Chemical Engineering Science 62 (2007) 71967204www.elsevier.com/locate/ces

Amulti-scale approach for CFD calculations of gasliquid owwithin largesize column equippedwith structured packing

L. Raynal, A. Royon-LebeaudIFP, BP 3, 69390 Vernaison, France

Received 19 April 2007; received in revised form 16 July 2007; accepted 12 August 2007Available online 19 August 2007

Abstract

This work has been carried out in the framework of post-combustion CO2 capture process development. Considering the huge amount ofgases to be treated and the constraints in terms of pressure drop, it appears that the absorption column will be equipped with high efciencyhigh capacity packings such as structured packings. The present paper focuses on the CFD modellisation of the two-phase ow within thiscomplex geometry. For limited computational resources reasons, it is presently impossible to run computations at large scales taking intoaccount the gasliquid interaction and the real geometry of the packing and original approaches must be developed. In the present work, amulti-scale approach is proposed. It rst considers liquidwall and liquidgas interaction at small scale via two-phase ow calculations usingthe VOF method. Second, the latter results are used in three-dimensional calculations run at a meso-scale corresponding to a periodic elementrepresentative of the real packing geometry. Last, those results are further used at large scale in three-dimensional calculations with a geometrycorresponding to a complete column. Results are compared with experimental data and with other CFD simulations in terms of liquid hold-up,pressure drop and unit operation. Some suggestions are made for further development. 2007 Elsevier Ltd. All rights reserved.

Keywords: Structured packing; CFD; Two-phase ow; Packed column; CO2 capture; Gas treatment

1. Introduction

This work has been carried out in the framework of post-combustion CO2 capture process development. It focuses onthe capture of CO2 emitted by power plants using gas, fuel orcoal. This process consists in using an amine based solvent,typically monoethanolamine (MEA), which removes CO2 frompost-combustion emitted gas within a packed tower, the ab-sorber, and in regenerating the solvent in a second column, thedesorber, by heat regeneration (see e.g. Freguia and Rochelle,2003). Post-combustion CO2 capture process is mainly char-acterised by three aspects. First, it deals with huge volume ofgas to be treated. For a typical 400MW power station, the cor-responding gas ux is approximately 1.106 Nm3 h1. Second,it is characterised by low CO2 partial pressure, in the rangeof 0.040.13 bar which is much less than the values of up to20 bar observed in natural gas treatment. This implies relatively

Corresponding author. Tel.: +33 4 78 02 25 27; fax: +33 4 78 02 20 08.E-mail address: ludovic.raynal@ifp.fr (L. Raynal).

0009-2509/$ - see front matter 2007 Elsevier Ltd. All rights reserved.doi:10.1016/j.ces.2007.08.010

low liquid load and large gas supercial velocity. Last, sincethis process is located just downstream the power plant operat-ing at a pressure close to the atmospheric pressure, the captureprocess must generate very low pressure drop. Since the maxi-mum pressure drop across the CO2 absorber would be around50100mbar for packings bed height of almost 30m, typicalvalues of admitted pressure drop per unit length would be inthe range of 1.5.3mbar m1. These three aspects call for highcapacity, high efciency and low pressure drop gasliquid con-tacting internals. Due to their geometric characteristics, highspecic geometric area and high void fraction, structured pack-ings are good candidates to meet these three requirements.

Since the work carried out in the group of Pr. Fair atUniversity of Austin, many studies have been devoted tostructured packings in order to develop models for pressuredrop and liquid hold-up, the latter being further used in masstransfer models (Bravo et al., 1985; Fair and Bravo, 1990).Until recently, most of the experimental studies on structuredpackings have been carried out for distillation. In most cases,uids are water or light hydrocarbons and less work has been

