layer conductivity as a functionof the frost density were derived. Using these correlations in the frostgrowth model, the developed model could well describe the experimentalfindings of the CO2 mass deposition rate as a function of time at differentCO2 concentrations and gas and cooled surface temperatures. To showthe significance of the results for the cryogenic packed bed concept, thetemperature increase of a cold particle suddenly positioned in a N2/CO2

mixture was calculated using the frost growth model. It was shown thatunder the packed bed conditions, heat transfer through the solid frostlayer plays no significant role and the process was mainly determinedby mass transfer of CO2 towards the particle, which coincides with theearlier findings reported in Chapter 3. It should be noted that for othercryogenic CO2 capture technologies based on continuously cooled heatexchanging surfaces (such as proposed by Clodic and Younes (2002)), theincreasing frost layer does play a significant role for the frost formationrates. Therefore the frost growth model developed in this chapter couldbe a valuable tool in the design and study of these types of equipment.

Acknowledgment

The contributions of Nhi Dang and Niels Hietberg to this chapter arehighly appreciated.

Notation

A area, [m2]a thermal diffusivity, [m2/s]c concentration, [mol/m3]Cp heat capacity, [J/kg/K]D diffusion coefficient, [m2/s]d diameter, [m]g mass deposition rate constant, [s/m]

Page 99: Novel process concept for cryogenic CO2 capture

90 Mass deposition rates of carbon dioxide

kg gas solid mass transfer coefficient, [m/s]M molecular weight, [kg/mol]N molar flux, [mol/m2/s]Ns stirrer rotation speed, [1/s]Nu Nusselt number (αgd/λg)P pressure, [Pa, mbar]Pr Prandtl number (Cp,gηg/λg)R gas constant, [J/mol/K]Re Reynolds number

(

ρgNsd2s/ηg

)

or (ρgvgdp/ηg)Sc Schmidt number (ηg/ρg/D)Sh Sherwood number (kgd/D)t time, [s]T temperature, [K], [◦C]V volume, [m3]v superficial velocity, [m/s]x mole fraction, [-]

Greek letters

αg gas solid heat transfer coefficient, [W/m2/K]δ layer thickness, [m]∆H sublimation enthalpy, [J/mol]ǫ emissivity, [-]η viscosity, [kg/m/s]λ thermal conductivity, [W/m/K]λeff effective conductivity frost layer , [W/m/K]ξ correction factor for drift flux, [-]ρ density, [kg/m3]υ atomic diffusion volume, [-]σ Stefan-Boltzmann constant, [J/s/m2/K4]Φ correction factor for drift flux, [-]Φ′′

h gas solid heat flux, [J/m2/s]

Subscripts

0 cold plate, initialg gas phasep particles solid phasetot total

Page 100: Novel process concept for cryogenic CO2 capture

4.5 Discussion and conclusions 91

Superscripts

σ equilibrium

Page 101: Novel process concept for cryogenic CO2 capture

92 Mass deposition rates of carbon dioxide

Page 102: Novel process concept for cryogenic CO2 capture

5Experimental demonstration of the novelprocess concept in a pilot scale setup

Abstract

The proposed novel process concept to cryogenically capture CO2 fromflue gases using dynamically operated packed beds consists of a cap-ture, recovery and cooling step. The capture step was extensively studiedboth by experiments in a small scale setup as well as simulations inChapter 3. This chapter describes a fully automated and continuouslyoperated experimental pilot plant, in which the entire process cycle isdemonstrated, including the capture, recovery and cooling step. Tests ofmore than 10 hours showed that the process can be continuously oper-ated for long periods. The temperatures measured in the beds during thethree process steps have been compared to simulation results obtainedwith the one-dimensional pseudo-homogeneous numerical model devel-oped in Chapter 2. Breakthrough times can be well described. However,it was observed that the temperatures measured at a radial position closeto the wall differed from the temperatures measured in the radial center.Therefore, the model was extended with an additional energy balance forthe wall, where heat exchange with the packing is taken into account. It

Page 103: Novel process concept for cryogenic CO2 capture

94 Experimental demonstration in a pilot scale setup

is shown for the cooling step that the high conductivity of the wall causesthe temperature profiles to be very smooth close to the wall relative to theprofiles in the center. When including a separate energy balance for thewall the measured profiles could be well described by the model.

5.1 Introduction

Cryogenic CO2 capture using dynamically operated packed beds requiresthree process steps: the capture, recovery and cooling step. The capturestep was demonstrated in Chapter 3 in a small scale experimental setup.This chapter gives a demonstration of the entire process cycle includingall required process steps. First, the experimental setup and procedureare described, followed by the results for long term experiments. Subse-quently, the temperature profiles at different locations within the beds arestudied in more detail for all three individual process steps and are com-pared to simulation results of the one-dimensional pseudo-homogeneousnumerical model described in Chapter 2. Finally, the chapter concludeswith the description of an extension of the numerical model, in order toprovide better understanding of the observed temperature profiles duringthe experiments related to the energy transport to and in the steel vesselwall.

5.2 Experimental

This section provides a description of the constructed experimental setupand the experimental procedure followed in the experiments.

5.2.1 Experimental setup

A schematic flow diagram of the experimental rig is shown in Fig. 5.1.Three identical stainless steel vacuum insulated vessels (OD × ID × L =106 × 100 × 500 mm) were filled with spherical glass beads (dp = 3 mm).All three beds were equipped with 9 K-type thermocouples at several ax-ial and radial positions (see Fig. 5.2) in order to monitor the developmentof temperature profiles within the beds. Pressure sensors were installedboth at the inlet as outlet of all three beds to measure the pressure dropsover the beds. Cooling of the beds was performed by feeding a cold N2

Page 104: Novel process concept for cryogenic CO2 capture

5.2 Experimental 95

gas flow, which was refrigerated before using a cryogenic cooler (StirlingSPC-1). The cooler was connected to the beds using polystyrene insu-lated pipes. Mixtures of N2 and CO2 were fed to the refrigerated bed, ofwhich the composition was controlled by using two mass flow controllers(Brooks). Before feeding the gas mixture to a bed, the gas was heatedup using a tracing line winded around the feed tube. The recovery stepwas performed by recycling CO2 using a 3kW side channel blower (VA-COM SC 425). At initial stages of recovery steps, some additional CO2

deposited onto the packing as explained in Chapter 2. Therefore, a CO2

buffer was required, which consists of three stainless steel vessels (totalvolume of 36 liter). Low temperature CO2 was not allowed to enter theblower, therefore a tracing line has been installed around the blower inlettube to heat up the incoming gas to ambient temperature. Furthermore,all beds could be flushed with heated N2, if so required. All valves arepneumatic ball valves, special cryogenic valves (Meca-inox) were used atthose positions where low temperatures were prevailing. Signal process-ing and process control was carried out by NI Labview. A picture of theexperimental setup is shown in Fig. 5.3.

5.2.2 Experimental procedure

Before an actual measurement was started, the N2/CO2 mixtures wereprepared and heated up, the cooler was started up and the CO2 buffervessels were filled. After these preparations, the first bed (bed #1) wascooled down until the average temperature of two thermocouples (r =0 cm and r = 4 cm) at the outlet of the bed reached a certain set pointtemperature. At that point, the second bed (bed #2) was fed with thecold N2 flow, and at the same time bed #1 was fed with the N2/CO2 gasmixture. The temperature at the outlet of a bed will increase when CO2

starts to break through. Therefore again a certain set point temperaturewas used to monitor whether a bed finished a capture step. At the pointthat CO2 breakthrough occurred in bed #1, and bed #2 was cooled down,bed #1 was switched to the recovery step, bed #2 to the capture step andbed #3 to the cooling step. The recovery step was started by switchingon the recycle blower for CO2. Due to the additional CO2 which desubli-mates initially onto the packing surface in a relative short time, the pres-sure decreases. This was counteracted by opening the valve connectedto the CO2 buffer vessels. Fresh CO2 flows into the recovery piping, until

Page 105: Novel process concept for cryogenic CO2 capture

96 Experimental demonstration in a pilot scale setup

the additional desublimation of CO2 stopped and the pressure reacheda certain set point. At that point CO2 was recycled, and the previouslydeposited CO2 in the bed was evaporated and removed from the bed. Theoutlet flow is then larger than the inlet flow, therefore the pressure startsto increase. This is again counteracted, in this case by opening a massflow controller, controlled using a PID control loop. Again the temper-ature at the outlet of the bed is used to determine the point to switch.When all CO2 is removed from the packing, the temperature will imme-diately start to increase and at that point the bed was switched again toa cooling step. In case one of the beds finished its step before the otherbeds were finished, the bed was put on a standby mode and switchedagain to the next process step as soon as the other beds were finished.

Page 106: Novel process concept for cryogenic CO2 capture

5.2

Exp

erim

en

tal

97

Figure 5.1: Simplified flow scheme of the pilot test setup, in which TI, PI and FI stand for temperature,pressure and flow indication respectively.

Page 107: Novel process concept for cryogenic CO2 capture

98 Experimental demonstration in a pilot scale setup

5.3 Experimental results

Experiments have been carried out using the conditions as listed in Ta-ble 5.1. Long term runs of more than 10 hours proved that the systemis able to run stable for a long period of time. The average temperatureof the thermocouples at r = 0 cm and r = 2 cm at the outlets of the threebeds as function of the time is plotted for 6 hours in Fig. 5.4a. Similar re-peating temperature patterns for the three beds can be observed. To havea closer look, the temperature profile is zoomed in for one bed, showing acooling, capture and recovery step in more detail (see Fig. 5.4b). Duringthe capture step, the temperature at the outlet of the bed is almost sta-ble at the initial bed temperature (approximately -130◦C), but increasesstrongly as soon as CO2 starts to breakthrough. When the bed is at atemperature of -100◦C the capture step is stopped. When the recoverystep is started, it can be observed that the temperature increases quicklyto a temperature of approximately -76◦C corresponding to a saturationtemperature of pure CO2 at the operating pressure (1.2 bar). Again itcan be observed that the temperature increases at the end of the recov-ery step. At the point that the average outlet temperature reaches -70◦C,the bed is switched to a cooling step. During the cooling step, the out-let temperature first rises to a temperature of about 20◦C. This is thetemperature at which the CO2 is fed to the bed during the recovery step.Note that for commercial operation it is better to feed the recovery CO2

gas flow at lower temperatures (-70◦C) to save cooling duty, but due topractical limitations of the used blower, this could not be realized in theexperimental situation. After this relatively hot zone of 20◦C has movedthrough the bed, the temperature at the outlet of the bed decreases, untilthe set point is reached (-130◦C).

Fig. 5.5a shows the pressure at the outlet of one bed, again zoomedin for one process cycle with the three steps. During the capture step,the bed is at atmospheric pressure. As soon as the bed is switched to therecovery step, it can be observed that the pressure suddenly drops to alow value. As explained earlier, this is related to the gaseous CO2 presentin the piping of the recovery step which is depositing onto the packingsurface. Immediately the valve connected to the CO2 buffer vessels isopened to allow the pressure to increase again and additional CO2 is de-posited at the packing. At that point the valve is closed, and the pressureincreases and is controlled at 1.2 bar, by allowing CO2 to leave the system

Page 108: Novel process concept for cryogenic CO2 capture

5.3 Experimental results 99

Figure 5.2: Thermocouple positioning in the packed bed.

Page 109: Novel process concept for cryogenic CO2 capture

100 Experimental demonstration in a pilot scale setup

Figure 5.3: Picture of the experimental pilot setup.

Page 110: Novel process concept for cryogenic CO2 capture

5.4 Simulations 101

Table 5.1: Conditions as used for experiments in the pilot setup.

Capture Recovery Cooling

N2 flow [NL/min] 22 - 285CO2 flow [NL/min] 5.5 220 -Total flow [NL/min] 27.5 220 285

CO2 gas mole fraction [-] 0.2 1 0

Inlet gas temperature [◦C] 100 20 -155Set point temperature [◦C] -100 -70 -130

Step time [min] 15 15 15

via a mass flow controller. The flow through this mass flow controller isplotted in Fig. 5.5b. Note that initially the pressure is well maintained at1.2 bar, but that at some point the pressure increases further to approx-imately 1.28 bar. This is related to the fact that the mass flow controlleris fully opened (15 NL/min) as observed in Fig. 5.5b. During the coolingstep, a pressure of 1.2 bar is measured, which is slightly changing duringthe step, which is likely related to the changing gas phase temperature inthe bed.

During all steps, the pressures at the inlets and outlets of the bed havebeen monitored. It was observed that the pressures at the inlet and outletshow virtually equal pressures during all three steps, meaning that thepressure drop over the packed bed is negligible compared to the pressuredrop over piping/valves etc.

5.4 Simulations

In this section the experimental results are studied in more detail andhave been compared to the simulation results of the numerical modeldeveloped in Chapter 2. Subsequently, the model will be extended with aseparate heat balance for the wall, to provinde a better understanding ofthe observed behavior.

Page 111: Novel process concept for cryogenic CO2 capture

102 Experimental demonstration in a pilot scale setup

0 50 100 150 200 250 300 350

-120

-80

-40

0

40 Bed #1 Bed #2 Bed #3

Tem

pera

ture

[°C

]

Time [min]

(a)

110 120 130 140 150 160

-120

-80

-40

0

40 CoolingRecovery

Tem

pera

ture

[°C

]

Time [min]

Capture

(b)

Figure 5.4: Evolution of the temperature at the outlets of the three beds for a longrun (a) and zoomed in for one process cycle for one bed (b).

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5.4 Simulations 103

110 120 130 140 150 160

0.6

0.8

1.0

1.2

1.4 CoolingRecoveryCapture

Pres

sure

[bar

]

Time [min](a)

110 120 130 140 150 1600

5

10

15

20

25Recovery bed #2

Recovery bed #1

Recovery bed #3

CO

2 ven

t [N

L/m

in]

Time [min](b)

Figure 5.5: Pressure at the outlet of bed #1 during the three process steps (a) andthe CO2 gas flow towards the vent during the recovery steps of the three beds (b).

5.4.1 Radial temperature profiles

The experimental temperature development during the capture step isshown for two different radial positions (r = 0 cm and r = 4 cm) for twoaxial locations: z = 16 cm (Fig. 5.6a) and z = 28 cm (Fig. 5.6b). It can beobserved that the temperature close to the wall (r = 4 cm) is initially higherthan the temperature in the radial center. The effect on the temperaturedevelopment is that the front of freezing CO2 is moving faster throughthe bed close to the wall. These experimental results are compared withsimulations with the one-dimensional model (see Chapter 2, assumingan initial bed temperature averaged over the two radial positions. Theresulting breakthrough times match well with the experimental outcomesand is exactly positioned between the two measured temperatures.

The evolution of the temperature has also been plotted for the recov-ery step for z = 16 cm (Fig. 5.7a) and z = 28 cm (Fig. 5.7b). A largedifference in the temperature profiles between the two radial positionscan be observed. The simulated temperature profile is again in betweenthe measured profiles. Remarkably, the CO2 is removed faster from theradial center than from the zone closer to the walls. Based on the cap-ture step, one would expect lower mass deposition of CO2 close to thewalls (because of the lower initial bed temperature) and therefore faster

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104 Experimental demonstration in a pilot scale setup

0 2 4 6 8 10 12 14

-160

-140

-120

-100

-80

-60

-40 measured, r = 0 cm measured, r = 4 cm simulated

Tem

pera

ture

[°C

]

Time [min](a) z = 16 cm

0 2 4 6 8 10 12 14

-160

-140

-120

-100

-80

-60

-40 measured, r = 0 cm measured, r = 4 cm simulated

Tem

pera

ture

[°C

]Time [min]

(b) z = 28 cm

Figure 5.6: Temperatures measured in the radial center (r = 0 cm) and close tothe wall (r = 4 cm) and the simulated temperature for two axial locations: z = 16cm (a) and z = 28 cm (b) during the capture step.

removal. Apparently, the wall is playing a significant role, which is alsovisible for the cooling step, discussed below.

Finally, the temperature development is shown for two axial locations:z = 16 cm (Fig. 5.8a) and z = 28 cm (Fig. 5.8b) for the cooling step. Thetemperature measured close to the wall shows a much more dispersedtemperature profile. Therefore, the zones close to the wall require muchlonger cooling times in order to reach a low temperature. The simulationoutcomes show that the temperature development is again in betweenthe profiles measured at the two radial positions, but the profile is assteep as the profile measured in the radial center. This is attributed tothe effect of the vessel wall, which will be studied in more detail in thefollowing section.

5.4.2 Influence of the wall

The previous section showed that the one-dimensional numerical modelcan predict the average breakthrough times, but is not able to explain thedifferences between the temperature measurements in the radial centerand close to the wall of the bed. The difference in temperature profiles be-tween these locations is particularly pronounced for the cooling step. Thetemperature development close to the wall is very disperse, which cannot

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5.4 Simulations 105

0 2 4 6 8 10 12 14-120

-80

-40

0

40

80 measured, r = 0 cm measured, r = 4 cm simulated

Tem

pera

ture

[°C

]

Time [min](a) z = 16 cm

0 2 4 6 8 10 12 14-120

-80

-40

0

40

80 measured, r = 0 cm measured, r = 4 cm simulated

Tem

pera

ture

[°C

]Time [min]

(b) z = 28 cm

Figure 5.7: Temperatures measured in the radial center (r = 0 cm) and close tothe wall (r = 4cm) and the simulated temperature for two axial locations: z = 16cm (a) and z = 28 cm (b) during the recovery step.

0 2 4 6 8 10 12 14-160

-120

-80

-40

0

40

80

120 measured, r = 0 cm measured, r = 4 cm simulated

Tem

pera

ture

[°C

]

Time [min](a) z = 16 cm

0 2 4 6 8 10 12 14-160

-120

-80

-40

0

40

80

120 measured, r = 0 cm measured, r = 4 cm simulated

Tem

pera

ture

[°C

]

Time [min](b) z = 28 cm

Figure 5.8: Temperatures measured in the radial center (r = 0 cm) and close tothe wall (r = 4 cm) and the simulated temperature for two axial locations: z = 16cm (a) and z = 28 cm (b) during the cooling step.

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106 Experimental demonstration in a pilot scale setup

be predicted by the one-dimensional model. It is very well possible thatthe axial heat conduction in the wall has an influence on the temperaturein the region close to the wall. A detailed description of the temperaturein the bed as function of the radial and axial position can be obtained byextending the model to two dimensions. However, it is also possible to getmore insight in the behavior by adjusting the one-dimensional model. Instead of accounting for the heat capacity of the wall in the accumulationterm, an extra energy balance for the wall is included in the model, inwhich heat is exchanged with the packed bed:

(ρwCp,w)∂Tw

∂t=

∂z

(

λw∂T

∂z

)

− awαgw (Tw − Tg) (5.1)

The energy balance for the bed (both the gas and the solid phase) willtherefore also be extended with an extra contribution due to the heatexchange with the wall:

(εgρgCp,g + ρs(1− εg)Cp,s)∂Tg

∂t= −ρgvgCp,g

∂Tg

∂z+

∂z

(

λeff

∂Tg

∂z

)

nc∑

i=1

m′′

i as∆Hi + abαgw (Tw − Tg) (5.2)

in which the expressions used to calculate the bed to wall heat transfercoefficient (αgw) are summarized by Tiemersma (2010).

Simulations have been carried out assuming that the initial wall tem-perature is equal to the temperature measured close to the wall, and thatthe packing has an initial temperature as measured in the radial center.The temperature development for both the wall as well as the bed havebeen plotted as function of the time for z = 16 cm in Fig. 5.9. It can beclearly observed that the temperature within the wall is very dispersed,which is explained by a much higher value of the conductivity of the solidsteel wall compared to the effective conductivity within the bed. Basedon this result, the dispersed measured temperature close to the wall canbe better understood. Accounting for the wall separately, also leads toa better prediction of the temperature within the packing in the radialcenter of the bed.

The wall will also influence the temperature development during thecapture and recovery step. It is very well possible that CO2 will depositat the tube wall (which is not accounted for in the extended model). This

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5.5 Discussion and conclusions 107

0 2 4 6 8 10 12 14-160

-120

-80

-40

0

40

80

120 measured, r = 0 cm measured, r = 4 cm simulated, r = 0 cm simulated wall

Tem

pera

ture

[°C

]

Time [min]

Figure 5.9: Effect of accounting for the wall separately in simulations for thetemperature development at z = 16 cm during the cooling step.

could well explain the slower removal of CO2 during the recovery stepclose to the wall. Locally there will be more CO2 deposited, due to theheat capacity of the wall and therefore it will take longer to remove solidCO2 from those zones. An additional effect may be that the pressure dropwill locally increase due to the higher amount of deposited CO2, resultingin lower flow rates and therefore even slower solid CO2 removal rates.