L. Raynal, A. Royon-Lebeaud / Chemical Engineering Science 62 (2007) 71967204 7197

carried out with liquids more viscous than water. Since globalwarming and greenhouse gas emission reduction have becamea major worldwide issue, recent works have been focused onamine based solvent for CO2 capture (see e.g. Aroonwilas et al.,1999; Tobiesen and Svendsen, 2006). Such solvents have vis-cosities which range from 2 up to 10 cP and little work hasbeen done on liquid viscosity effects for such viscosity val-ues. Besides, it might be interesting to develop new packingsthat would be optimised for such process. The present studyaims at showing how CFD can be an interesting tool to com-plete experimental work for pressure drop and liquid hold-updetermination for CO2 capture process development. For suchdevelopment the two-phase ow within the packing has to bewell understood at local scale to be further used at large scale.Petre et al. (2003) proposed a rst approach based on meso-scale CFD three-dimensional calculations within said repre-sentative elementary units, called REU. The obtained resultsin REU are further used in a zero-dimensional model for drypressure drop calculations at large scale. In the latter work, itis shown that most of the pressure drop is caused at the criss-crossing junctions within packing metal sheets. Such resultsare of qualitatively interest but, as discussed in Raynal et al.(2004), they may not be considered for quantitative purposessince the considered REU geometry does not correspond to afully representative geometry. It is believed that in order to berepresentative, the geometry must correspond to the smallestperiodic element which can be found in the real packing ge-ometry. Besides those calculations were limited to wet pres-sure drop that is without liquid. More recently, Ataki and Bart(2006) made CFD simulations for a metal structured packing,the Rombopak 4M, with an interesting approach consideringfully periodic elements. The latter packing is made of lamel-las and differs signicantly from common structured packingsmade of continuous metal sheets. In the present study, the cho-sen geometry corresponds to the well-known Mellapak 250.Ypacking from Sulzer. Due to the complexity of the structuredpacking geometry and the limited CPU resources of calcula-tors, it is impossible to run CFD simulations at the scale ofa column while taking gasliquidwall interactions into ac-count at the scale of the liquid lm. It is thus important topropose original approaches that enable calculations at largesize with models representative of phenomena occurring atsmall scales.

In the present study, CFD calculations have been carried outin three steps at different scales for the Mellapak 250.Y struc-tured packing. The calculation strategy is described in Section2. Section 3 shows the results obtained at small scale, dealingwith gasliquidwall interactions at the liquid lm scale. Sec-tion 4 shows the results obtained at the so-called meso-scalewhich are based on a geometry corresponding to the smallestperiodic element in a real three-dimensional packing. Section5 is dedicated to calculations run at large scale which enablediscussion about internals and packing designs. Indeed, lesswork has been made on gas distributor devices and their im-pact on the ow eld within the column. Recent studies haveshown that CFD could be an appropriate tool to determine thegas ow eld above the distributor and below the packed bed

(see Wehrli et al., 2003); however, CFD calculations in suchstudies were done around the distributing devices only and notwithin the packed bed. One of the objectives of this study is toshow that the present strategy is able to determine gas distrib-utor effects on the ow eld within the layers of packings andallows for internals design optimisation.

2. Calculation strategy

Fig. 1a shows a sketch of an industrial gas treating columnoperating in a counter-current ow mode. It contains liquid andgas distributors at top and bottom of the column, respectively.In between distributors, the packed bed is made of layers ofpacking elements which are turned 90 from each other asshown in Fig. 1b. The column diameter for a CO2 absorbermay reach up to 10m, the total height being about 2030m.Fig. 2 shows close views of the packing. Fig. 2a shows thetriangular based channel geometry; the base, B, and the height,h, of the channel are 22.4 and 11.6mm, respectively, the anglewith horizontal, , being 45. Fig. 2b shows a closer view ofthe metal sheet on which one observes the small scale walltexture. The latter can be represented as a sinusoidal-shapedwall as shown in Fig. 2c, the periodic length, , and amplitude,A, being, respectively, 2.8 and 0.3mm. Solving the gasliquidow on such a large range of scales is not possible so farfor computing resources reasons. However, it is shown in thepresent study that the combination of three types of calculationmakes simulations for a whole column possible.

Fig. 1. (a) Sketch of an industrial gas treating column, (b) picture of structuredpacking elements as installed within the column and (c) sketch of the smallestperiodic element within the structured packing.

7198 L. Raynal, A. Royon-Lebeaud / Chemical Engineering Science 62 (2007) 71967204

Fig. 2. Pictures of the Mellapak 250.Y (a) at channel scales, (b) close view and (c) details of wall texture cut-view.

Fig. 3. Sketch of the present calculation strategy. The type of calculations,the input and output used at each step and the corresponding geometry andcharacteristic length scales are given.