5.5 Discussion and conclusions

Chapter 3 showed the experimental demonstration of the capture step ina small scale setup. In this chapter the entire process cycle including thecooling, recovery and capture step was demonstrated in an advanced fullyautomated experimental pilot setup. Three beds were operated in paral-lel, therefore enabling the option to operate the process continuously.The numerical model developed in Chapter 2 was again able to describethe temperature profiles within the beds, also during the recovery andcooling step. However, the walls of the packed bed caused temperaturedifferences measured in the radial center and close to the wall and also

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108 Experimental demonstration in a pilot scale setup

a much more dispersed temperature front closer to the wall, especiallyfor the cooling step. To get better understanding of the wall effects, thenumerical model was extended for the cooling step with an additionalenergy balance for the wall, accounting for the heat exchange with thepacking. The simulation results showed that the temperature develop-ment within the bed can indeed be better described. Note that these walleffects are important in the interpretation of the experimental results ofthe pilot setup, but can be neglected for beds with large diameters, wherethe volume of the wall is much smaller compared to the bed volume.

Acknowledgment

Jeroen Zijp is kindly acknowledged for his substantial contribution to thedesign of the experimental setup described in this chapter.

Notation

a specific surface area, [m2/m3]Cp heat capacity, [J/kg/K]dp particle diameter, [mm]ID inner diameter, [mm]L bed length, [mm]m′′

i mass deposition rate per unit surface area for component i,[kg/m2/s]

nc number of components, [-]OD outer diameter, [mm]r radial coordinate, [m]t time, [s]T temperature, [K], [◦C]v superficial velocity, [m/s]z axial coordinate, [m]

Greek letters

αgw heat transfer coefficient gas to wall, [W/m2/K]∆Hi enthalpy change related to the phase change of component i,

[J/kg]εg bed void fraction, [-]

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5.5 Discussion and conclusions 109

λ thermal conductivity, [W/m/K]λeff effective conductivity, [W/m/K]ρ density, [kg/m3]

Subscripts

g gas phasei component is solid phasew wall

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110 Experimental demonstration in a pilot scale setup

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6Techno-economic evaluation

Abstract

A techno-economic evaluation of the proposed process cryogenic conceptis presented in this chapter. A basic process design focusing on theCO2/N2 separation for a 600 MW coal fired power plant is given and theCO2 avoidance costs have been calculated. The influence of several pro-cess parameters have been investigated: lower initial bed temperaturesand higher CO2 concentrations in the feed result in more efficient use ofthe bed volume. The pressure drop over the system plays an importantrole in the process economics, due to the high flow rates required in theprocess. The cryogenic concept is compared to two competing technolo-gies: amine absorption and membrane separation. The results show thatthe preferred technology highly depends on assumptions related to theavailability of utilities. If the required cold duty is available at a rela-tively low price, e.g. via integration with a LNG regasification station, the

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112 Techno-economic evaluation

proposed cryogenic process concept can well compete with alternativetechnologies.1

6.1 Introduction

Several economic studies to CO2 capture methods have been publishedin literature. Resulting costs vary strongly, as they are highly influencedby the system boundaries such as the CO2 source and therefore inletconcentration, whether or not transport and storage is included, the levelof maturity and cost measures and assumptions. For example, the opti-mized costs in a study by Abu-Zahra et al. (2007a,b) to CO2 capture by30% MEA absorption from a 600 MW bituminous coal-fired power planthave been estimated at 33 e/ton CO2 avoided. On the other hand, in astudy by van Straelen et al. (2010) to CO2 capture from a refinery also us-ing 30% MEA, costs of 90-120 e/ton CO2 avoided were reported, depen-dent on the scale and the CO2 concentration in the flue gas. Merkel et al.(2010) evaluated a process based on CO2 capture using membranes andcalculated CO2 capture costs of $39/ton CO2. In a report by McKinsey(2008) the development of costs for CCS (including storage costs) is an-alyzed over the next twenty years. They expect that early demonstrationprojects will operate at 60-90 e/ton, but that costs could come down to30-45 e/ton in 2030, a price level which is expected to make CCS eco-nomically self sustaining. More research is required to bring down thecosts. Although many studies focus on reducing operational costs, e.g.by finding novel more efficient solvents for amine scrubbing, it is at leastas important to reduce capital costs in order to reduce CO2 avoidancecosts (Schach et al., 2010).

The aim of the work described in this chapter is to evaluate the Cryo-genic Packed Bed (CPB) concept, both on technical aspects as well as oneconomic performance. Furthermore, the economics of the CPB conceptare compared to other post-combustion technologies, viz. amine scrub-bing and membrane technology, investigating the importance of variousprocess assumptions. The chapter is organized as follows: first a basecase is defined and the costs per ton of CO2 emissions avoided are calcu-lated. Subsequently, a sensitivity analysis of some key process parame-

1This chapter is based on: Tuinier et al., Techno-economic evaluation of cryogenic CO2capture - A comparison with absorption and membrane technology. Accepted for publica-tion in the Int. J. of Greenhouse Gas Control, doi:10.1016/j.ijggc.2011.08.013.

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6.2 Process evaluation 113

ters is discussed. Finally, the results will be compared to the results ofeconomic studies on CO2 capture via amine scrubbing (Rubin, 2010) andmembrane technology (Merkel et al., 2010).

6.2 Process evaluation

In order to be able to compare the CPB concept with other technologies,a basic design was made for a capture plant treating flue gas typicallygenerated by a 600 MW coal fired power plant, which is often used as abase case in literature studies. The bed dimensions and required processconditions are obtained by carrying out simulations using the detailedmodel, which is described in Chapter 2. The capital and operation costsare then estimated and the costs per ton of CO2 avoided are calculated.This section ends with a parameter study, in which the influence of sev-eral key parameters on the capture costs is evaluated.

6.2.1 Base case

To simplify the comparison, only CO2 capture is taken into account with-out impurities and H2O removal. The assumed flue gas conditions andcomposition are shown in Table 6.1. The bed dimensions and propertiesfor the base case are detailed in Table 6.2. The initial bed temperaturewas set at -150◦C, which results in more than 99.9% CO2 recovery. Abreakthrough time (duration of each step) of 600 seconds was chosen.The required flow rates, pressures and inlet temperatures are listed inTable 6.3. The resulting pressure drops over the beds (also shown inTable 6.3) are rather small due to the nature of the selected packing;a structured monolith. However, gas distribution over the beds, pipingand valves will cause an additional pressure drop, therefore a total pres-sure drop of 100 mbar is assumed for the capture step and 200 mbarfor the cooling and recovery step (because of the higher volumetric flowrates). During the recovery step, the outgoing CO2 flow has a tempera-ture of -78◦C and will be partly recycled to the inlet. Due to a temperatureincrease during the compression by the recycle blower the inlet tempera-ture during the recovery step is increased to -66◦C. The flue gas temper-ature is estimated at 150◦C, but will increase in temperature to 162◦C,also because of compression. The resulting axial temperature and massdeposition profiles are shown in Fig. 6.1. It can be observed that the heat

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114 Techno-economic evaluation

stored in the bed during the capture step is being used during the recov-ery step to evaporate previously deposited CO2. Furthermore, it can beobserved that during the cooling step not the entire bed has to be cooleddown, as the last part of the bed will be cooled down during the capturestep.

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6.2

Pro

cess

eva

lua

tion

115

0 1 2 3 4

-150

-100

-50

0

50

100

150 0 s 300 s 600 s

Tem

pera

ture

[°C

]

Axial position [m](a)

0 1 2 3 4

-150

-100

-50

0

50

100

150 0 s 300 s 600 s

Tem

pera

ture

[°C

]

Axial position [m](b)

0 1 2 3 4

-150

-100

-50

0

50

100

150 0 s 300 s 600 s

Tem

pera

ture

[°C

]

Axial position [m](c)

0 1 2 3 40

20

40

60

80

100

0 s 300 s 600 s

Mas

s dep

ositi

on [k

g·m

-3]

Axial position [m](d)

0 1 2 3 40

20

40

60

80

100

0 s 300 s 600 s

Mas

s dep

ositi

on [k

g·m

-3]

Axial position [m](e)

0 1 2 3 40

20

40

60

80

100

Mas

s dep

ositi

on [k

g·m

-3]

Axial position [m](f)

Figure 6.1: Simulated axial temperature (a - c) and mass deposition (d - f) profiles for the capture, recoveryand cooling step. Operating conditions and bed properties can be found in Table 6.1, 6.2 and 6.3.

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116 Techno-economic evaluation

Table 6.1: Flue gas conditions and composition.

Temperature [◦C] 150Pressure [bar] 1.013

Vol.% Flow [kg/s]N2 86.5 510CO2 13.5 125

Total 635

Table 6.2: Bed dimensions and properties.

Diameter [m] 8.5Length [m] 4.25Number of beds [-] 21 (7 per step)Packing Steel monolith structureDensity solids [kg/m3] 7750Porosity [-] 0.697

6.2.2 Costs base case

In order to calculate the CO2 avoidance costs, the capital investmentcosts are first calculated using a conceptual cost estimation method withan accuracy of 40%. In this method, the main equipment costs areestimated. Fig. 6.2 shows a simplified process scheme with all mainequipment. The costs for blowers, the heat exchanger and the columnshave been calculated using correlations reported by Seider et al. (2004)and Loh and Lyons (2002) and were updated to costs in 2010 using

Table 6.3: Process parameters base case.

Capture Recovery CoolingTin [◦C] 162 -66 -150Pin [mbar] 1100 1200 1200Flow/bed [kg/s] 91 564 357∆P packing [mbar] 16.9 82.5 56.7Total ∆P[mbar] 100 200 200

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6.2 Process evaluation 117

the Chemical Engineering Plant Cost Index (CEPCI). The packing costsare calculated using a steel price of $1200/ton steel (market price of$600/ton multiplied with factor two for packing construction). The mod-ule costs, including piping, installation etc. are then calculated by multi-plying the equipment costs with a Hand factor. When all the module costsare summed up, 25% is added for contingencies. The total direct invest-ment is subsequently calculated and an allocated investment (for storage,utilities and environmental provisions), start up investments and work-ing capital are added. Finally, the total fixed capital is calculated, resultsare shown in Table 6.4. The operational costs consist of the electricitycosts required for the blowers. The CO2 emitted due to the additionalpower required by the blowers could be captured as well, but is assumedto be emitted into atmosphere in this study. For this base case the cool-ing is provided by the evaporation of LNG and no additional costs areassumed. Depreciation, interest, labor and maintenance are calculatedusing 20% of the total capital charge per year. The used cost parameterscan be found in Table 6.5. The operational and final CO2 avoidance costsare summarized in Table 6.6. The total costs per ton of CO2 emissionavoided amounts $52.8. The capital costs ($28.9/ton CO2 avoided) andthe operational costs ($23.9/ton CO2 avoided) have a similar share in theavoidance costs.

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118

Tech

no-e

con

om

iceva

lua

tion

Figure 6.2: Simplified process scheme.

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6.2

Pro

cess

eva

lua

tion

119

Table 6.4: Capital investment costs for the base case.

Equipment Equipment costs [M$] Hand factor Module costs [M$]

Column for packed bed (21) 0.39 4 32.9Packing (21) 0.67 4 56.5Flue gas compressor 1.12 2.5 2.8CO2 recycle blower 10.09 2.5 25.2N2 cooling blower 10.55 2.5 26.4CO2 product compressor 15.09 2.5 37.7LNG heat exchanger 1.03 4.8 4.9Contingencies 25% 46.6

Total direct investment (TDI) 233Total allocated investment 40% of TDI 93Start up investment 5% of TDI 12

Total process investment (TPI) 338Working capital 2% of TPI 7

Total fixed capital 345

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120 Techno-economic evaluation

Table 6.5: Cost evaluation parameters.

Operation hours per year 7000Capital charge 0.2Blower/compressor/pump efficiencies 0.72Electricity price [$/kWh] 0.06CO2 emission due to additional power [ton/MWh] 0.8042CO2 product pressure [bar] 140

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6.2

Pro

cess

eva

lua

tion

121

Table 6.6: Operational and total costs for the base case.

CO2 captured [ton/hr] 450CO2 emitted due to additional power [ton/hr] 109.2CO2 avoided [ton/hr] 340.8

MW $/hr $/ton CO2 avoided

Flue gas blower 8.2 490 1.4CO2 recycle blower 34.4 2063 6.1N2 cooling blower 36.4 2181 6.4CO2 product compressor 56.9 3412 10.0

Total electricity blowers 135.8 8145 23.9

Capital/maintenance/labor charge 9850 28.9

Total costs 52.8

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122 Techno-economic evaluation

6.2.3 Parameter study

Initial bed temperature

The process has been evaluated for different initial bed temperatures.Both the CO2 avoidance costs and LNG consumption are shown inFig. 6.3. At a higher initial bed temperature, less CO2 is deposited perunit of bed volume. Therefore, more bed volume is required to maintainsimilar breakthrough times, resulting in increasing capital costs. Largerflow rates during the recovery and cooling step are required as well, inorder to finish in 600 seconds. A larger flow rate during the cooling cycleresults in a higher LNG consumption. An initial bed temperature of -160◦C results in even more efficient use of the beds and therefore slightlylower costs and LNG consumption. However, the temperature differencebetween LNG (-162◦C) and the refrigerated N2 becomes too small. It canbe concluded that a lower bed temperature results in more efficient CO2

capture. It should be noted that this conclusion cannot be drawn when(part) of the cooling is generated by refrigeration. The efficiency of a re-frigerator decreases with decreasing temperatures and results in highercooling costs.

CO2 concentration in flue gas

The CO2 concentration in flue gases depend on the used feedstock andprocess. A concentration of about 5 vol.% is for example encounteredin natural gas fired combined cycle power plants. The effect of theCO2 concentration on the performance of the cryogenic packed bed con-cept is summarized in Fig. 6.4. The amount of flue gas (635 kg/s) iskept constant for all cases. The front of desublimating CO2 will moveslower through the bed at decreasing inlet CO2 concentrations. There-fore smaller equipment can be used (when maintaining an equal break-through time) and consequently lower flow rates are required for therecovery and cooling steps. However, at the same time the amount ofCO2 captured will decrease, due to the lower CO2 content in the gas.The reduction in equipment size and required flows is cancelled out bythis decrease. An inlet concentration of 5 vol.% results in avoidancecosts of $95.7/ton, which are substantially higher than for the base case($52.8/ton). The increase in costs is especially strong when going to evenlower concentrations, which is related to the CO2 emissions caused by

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6.2 Process evaluation 123

-160 -150 -140 -1300

20

40

60

80

100 LNG

evaporation [ton/ton CO

2 ]

Avo

idan

ce c

osts

[$/to

n C

O2]

Initial bed temperature [°C]

Operational costs Capital costs LNG evaporation

0

1

2

3

4

Figure 6.3: Avoidance costs and LNG consumption as function of the initial bedtemperature.

the extra power required. The ratio of the additional required power tothe amount of CO2 captured becomes high at low concentrations. Fig. 6.4also shows that a CO2 inlet concentration of 15% results in lower avoid-ance costs. At even higher CO2 concentrations, recovery of the beds be-comes more difficult, as the heat stored in the first zone of the bed duringthe capture step becomes insufficient. Additional heat has to be suppliedto the process to recover CO2 in those cases.

Pressure drop

A pressure drop of 100 mbar for the capture step and 200 mbar for the re-covery and cooling step were assumed. The actual pressure drop dependson packing type, tubing diameters but possibly also to a large extent onthe gas distribution over the shallow packed beds. For a better distribu-tion a larger pressure drop is required. A non-uniform distributed feedmight result in different freezing/evaporating front velocities at differentradial positions and therefore in a non-optimal use of the bed volume.Earlier or less sharp breakthrough might be observed, resulting in alower capture rate of CO2 or higher LNG consumption. The amount of

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124 Techno-economic evaluation

0

20

40

60

80

100

120

140

1513.510

LNG

evaporation [ton/ton CO

2 ]

Avo

idan

ce c

osts

[$/to

n C

O2]

CO2 concentration [vol.%]

Operational costs Capital costs LNG evaporation

50

1

2

3

4

5

6

Figure 6.4: Avoidance costs and LNG consumption as function of the CO2 inletconcentration.

maldistribution which is still acceptable is unknown and requires morestudy. To indicate the significance of the pressure drop on the processperformance and economics, two cases with 50% higher and lower pres-sure drops were evaluated. Fig. 6.5 shows that the pressure drop overthe beds has a significant effect on the operational costs, which is ex-plained by higher compression costs. Also the amount of required LNGchanges slightly, which is related to the heat generated by compression.It should be noted that some CO2 bypass might be tolerated, since often90% capture is deemed sufficient.

6.3 Comparison with absorption and membrane techno-logy

The economics of the cryogenic packed bed concept is compared to ab-sorption and membrane technology in this section. In order to present acomparison as fair as possible, costs are calculated based on the same

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6.3 Comparison with absorption and membrane technology 125

0

20

40

60

80

100

150/300100/200

LNG

evaporation [ton/ton CO

2 ]

Avo

idan

ce c

osts

[$/to

n C

O2]

Pressure drop [mbar]

Operational costs Capital costs LNG evaporation

50/1000

1

2

3

4

Figure 6.5: Avoidance costs and LNG consumption as function of the pressuredrops during the capture step (left value) and recovery and cooling step (rightvalue).

cost parameters as used in the evaluation of the cryogenic concept (asshown in Table 6.5).

6.3.1 Absorption technology

The required input for the evaluation of CO2 capture costs via absorp-tion technology is obtained from the Integrated Environmental ControlModel developed by Rubin (2010). A 600 MW power plant with mo-noethanolamine (MEA) absorption was simulated, resulting in a flue gasof 666 kg/s containing 14 vol.% CO2 (on a dry basis), which is similarto the flue gas composition as used in the evaluation of the cryogenicconcept. The costs for all purification steps upstream the capture pro-cess (NOx, SO2 and particulates removal) are not taken into account. Theequipment costs and the electricity, steam, MEA and corrosion inhibitorconsumption are taken from the model. Steam required for strippingis generated with an auxiliary boiler in the simulation, but capital and

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126 Techno-economic evaluation

operational costs are not taken into account. The resulting values arepresented in Table 6.7 and 6.8.

6.3.2 Membrane technology

Merkel et al. (2010) carried out a basic study on the economics of CO2

capture with membrane technology, treating flue gas of 602 kg/s con-taining 12.9 vol.% CO2. In their evaluation two process alternatives wereevaluated. In the first option the driving force for permeation is generatedby a vacuum on the permeate side. In the second option an air sweepis used, which is then fed to the boiler of the power plant. Althoughthe second alternative is more efficient and looks promising, it will notbe taken into account in this study, as it will influence the combustionprocess and might be more difficult to retrofit to existing facilities. Thecalculated equipment costs only consist of the membrane costs and com-pressors/expanders, but are multiplied with an installation factor. Thecapital and operational costs adjusted with the parameters used in thisstudy are also shown in Table 6.7 and 6.8.

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6.3

Com

pa

rison

with

ab

sorp

tion

an

dm

em

bra

ne

tech

nolo

gy

127

Table 6.7: Capital investment costs for amine scrubbing and membrane technology.

Amine scrubbing Membranes

Equipment Module costs [M$] Equipment Module costs [M$]

Flue gas blower 5.7 Membranes 150CO2 absorber 81.3 Compressors/expanders 100Heat exchangers 6.2 Installation factor (60%) 150Circulation pumps 12.7Sorbent regenerator 46.7Reboiler 22.3Sorbent reclaimer/processing 15.4Drying/compression unit 49.1Contingencies 59.9 Contingencies 100

Total direct investment 299.3 Total direct investment 500Total allocated investment 119.7 Total allocated investment 200Start up investment 15.0 Start up investment 25

Total process investment 434 Total process investment 725Working capital 8.7 Working capital 15

Total fixed capital 443 Total fixed capital 740

Page 137: Novel process concept for cryogenic CO2 capture

128

Tech

no-e

con

om

iceva

lua

tion

Table 6.8: Operational and total costs for amine scrubbing and membrane technology.