The calculation strategy is illustrated by the dawning shownin Fig. 3. On the latter sketch, the type of simulation, theinput and the output of the simulations and the associated

geometry are given. The rst type of calculation consistsof two-dimensional gasliquid simulations using the VOFmodel as described in Raynal and Harter (2001) and used forgas/liquid ow simulations in packings by Raynal et al. (2004).In the latter study, the wall was considered to be smooth, andcalculations were made for very large liquid owrates corre-sponding to a co-current ow application. Present calculationsconsider the gasliquid ow at the scale of the corrugationwhich enables to take into account the wall texture and itsinuence on the liquid ow. From such calculations, one canessentially determine the liquid hold-up within the packingand the velocity at gasliquid interface (see Fig. 3, top). Thesecond type of calculations consists of three-dimensional gasow simulation in the smallest periodic element representativeof the real structured packing geometry (see Fig. 3, middle).Those calculations give relationships between pressure dropand gas supercial velocity. The presence of liquid is indirectlytaken into account by rst converting the supercial gas ve-locity into an interstitial gas velocity taking the liquid hold-upinto account, and second by choosing a moving boundary con-dition at walls via imposing the velocity corresponding to theliquid velocity at interface at walls. These two informations,liquid hold-up and liquid velocity at gasliquid interface, arethose obtained at the rst step. Last, simulations are carried outat a very large scale considering the packed bed as a porousmedia (see Fig. 3 bottom), the latter being characterised bypressure drop coefcients obtained at step two. From suchcalculation at large scale one is able to study the gasliquiddistributors/packings interactions for achieving optimumcolumn design. Each steps are described in details in thefollowing sections.

L. Raynal, A. Royon-Lebeaud / Chemical Engineering Science 62 (2007) 71967204 7199

Fig. 4. Views of the two-dimensional computational domain used for the simulations at small scale: (a) case of smooth walls and (b) case of walls withtexture, details.

2.1. Calculations at small scale

In the case of a uniformly wetted smooth vertical wall, andassuming that the liquid ows as a laminar and fully developedlm with no gas interaction, the thickness of the liquid lm, e,is given by

e =(3qLL

g

)1/3, (1)

where qL is the liquid ow rate per unit width of wetted surface,or specic ow rate, L the liquid kinematic viscosity and g theacceleration of gravity. The specic ow rate is given by theratio of the liquid load, QL, to the geometric area, aG. Sincethe work of Bravo et al. (1985), this model has been commonlyused to deduce the liquid holdup, hL, from the liquid thickness,e, using

hL = e aG. (2)For industrial conditions, the liquid load varies from approx-imately 10 to 50m3 m2 h1. This corresponds to liquid lmthickness varying from 0.2 to 0.4mm. Since these values areof the same order of magnitude as the amplitude of the surfacetexture at walls (A = 0.3mm), the latter has to be considered.The same model gives the following expression for the liquidvelocity at the interface, or effective liquid velocity as calledby Bravo et al. (1985), UL,eff :

UL,eff = 32UL = 32(

gq2L3L

)1/3, (3)

whereUL is the averaged liquid velocity of the liquid lm givenby the ratio of the specic liquid ow rate, qL, to the liquid lmthickness. The objectives of the rst step calculations are toprovide the liquid hold-up and the liquid velocity at the liquidlm interface. Both the informations are compared to the one-dimensional laminar falling lm model and are further used in

the second step of the present methodology for the calculationsat meso-scale.

The rst step calculations thus consists in simulating thegasliquidwall interaction at the liquid lm scale. It is donewith the VOF approach. This two-phase ow model allowsto capture the interface between two non-interpenetrating u-ids. It has been successfully used by Raynal et al. (2004) inthe case of co-current ow at very large liquid loads owingalong smooth walls in a two-dimensional geometry corre-sponding to a high geometric area packing. Valluri et al.(2005) used a similar technique to simulate liquid lms ow-ing down over two-dimensional structured wavy surfaces.While these authors discuss the impact of the sinusoidal-likewavy deformations amplitude and frequency on the effectivearea, they do not use their results for estimating the liquidhold-up.

Present calculations aim at determining the inuence of thewall structure on the liquid hold-up for the Mellapak.250Y ge-ometry as shown in Fig. 2. Simulations have been carried outwith the Fluent.6.2 commercial code. The VOF approach hasbeen used in the steady calculations approach, with the implicitscheme for the interface reconstruction and the HRIC mod-ule for the VOF solver. Two geometries were used. One withsmooth walls and the other with texture on walls which con-sists of a sinusoidal-like structure. The whole computationaldomain is similar to the one used by Raynal et al. (2004), andonly close views are shown in the present article. Fig. 4a showsthe central periodic element within which the liquid hold-upis determined for the case of smooth walls. Fig. 4b shows acloser view near the wall in the case of wall with texture.The grid is adapted to the computed case in such a way thatat least 810 grid points are within the liquid lm. Most ofthe grid is made of structured mesh, the number of cells be-ing about 10,000 cells. The liquid hold-up is determined viathe calculation of the liquid fraction in the central periodicelement.