Amine scrubbing Membranes

CO2 captured [ton/hr] 439 369CO2 emitted due to additional power [ton/hr] 58 120CO2 avoided [ton/hr] 381 249

$/ton CO2 avoided

Sorbent 7.0Inhibitor 1.4Reclaimer waste disposal 1.4

Total chemicals 9.9 -

Flue gas blower 2.1CO2 product compressor 9.1Solvent pump 0.2

Total power costs 11.3 36.0

Capital charge (20% total fixed capital/year) 33.2 84.9

Total costs 54.5 120.9

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6.3 Comparison with absorption and membrane technology 129

6.3.3 Comparison

The CO2 avoidance costs for all three technologies are compared inFig. 6.6. Amine scrubbing and the cryogenic concept have comparablecosts, while membranes are significantly more expensive in this evalua-tion. The results are highly dependent on the assumptions, especially onthe availability of utilities. In the amine case it was assumed that steamis available at no costs, which is unrealistic. When no steam is availableat all and the operational and capital costs and additional CO2 emis-sions related to steam generation in an auxiliary boiler are taken intoaccount, the avoidance costs for scrubbing become high ($133.4/ton).This is related to the large amount of heat required during the strip-ping of MEA (4.5 MJ/kg CO2). The model developed by Rubin also of-fers the possibility to select an advanced amine, which is used in FluorsEconamine R© FG+ process. The resulting avoidance costs are substan-tially lower ($84.2/ton), related to less steam required for regenerationand lower degradation rates.

The cryogenic concept is attractive when the cold exposed during theregasification of LNG could be used for free. If no LNG is available andthe entire required cooling capacity should be generated using cryogenicrefrigerators, the electricity consumption of the refrigerators would be inthe same order of magnitude as the electricity production of the powerplant, and can therefore be considered as unrealistic. Furthermore, theevaporation of LNG could be integrated with other processes, thereforeLNG might be only available at certain costs. When comparing to theavoidance costs of the advanced amine, a maximum price of $8.7/tonLNG can be allowed. The required cooling power is 248 MW, which cor-responds to an LNG consumption of 2.7 kg LNG/kg CO2 avoided. Anaverage sized LNG terminal evaporates about 5 million ton/year. Basedon an operation of 7000 hours per year, a total amount of 8.6 million tonof LNG would be required for cooling, which corresponds to more thanone terminal. When only one terminal is available and the remainingcooling duty has to be generated by refrigeration, the avoidance costs willbe $314.4/ton avoided, which is still excessively high. In the situationthat the pressure drops can be reduced as shown in Fig. 6.5, less coolingis required and the total avoidance costs when taking refrigeration intoaccount results in $126.5/ton, which makes it competitive with the othertechnologies.

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130 Techno-economic evaluation

0

20

40

60

80

100

120

140

160

180

Cryoge

nic

LNG + cooli

ng

Cryoge

nic

Membra

nes

Adv. a

mine

+ steam

Amine

+ steam

Avo

idan

ce c

osts

[$/to

n C

O2]

Cooling Electricity Steam Capital charge Chemicals

Amine

Figure 6.6: Avoidance costs for different technologies.

CO2 removal using membrane technology is more costly than cryo-genics and scrubbing, due to its high capital costs. When the costs ofthe membrane modules could be reduced in the future, this option maybecome competitive, especially when cold or steam utilities are not avail-able/expensive.

The cryogenic concept shows the advantage that deep CO2 removalcan be obtained, generating both a very pure cleaned flue gas as CO2

product. When cooling to -150◦C, the vapor pressure of CO2 is only 8Pa, resulting in more than 99.9% CO2 removal, compared to 90% for theother technologies. To quantify the exact advantage of this ‘deep’ CO2

removal, the costs for CO2 emissions should be known. The removal ofimpurities is not incorporated in this study. The cryogenic concept hasthe potential to remove water and for example sulfur containing impuri-ties simultaneously, as vapor pressures are low at the used temperatures.In that case it could be necessary to install a small separate bed, whichallows separate regeneration. Future work will focus on these aspects.

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6.4 Conclusions 131

6.4 Conclusions

The costs of cryogenic CO2 capture using dynamically operated packedbeds depend strongly on initial bed temperatures and CO2 concentra-tions in the feed gas. At lower initial temperatures the cold stored in thebed can be used more efficiently, resulting in more CO2 deposited perunit of bed volume. At low CO2 inlet concentrations, the relative costsfor the amount of CO2 avoided increase strongly. Due to high flow ratesrequired during the process, the pressure drops over the system sub-stantially influence the CO2 avoidance costs. It is expected that requiredgas distribution plays an important role in the resulting pressure drop.Future research should focus on the effects of gas (mal)distribution (andhence the required pressure drop over the gas distributor) on the processperformance.

In the comparison with other technologies it was found that the pre-ferred technology depends heavily on the availability of utilities. The cryo-genic concept requires a cold source, such as the evaporation of LNG ata regasification terminal, while amine scrubbing requires low pressuresteam in order to strip the solvent. When both LNG and steam are notavailable at low costs, membrane technology shows advantages. Whensteam is available at low costs, especially when using an advanced amine,scrubbing is the preferred technology. The cryogenic concept could bethe preferred option, when LNG is available at low costs. Especially whenpressure drops can be decreased and the simultaneous removal of im-purities can be incorporated in one process, the concept could become aserious candidate for capturing CO2 from flue gases.

Acknowledgment

Paul Hamers is kindly acknowledged for his excellent contribution to thework described in this chapter.

Notation

P pressure, [mbar]T temperature, [◦C]

Page 141: Novel process concept for cryogenic CO2 capture

132 Techno-economic evaluation

Abbreviations:

CEPCI Chemical Engineering Plant Cost IndexCPB Cryogenic Packed BedLNG Liquefied Natural GasMEA Mono Ethanol AmineTDI Total Direct InvestmentTPI Total Process Investment

Page 142: Novel process concept for cryogenic CO2 capture

7Biogas purification

Abstract

This chapter demonstrates with numerical simulations the option to usethe proposed process concept for biogas treatment. The performance iscompared to Vacuum Pressure Swing Adsorption (VPSA) on the basis ofseveral criteria: purity and recovery of the obtained product, bed dimen-sions and energy requirements. Simulation results reveal that the purityand recovery of CH4 are higher for the cryogenic packed bed (CPB) con-cept, while also the bed capacity is much higher: the productivity (definedas kgCH4

h-1 m-3packing) is a factor eight higher. The recovery is carried out

with air and when operated in reversed flow mode, the CPB technologyrequires a 22% lower energy duty (2.9 MJ/kgCH4

vs. 3.7 MJ/kgCH4for the

VPSA process). Furthermore, simultaneous deep H2S removal is possibleusing the proposed concept, although initial bed temperatures as low as-150◦C are required.1

1This chapter is based on a paper submitted to Industrial & Engineering Chemistry Re-search.

Page 143: Novel process concept for cryogenic CO2 capture

134 Biogas purification

7.1 Introduction

Biogas is formed at e.g. landfill sites, waste water treatment facilities oris produced by anaerobic fermentation of manure. Biogas mainly con-sists of CH4 (50-70 vol.%) and CO2 (25-45 vol.%) and furthermore maycontain contaminants such as H2O, H2S and siloxanes (Patterson et al.,2011). CO2 is a well known greenhouse gas, however, CH4 has a relativeglobal warming potential 25 times higher than CO2 (Sejian et al., 2011).It is reported that greenhouse gas emissions from the agricultural sec-tor account for about 25.5% of the total global anthropogenic emissions(Sejian et al., 2011). It is therefore critical to bring down CH4 emissions.The cheapest option to avoid emissions is to collect biogas and send itto a flare. However, biogas based methane has the potential to serveas a renewable energy resource, to generate power or to be used as atransportation fuel. To convert biogas to commercial grade CH4, severalseparation and purification steps are required.

H2S is formed by the anaerobic fermentation of sulfur containing pro-teins. H2S removal is necessary, as combustion will lead to the envi-ronmentally hazardous SO2 and could form H2SO4, causing corrosionof process equipment. H2S can be removed by oxidation to elementarysulfur or by scrubbing the biogas with an aqueous alkaline solution. Dis-advantage of the latter is that CO2 has a higher reactivity with the solvent,causing low selectivities for H2S removal (Abatzoglou and Boivin, 2009).

Siloxanes are a group of molecules containing Si, which are found inlandfill biogases. The combustion of siloxanes leads to silicates and microcrystalline quartz. These solids will damage engines and turbines, andare therefore highly undesirable. Several possible technologies to removesiloxanes are available: (reactive) absorption with liquids, adsorption andcryogenics (Abatzoglou and Boivin, 2009).

CO2 is present in high concentrations in biogas. Removal of CO2 isnecessary to increase the energy content of the biogas. The availabletechnologies are: scrubbing (using water, a physical or chemical sol-vent), cryogenic separation, membranes or Pressure Swing Adsorption(PSA) (Basu et al., 2010; Patterson et al., 2011). Similar technologies areapplied or studied for CO2 removal from flue gases. However, differenceswith flue gas treatment are the higher CO2 content in the feed, the lowertemperature at which the gas is available and the product requirements.The desired product in flue gas treatment is CO2, while biogas treat-

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7.2 Adsorption 135

ment is focused on obtaining CH4. The removed CO2 is normally notconsidered for sequestration, as the amounts of CO2 produced are lower,making CO2 collection and transportation relatively expensive.

All technologies listed above have been commercially applied to re-move CO2 from biogas, but energy requirements are normally high dueto required pressure differences or elevated temperatures for solvent re-covery. This chapter describes the possibilities of using the cryogeniccapture technology developed in this thesis and compares it to adsorp-tion technology on the basis of product purity and recovery, bed dimen-sions and energy requirements. To be able to compare the performanceand requirements of the two processes, a specific detailed study on ad-sorption from literature has been selected (Grande and Rodrigues, 2007),in which both dimensions and properties of the beds as well as energyrequirements have been provided. First, some details about the selectedadsorption process are given. Subsequently, the possibility of using theCPB concept is explored by carrying out simulations using the numericalmodel as described in Chapter 2. Finally, the two processes are com-pared for the removal of CO2 from a CH4/CO2 mixture. Finally, additionalsimulations are presented which show the possibility of simultaneouslyremoving H2S using the proposed process concept.

7.2 Adsorption

Adsorption technology is a widely applied separation technology. Sep-aration of CO2 from gas mixtures by adsorption is based on differ-ences in equilibrium capacities at the adsorbent surface (e.g. zeolite13X) or on differences in uptake rates (e.g. carbon molecular sieve3K). Regeneration of a bed is normally obtained by reducing the pres-sure. An adsorption process therefore consists of several packed bedswhich are operated simultaneously in different process steps, e.g. feed,recovery and pressurization. The removal of CO2 from biogas us-ing adsorption is widely discussed in literature (Grande and Rodrigues,2007; Kapoor and Yang, 1989; Kim et al., 2006; Esteves et al., 2008;Ribeiro et al., 2008). Grande and Rodrigues (2007) analyzed VacuumPressure Swing Adsorption (VPSA), in which the recovery step is carriedout at sub-atmospheric pressure. The purity and recovery of CH4 andenergy consumption by changing pressures and cycle times have beeninvestigated for two different adsorbents, Carbon Molecular Sieve (CMS)

Page 145: Novel process concept for cryogenic CO2 capture

136 Biogas purification

Table 7.1: Bed properties

Property CPB VPSALength [m] 1.65 4.667Radius [m] 0.3 0.4667Porosity [-] 0.7 0.33Packing material SS CMSType of packing monolith particlesBulk density [kg m-3] 2347 715.4Number of beds [-] 2 2

Table 7.2: Operating conditions

Property CPB VPSAFeed flow rate [SLPM] 16000 16000Feed pressure [bara] 5 8Recovery pressure [bara] 1.1 0.1Initial bed temperature [◦C] -110 20Cycle time [s] 140 140Purity [CH4%] 99.1 98.1Recovery [CH4%] 94.3 79.7Productivity [kg CH4 h-1 m-3

packing] 350.2 43.1

3K and zeolite 13X. Lowest energy requirements have been obtained in arun with CMS as adsorbent. In this run a gas mixture of 16 000 SLPM(0.312 kg/sec), containing 45 vol.% CO2 and 55 vol.% CH4 is treated. Itis assumed that the biogas is available at 2 bara and 25◦C. Furthermoreit is assumed that water and other contaminants have been previouslyremoved. The bed properties and operating conditions are listed in Ta-ble 7.1 and 7.2 respectively. For further details the reader is referred totheir work.

7.3 Cryogenic packed bed concept

Although the process is originally developed to capture CO2 from fluegases, it can also be applied to a wide range of other gas separations, suchas biogas treatment. In case of CO2 capture from flue gases, gaseous CO2

Page 146: Novel process concept for cryogenic CO2 capture

7.3 Cryogenic packed bed concept 137

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6-120

-100

-80

-60

-40

-20

0

20

40 0 s 70 s 140 s

Tem

pera

ture

[°C

]

Axial position [m]

(a)

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6

0

20

40

60

80

100 70 s 140 s

Mas

s dep

ositi

on [k

g·m

-3]

Axial position [m]

(b)

Figure 7.1: Simulated axial temperature (a) and mass deposition (b) profiles dur-ing the capture step.

is fed to a bed during the recovery step in order to obtain pure CO2 at theoutlet. For biogas treatment, CO2 capture is not required, therefore aircan be used in the recovery step. The evolution of axial temperature,mass deposition and concentration profiles within the packed beds canbe described using the pseudo-homogeneous one dimensional axially dis-persed plug flow model described in Chapter 2.

In order to compare the CPB concept with VPSA, simulations havebeen carried out for the cryogenic concept using an equal flow rate andgas composition as used by Grande and Rodrigues (2007). The bed iscooled down to -110◦C and the feed is pressurized to 5 bar. Under theseconditions the outlet CH4 purity is 99.1%. The dimensions of the beds(see Table 7.1) are chosen in such a way, that the capture/feed step takes140 s, similar to the VPSA case. The recovery and cooling step togetheralso take 140 s, therefore the process can be operated continuously whentwo beds are operated in paralel. Axial temperature and mass depositionprofiles during the capture step are shown in Fig. 7.1a and 7.1b respec-tively. It should be noted that the process cycle for the VPSA case consistsof a (1) feed (60 s), (2) depressurization (10 s), (3) blow down (140 s), (4)purge (50 sec) and (5) pressurization step (70 s). Although their analysisis based on two beds, three beds are required if continuous operation isdesired.

Page 147: Novel process concept for cryogenic CO2 capture

138 Biogas purification

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6-120

-100

-80

-60

-40

-20

0

20

40 0 s 3 s 6.5 s

Tem

pera

ture

[°C

]

Axial position [m]

(a)

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6

0

20

40

60

80

100 0 s 3 s

Mas

s dep

ositi

on [k

g·m

-3]

Axial position [m]

(b)

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6

-120

-110

-100

-90

-80 30 s 70 s 140 s

Tem

pera

ture

[°C

]

Axial position [m]

(c)

Figure 7.2: Simulated axial temperature (a) and mass deposition (b) profiles dur-ing the recovery step in the first seconds and the temperature profiles later intime (c).

Page 148: Novel process concept for cryogenic CO2 capture

7.3 Cryogenic packed bed concept 139

When CO2 breakthrough is observed, the bed should be recovered.By reducing the pressure to atmospheric pressures and by flushing thesystem with (dry) air, very fast recovery can be obtained. Therefore therecovery and cooling step can be integrated in one step by using an airflow of 5 kg/s, refrigerated to -110◦C. All CO2 is recovered after 6.5seconds, as can be observed from Fig. 7.2a and 7.2b. In this first 6.5seconds, the outlet flow cannot be recycled to the inlet of the bed viathe cooler, due to the presence of CO2. Therefore an amount of 32.5 kgof dry air is required per step (average flow of 0.23 kg/s). After all CO2

is recovered, the entire bed is further cooled down to -110◦C, in orderto start a new capture step again (Fig. 7.2c). A simplified process flowdiagram including flow compositions and process conditions is shown inFig. 7.3.

Page 149: Novel process concept for cryogenic CO2 capture

140

Bio

ga

sp

urifi

ca

tion

Figure 7.3: Simplified process scheme including conditions for the base case.

Page 150: Novel process concept for cryogenic CO2 capture

7.3

Cry

ogen

icp

ack

ed

bed

con

cep

t141

Figure 7.4: Simplified process scheme including conditions, recovery/cooling step in reversed flow mode.

Page 151: Novel process concept for cryogenic CO2 capture

142 Biogas purification

7.4 Adsorption versus cryogenic packed bed concept

The two processes are compared on:

• Product purity and recovery of CH4 from the biogas

• Required bed dimensions

• Energy requirements

Table 7.2 shows that the product purity in VPSA is 98.1% CH4 against99.1% in the CPB concept. The recovery of CH4 in VPSA is lower (79.4%),due to losses in purge steps. The losses of CH4 in the CPB concept arelimited to the amount of CH4 present in the bed before switching to therecovery. When reducing the pressure and flushing the system with airthis amount of CH4 is lost: the recovery is 94.3%. The beds required inthe cryogenic concept are much smaller (Table 7.1). The productivity iscalculated for both processes: this value is more than 8 times higher forthe CPB concept (350.2 vs. 43.1 kgCH4

h-1 m-3packing).

Grande and Rodrigues (2007) estimated the energy requirements forthe adsorption process, which are summarized in Table 7.3. The totalpower required amounted 3.7 MJ/kg CH4. The energy requirements forthe CPB concept are based on the process flow scheme shown in Fig. 7.3.Three compressors are required, the first one to pressurize the feed to5 bar (C-1), the second one to pressurize the product to 200 bar (C-2) and a third one to recycle air through the bed and the cooler (C-3).Simulations showed that the pressure drop in the packed bed is only0.04 bar, which is explained by the selected nature of the structuredpacking (monolith). It is assumed that the cooler causes another pressuredrop of 0.06 bar, therefore an outlet pressure of 1.1 bar is required forcompressor C-3. All compressor duties have been calculated assumingisentropic compression and an efficiency of 72%. Results are shown inTable 7.4.

Furthermore, energy is required by the refrigerator in order to cooldown the air. As described earlier, an average amount of 0.23 kg/s offresh dry air is required to recover the bed. It is assumed that this airis available at 25◦C. After CO2 is removed from the bed, the refrigeratedair can be recycled. The temperature of the air at the outlet of the bedis changing during the recovery/cooling step, as shown in Fig. 7.5. Theaverage outlet temperature is -101.7◦C, which is mixed with fresh air

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7.4 Adsorption versus cryogenic packed bed concept 143

Table 7.3: Energy requirements for VPSA

Step Power (kW)Compression of feed to 8 bar 57.7Compression of product to 200 bar 73.2Power for purge step 28.1Blowdown step 129.8Total power 288.8

of 25◦C, resulting in an average inlet air temperature of -95.8◦C. Dueto compression in C-3, the stream heats up and will be cooled downin the refrigerator to -110◦C. The required cooling duty is 102.4 kW.However, the efficiency of a cooler in this power and temperature rangeis approximately 40% Timmerhaus and Flynn (1989), and therefore thepower input for the refrigerator is higher. The resulting cooling costs arealso listed in Table 7.4. The total energy requirement for the processthen amounts to 4.3 MJ/kg CH4, which is 16% higher than for the VPSAprocess.

The energy requirements are mainly caused by the cooling step. A rel-atively large air flow of 5 kg/s is required to cool down the bed to -110◦C.When taking a closer look at the temperature profiles in Fig. 7.2, it canbe observed that the hot zone at the inlet of the bed (which is at the bio-gas inlet temperature, 25◦C), is moved through the entire bed. This isalso visible in the outlet temperature as function of the time in Fig. 7.5.Therefore, it is more efficient to reverse the flow direction during the re-covery/cooling step, as illustrated in the process flow scheme in Fig. 7.4.Fig. 7.6 shows the resulting axial temperature and mass deposition pro-files. In this case a lower total air flow is required (3 kg/s). All CO2 isrecovered after 10 seconds, therefore an average amount of 0.21 kg/s offresh dry air is required. The resulting outlet temperature as function ofthe time is also plotted in Fig. 7.5. When comparing the two profiles forthe base case and the reversed flow mode, it can be concluded that theaverage outlet temperature is lower in the reversed flow mode. Further-more, the waste stream containing CO2 is emitted at a relatively highertemperature, avoiding cold losses. Finally, the lower flow rate causes alower power consumption by compressor C-3. The required energy for allsteps are also included in Table 7.4. A total amount of 2.9 MJ/kgCH4

isrequired, which is 22% lower than for the VPSA process.

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144 Biogas purification

0 20 40 60 80 100 120 140

-120

-100

-80

-60

-40

-20

0

20 Base case Reverse recovery

Out

let t

empe

ratu

re [°

C]

Time [s]

Figure 7.5: Outlet temperature as function of time

Table 7.4: Energy requirements for cryogenic packed beds

Power (kW)Step Base case Reversed flowCompression of feed to 5 bar 37.9 37.9Compression of product to 200 bar 62.6 62.6Air recycling 34.2 20.4Cooling power 256 142.5Total power 390.7 263.4

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7.4 Adsorption versus cryogenic packed bed concept 145

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6-120

-100

-80

-60

-40

-20

0

20

40 0 s 5 s 10 s

Tem

pera

ture

[°C

]

Axial position [m]

(a)

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6

0

20

40

60

80

100 0 s 5 s

Mas

s dep

ositi

on [k

g·m

-3]

Axial position [m]

(b)

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6

-120

-110

-100

-90

-80 30 s 70 s 140 s

Tem

pera

ture

[°C

]

Axial position [m]

(c)

Figure 7.6: Simulated axial temperature (a) and mass deposition (b) profiles dur-ing the recovery step in the first seconds and the temperature profiles later intime (c). The bed is operated in reversed flow mode.

Page 155: Novel process concept for cryogenic CO2 capture

146 Biogas purification

7.5 Hydrogen sulfide removal

The H2S content in biogas can range from 0.0001 - 1 vol.%(Abatzoglou and Boivin, 2009). When feeding a mixture containing1 vol.% H2S, 45 vol.% CO2 and 54 vol. % CH4 to a cryogenically refriger-ated bed (equal conditions and properties as those listed in Table 7.1 and7.2), CO2 and H2S can be separated simultaneously. However, with a bedtemperature of -110◦C the H2S content can only be reduced to 0.6 vol.%.If deep H2S removal is required, the initial bed temperature should belower. The H2S content can be reduced to 40 ppmv when the initial bedtemperature is -150◦C. Fig. 7.7 shows the resulting axial temperatureand mass deposition profiles when feeding the mixture to a bed initiallycooled to -150◦C.

It can be observed, that both CO2 as well as H2S will deposit ontothe packing surface, although the zone where H2S deposits also containsCO2. This can be explained by having a closer look at the saturation va-por pressures of both components in Fig. 7.8. When the mixture is cooleddown, CO2 starts to deposit at -88◦C, as indicated by the dotted line. Thetemperature will decrease further and below -105.1◦C the saturation va-por pressure of H2S becomes lower than its partial pressure in the feed(0.05 bar) and H2S starts to deposit as well. During the cooling of themixture from -105.1◦C to the initial bed temperature, both componentswill deposit onto the packing. Therefore, it can be concluded that compo-nents with saturation vapor pressures close together will deposit in thesame region. Nevertheless, these results show that both components canbe efficiently removed from the feed gas.

7.6 Discussion and conclusions

In this chapter it is proposed to purify biogas using the process conceptstudied in this thesis. The possibility has been studied with numericalsimulations. The cryogenic packed bed concept shows to be favorablecompared to adsorption technology on several aspects. The purity andthe recovery of the obtained CH4 are higher than in VPSA. Higher purityis possible (up to 99.99%) at lower initial bed temperatures. Furthermore,required bed sizes are significantly smaller, resulting in much lower cap-ital investments. However, good insulation of the cryogenic packed bedsis required to avoid heat leaks into the system from the surroundings.

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7.6 Discussion and conclusions 147

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6-160-140-120-100-80-60-40-200

2040

Tem

pera

ture

[°C

]

Axial position [m]

(a)

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.60

20

40

60

80

100

120

CO2 + H

2S

CO2

H2S

Mas

s dep

ositi

on [k

g·m

-3]

Axial position [m]

CO2

(b)

Figure 7.7: Simulated axial temperature and mass deposition profiles after 140seconds when feeding a CH4/H2S/CO2 mixture to an initially refrigerated packedbed.

-160 -140 -120 -100 -80 -60 -4010-2

10-1

100

101

102

103

104

105

106

107

108

109

Initial bed T

Feed pressure H2S

Vapor pressure CO2

Vapor pressure H2S

Vap

or p

ress

ure

[Pa]

Temperature [°C]

Feed pressure CO2

Figure 7.8: Saturation vapor pressures of CO2 and H2S as function of the tem-perature.

Page 157: Novel process concept for cryogenic CO2 capture

148 Biogas purification

Additionally, energy requirements are quite competitive with VPSA, evenif a cooler efficiency of 40% is taken into account. Energy requirementsare especially low when the recovery/cooling mode is operated in reversedflow mode. Furthermore, dried air (or nitrogen) is required, which mayraise operating costs somewhat. The amount of dried air can be reducedby operating the recovery/cooling step at sub-atmospheric pressures, atthe expense of higher operation costs. Furthermore, it should be notedthat the upgraded gas is compressed to 200 bar in this study. However,if the cleaned product would be liquefied, the CPB technology shows theadvantage that the product gas leaving the packed beds is already at alow temperature of -105◦C, reducing further cooling costs. This studyfocused on the CH4/CO2 separation and showed the possibility to simul-taneously remove H2S. Results in Chapter 3 showed that H2O can beseparated simultaneously using the CPB concept. Furthermore, silox-anes could also be separated in the same or a separate bed. Therefore, itcan be concluded that the proposed CPB technology is a very promisingprocess for biogas treatment.

Notation

P pressure, [bar]T temperature, [◦C]

Abbreviations

CMS Carbon Molecular SieveCPB Cryogenic Packed BedPSA Pressure Swing AdsorptionVPSA Vacuum Pressure Swing Adsorption

Page 158: Novel process concept for cryogenic CO2 capture

8Epilogue and outlook

This dissertation discusses a novel process concept for cryogenic CO2

capture based on dynamically operated packed beds. When feeding a fluegas to a refrigerated packed bed, a separation between CO2 and N2 canbe obtained. Two models to describe this process have been developed inthis work, a detailed one-dimensional pseudo-homogeneous model tak-ing axial dispersion effects into account and a simplified, but fast ‘sharpfront’ model. These models have been validated using experiments in asmall scale experimental setup. This setup was used to demonstrate thecapture step for N2/CO2/H2O gas mixtures, showing simultaneous sep-aration of CO2 and H2O. The entire process, including the recovery andcooling step, was demonstrated in a continuous fully automated experi-mental pilot setup. An unknown parameter required for the design wasthe desublimation rate of CO2. Therefore, a dedicated setup was designedto measure these desublimation rates under a wide range of conditions.A model to describe frost growth rates was successfully developed. It wasdemonstrated that mass transfer towards the cold packing is the rate de-termining step under the conditions in the packed beds. The economicsof the novel concept were studied and compared with competing tech-nologies. Especially when a cold source, such as the regasification ofLNG is available at low costs, the process could compete well with other

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post-combustion CO2 capture technologies. Finally, it was demonstratedthat the concept could be very promising for biogas purification, evenwhen cooling needs to be provided by refrigeration. By using a simulta-neous temperature and pressure swing after the capture step, depositedCO2 was very efficiently removed from the bed.

It can be concluded that already many important aspects of the pro-cess concept have been discussed in this dissertation. It is attempted inthis chapter to give a realistic future outlook for the concept and to listsome important points of attention for future research activities.

8.1 Important aspects for future development

This section discusses some important aspects of the proposed concept.First the effects of the recovery step on the process performance are dis-cussed, followed by a discussion on the possible consequences of radialprofiles within the beds.

The cooling duty required to capture a certain amount of CO2 dependson aspects such as the CO2 concentration in the feed gas and the initialbed temperature before a capture step is started. However, it is also indi-rectly related to the operation mode of the recovery step, as the recoverymode dictates the temperature of the bed before a cooling step is started.During the recovery step, a driving force is required for the sublimation ofthe solid CO2 deposited during the capture step. This driving force couldbe in the form of heat, which was proposed for flue gas treatment. PureCO2 is fed to the system and the heat stored in the initial zone of the bedduring the capture step is used to heat up the rest of the bed and evapo-rate CO2. The result of this way of operation is that the bed temperaturewill be at least -78◦C before starting the cooling step. In chapter 7 it wasshown for biogas treatment that it is possible to provide the driving forcefor sublimation in another way: by reducing the pressure. This reductionis in the first place attained by letting the total pressure decrease, butat the same time feeding a CO2 lean gas (such as air) to the bed. Dueto the large driving force, CO2 will evaporate and is removed very fast bythe air flow. A coinciding advantage of this way of recovery is that thesublimation heat is returned to the packing material. Therefore, coolingof the bed to the required initial temperature is for a large extent coveredby the evaporation of CO2, causing the cooling costs to be minimal. Thisway of recovery is difficult for flue gas treatment. In the first place be-

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cause the capture step cannot be performed at elevated pressures, dueto the large amounts of flue gas and related high compression costs. An-other difference with biogas treatment is that in flue gas treatment thecaptured CO2 is the product, while for biogas treatment the gas passingthrough the bed (CH4) is the product. When CO2 capture is not requiredfor biogas treatment air can be used to recover the bed.

The recovery step is not only having effect on the required coolingduty, but is also for a large extent responsible for the electricity costsdue to the recycle blower. The gas flow required during the recoverystep is five to six times as large as the flue gas flow. Especially for powerplant scale, these flows will be large and even a small increase in pressuredrop will have a substantial effect on the process economics, as explainedin Chapter 6. It could be very advantageous to develop an alternativerecovery mode in future work. An alternative option was discussed inChapter 2, in which the deposited CO2 was molten to the liquid phase byblocking a vessel after the capture step and introducing heat by internalheat pipes. It is expected that increased capital costs will not balance theenergy savings. However, it can be recommended to carry out an accurateeconomic evaluation of this concept.

Another alternative could be to operate the process using moving beds,in which the packing materials will be circulated. After CO2 has beendeposited onto the packing, it could be removed from the bed and trans-ported to a separate vessel where CO2 could be removed (possibly underpressure) to finally transport the particles to a cooling step. In this waya continuously operated process could be obtained, at the expense ofoperational complications related to the handling of moving solids.

Radial temperature profiles were observed during experiments withthe pilot setup, as described in Chapter 5. These radial profiles could beexplained by the heat transfer in and to the vessel wall, an effect which isexpected to play a minor role for large scale beds, where the volume of thevessel wall is small compared to the bed volume. However, it could verywell be that on large scale operation still radial profiles may be impor-tant in the beds. To minimize pressure drops over the packing, shallowpacked beds were chosen in the economic analysis in Chapter 6, havinga diameter larger than the length of the beds. Gas distribution becomesan important aspect for such kind of beds. A perfectly distributed flowwill result in a possibly unacceptably high pressure drop. It might verywell be that a certain amount of maldistribution of the feed has to betolerated in order to minimize the pressure drop. The effects on the front

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development is unknown. It could be that due to flow maldistribution,not all parts of the bed will have the same temperature after a coolingstep, before starting the capture step. The consequence could be that thefrost fronts will move faster through the zone in which the temperature islower, and early breakthrough of CO2 might be the result. That will causean inefficient use of the cold stored in the packing, as the bed should beswitched to a recovery step as soon as breakthrough is observed. How-ever, it could also be that heat transfer within the bed will level out thetemperature differences. Another option is that the zones in which moreCO2 is deposited will cause an additional pressure drop, and gas will beautomatically redirected to zones in which no CO2 is deposited yet. Toget a better insight in the possible consequences, the model should beextended to two dimensions and the Navier-Stokes equations should beincorporated to calculate the velocity profiles within the bed.

8.2 Future of the proposed concept

Scaling up a new technology such as discussed in this thesis directly fromlaboratory scale to refinery or power plant scale is not realistic. Thereare still several aspects to be studied in more detail in future work, asdiscussed in the previous section. Biogas purification is identified asa very interesting application of the developed technology. The smallerscale compared to flue gas treatment and therefore the lower risks, couldmake biogas treatment an ideal intermediate scale up step before goingto flue gas scale.

It is concluded in the techno-economic evaluation in Chapter 6 thatfor flue gas treatment the concept could be economically viable when aLNG regassification terminal is available to provide the required coolingduty at low cost. If this is not the case, refrigeration is required, whichwill become expensive and other capture technologies might be preferred.Therefore, the technology presented in this dissertation is not providingthe single solution to bring down greenhouse gas emissions, but could of-fer a promising option in some situations. At this stage no capture tech-nology is clearly outperforming competing technologies. The preferredtechnology will likely depend on the local conditions, such as the LNGavailability for the cryogenic concept or low steam costs for scrubbing.Therefore it is necessary to progress research on all capture technologies,including cryogenic CO2 capture. Due to the very high degree of CO2 re-

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moval possible with the cryogenic process, an interesting option could beto used the cryogenic capture process as a second stage for deep CO2

removal after a first capture stage based on scrubbing to remove about90%.

However, the progress in CCS (Carbon Capture and Storage) devel-opment is taking place at a slower pace than required. This is mainlyrelated to the limited public acceptance of CCS, which is again relatedto the uncertainties in safety risks or at least the inability to judge thesignificance of these risks. The successfulness of CCS may therefore bemuch more dependent on social and political aspects, in stead of techno-logy developments.

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List of publications

Journal papers

1. M.J. Tuinier, M. van Sint Annaland, G.J. Kramer, and J.A.M.Kuipers. Cryogenic CO2 capture using dynamically operated packedbeds. Chemical Engineering Science, 65(1):114-119, 2010.

2. M.J. Tuinier, M. van Sint Annaland, and J.A.M. Kuipers. A novelprocess for cryogenic CO2 capture using dynamically operatedpacked beds - An experimental and numerical study. International

Journal of Greenhouse Gas Control, 5(4):694-701, 2011.

3. M.J. Tuinier, H.P. Hamers, and M. van Sint Annaland. Techno-economic evaluation of cryogenic CO2 capture - A comparison withabsorption and membrane technology. International Journal of

Greenhouse Gas Control, In press, doi:10.1016/j.ijggc.2011.08.013,2011.

4. M.J. Tuinier, and M. van Sint Annaland. Biogas Purification usingCryogenic Packed Beds. Submitted to Industrial and EngineeringChemistry Research, 2011.

5. M.J. Tuinier, and M. van Sint Annaland. Desublimation rates ofCO2 onto cryogenically cooled surfaces. Prepared for submission toInternational Journal of Heat and Mass Transfer.

6. M.J. Tuinier, and M. van Sint Annaland. Experimental proof-of-concept for cryogenic CO2 capture using dynamically operatedpacked beds. To be prepared for submission.

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7. M.J. Tuinier, and M. van Sint Annaland. A short-cut model to evalu-ate the conceptual feasibility for the cryogenic separation of contam-inants from a gaseous stream using dynamically operated packedbeds. To be prepared for submission.

Patents

1. G.J.B. Assink, G.J. Kramer, M. van Sint Annaland, and M.J. Tuinier.Process for the separation of CO2 from a gaseous feed stream, patentWO/2009/047341.

2. M. van Sint Annaland, and M.J. Tuinier. Process for the separationof contaminant or mixture of contaminants from a CH4 comprisinggaseous feed stream. European patent application No. 10188933.5,filed 26th of October 2010 (unpublished).

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Curriculum Vitae

Martin Tuinier was born on the 10th of August in Wijhe, The Nether-lands. He grew up in Deventer, where he attended elementary school andobtained his pre-university diploma in 2000 from the Alexander HegiusLyceum. In September 2000 he started studying Chemical Engineeringat the University of Twente in Enschede, The Netherlands. He special-ized in Process Engineering and carried out an internship for Akzo NobelPolymer Chemicals in Ningbo, China. During this internship, he workedon the optimization of a crystallization section of an organic peroxide pro-duction plant. He graduated in May 2007 in the group ‘Fundamentals ofChemical Reaction Engineering’ on the ‘Kinetic aspects of Oxidative Cou-pling of Methane on a Mn/Na2WO4/SiO2 Catalyst’. After obtaining hisMaster degree, he started a PhD project in the same group, under thesupervision of prof.dr.ir. Martin van Sint Annaland. In September 2010,the project was continued at the Technical University of Eindhoven, TheNetherlands. The results of this project are described in this dissertation.

In September 2011 Martin started working as a process engineer inthe Chemical Reaction Technology group of Evonik Industries A.G. inMarl, Germany.

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Dankwoord

Promoveren of de industrie in. Een lastige keuze, die niet alleen ik, maarmeerdere collega’s hebben moeten maken aan het einde van hun studie.Eindelijk je kennis toepassen in het bedrijfsleven, of je toch nog vier jaarlang verdiepen in een bepaald onderwerp? De keuze werd in mijn gevalmakkelijker (of was het nou juist moeilijker?) gemaakt door de interes-sante opdracht die me werd aangeboden door Martin van Sint Annaland:het ontwikkelen van een nieuw proces concept voor het cryogeen afvan-gen van CO2. Hier had ik wel oren naar en dacht: ik grijp deze kans.En zonder enige spijt. Met het schrijven van dit dankwoord sluit ik eenleerzame, nuttige en ook leuke periode af. Dat heb ik aan vele mensen tedanken, waarvan ik er een aantal in het bijzonder wil bedanken.

Martin van Sint Annaland wil ik in de eerste plaats hartelijk bedankenvoor zijn vertrouwen in mij en het aanbieden van deze promotie opdracht.Zijn voorverkennende werk bij Shell was erg waardevol, waardoor we alsnel met de eerste resultaten naar buiten konden komen. Tijdens hetpromotie traject heeft Martin me subliem geholpen met deskundig adviesen als het nodig was met motiverende woorden. Martin, heel erg be-dankt! Daarnaast wil ik Hans Kuipers bedanken, die vooral initieel nauwbetrokken was bij het project. Het bewaken van de grote lijn was eenbelangrijke bijdrage.

Vanuit Shell is Gert Jan Kramer betrokken geweest bij het project. Ikben erg dankbaar voor zijn input tijdens de project besprekingen, welkezonder uitzondering nuttig en inspirerend waren. Ook Jonathan Barsemaen Marijke Hogenbirk van Shell wil ik hartelijk danken voor het beoorde-len van ons werk op eventueel patenteerbare resultaten en het vervolgensvrijgeven ervan.

Experimenteel werk besloeg een aanzienlijk deel van mijn promotie-werk. De daarom erg belangrijke technische ondersteuning van Johan

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168 Dankwoord

Agterhorst en Robert Meijer was van hoge kwaliteit. Ik vond dat weeen sterk team vormden en waardeer het zeer dat jullie ook tijdens deonzekere periode rond de verhuizing van Twente naar Eindhoven hardzijn blijven doorwerken aan mijn opstellingen. Ook ben ik erg dankbaarvoor de hulp van Erik Analbers, Wim Leppink en Gerrit Schorfhaar bijhet snel oplossen van problemen met de opstellingen.

Ik heb het geluk gehad een aantal sterke studenten te hebben mo-gen begeleiden, die een goede bijdrage hebben geleverd aan de resultatenbeschreven in dit proefschrift. Nhi Dang, thank you very much for thehelp during the design of the kinetic setup and carrying out the firstexploring experiments. Niels Hietberg ben ik dankbaar voor het verderoptimaliseren van deze opstelling en voor het uitvoeren van de vele ex-perimenten. Jeroen Zijp heeft een wezenlijke bijdrage geleverd aan hetontwerp van de pilot plant: Jeroen, bedankt! Paul Hamers heeft erg goedwerk verricht in het kader van de techno-economische evaluatie, waar ikhem erg dankbaar voor ben.

De secretariele ondersteuning van Nicole Haitjema in Twente en Ju-dith Wagters in Eindhoven was altijd goed geregeld, bedankt daarvoor!Ik wil ook graag Quint Segers en Paul Hamers alvast bedanken voor hetbijstaan tijdens de verdediging als mijn paranimfen. Tenslotte, wil ikalle collega’s van FCRE/SMR bedanken voor de goede samenwerking engezellige pauzes, borrels en uitjes!

-Martin Tuinier