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2.2. Calculations at meso-scale

The second step is done in a similar way as performed byPetre et al. (2003). It consists of calculations in a computa-tional domain corresponding to the smallest periodic geometrycharacteristic of structured packings as shown in Fig. 1c. Thecomputational domain corresponds to the volume between thetwo opposite smooth metal sheets. Flow is considered periodicin the z and y direction, a lineic pressure drop, DP/L, beingimposed along the z direction. Calculations are run in steadymode until the surface averaged gas velocity z component isconstant. This requirement is much more demanding than justthe standard 103 requirement on residuals values. Conver-gence in terms of mass ux often requires about 1000 iterationsand residuals values are less than 105. The computational do-main contains 150,000 tetrahedral cells. To study the viscousow model inuence, laminar and standard k. turbulent mod-els have been used. VOF simulations in such a complex three-dimensional geometry are considered to be too demanding interms of CPU resources and simulations with gas ow only areconsidered. However, boundary conditions have been adaptedin two ways in order to take indirectly the liquid inuence intoaccount. First, since part of the volume should be occupied bythe liquid, the velocity obtained from the calculations, UCFD,has to be corrected by the liquid hold-up. Indeed a given valueof pressure drop would be reached at a lower gas velocity withliquid ow than without any liquid ow. To determine the gascapacity F -factor, F, for comparison with literature data, thefollowing expression is thus used:

F = G UCFD (1 hL), (4)where the liquid hold-up is the result of the calculations run atrst step of the present strategy.

The second adaptation consists in modifying the boundaryconditions at walls. First, it is commonly accepted that, belowthe loading point, the liquid hold-up in packed columns is notaffected by the gas ow (Stichlmair et al., 1989; Suess andSpiegel, 1992). Second, as discussed by Sidi-Boumedine andRaynal (2005) or Alix and Raynal (2006), one observes from to-mographic measurements across structured packings that thereis almost no axial evolution of the liquid ow distribution. Thismeans that the main liquid velocity component is in the verti-cal direction. From these two observations, it is considered thatthe presence of liquid can be modelled by imposing a movingwall condition as boundary conditions at walls in the verticaldownward direction. The velocity that is imposed is given bythe value of the liquid velocity at interface obtained at rst stepof present methodology.

From such calculations, one can determine the relationshipbetween the pressure drop coefcient in the vertical direction,Kz, and the gas and liquid ow characteristics.

K = P1/2GU2G

1L

= K(QG,QL, L),

KZ = P1/2GU2G 1

L= KZ(QG,QL, L). (5)

2.3. Calculations at large scale

Once liquid hold-up, gasliquid interactions and pressuredrop characteristics are known at small scale, they are mergedtogether and it enables the determination of pressure drop char-acteristics of an equivalent porous media which is used in thisthird step consisting of CFD calculations at large scale. This laststep consists in considering a computational domain which cor-responds to the whole absorber at industrial scale as sketchedin Fig. 1a. The packed bed is considered as successive layersof porous volumes. Layers of packings, made of parallel metalsheets, are considered as an anisotropic porous media with twoequal pressure loss coefcients, Kz = K(x or y), in two direc-tions, and one pressure loss coefcient in the third direction,K(y or x), which is innite to model the inuence of the pres-ence of perpendicular plates. In practice, it has been consideredthat the third pressure loss coefcients would be 1000 timeslarger than the two others. Since successive layers are turnedby 90 between each other when installed in columns (seeFig. 1b), Kx and Ky pressure loss coefcients are switched be-tween each other for odd or even layers, one being equal to Kzthe other being equal to 1000 Kz. The height of these layersis 210mm corresponding to the height of industrial Mellapak250.Y packing elements.

Inlet and outlet effects may be considered via the use ofdifferent distributors. The inuence of the liquid distributor isquite well known and many articles have shown the importanceof distributor design in particular in terms of number of drippoints per unit surface (see Bonilla, 1993; Billet, 1995). In thepresent study, we assume that the liquid distributor is ideal. Theliquid distributor is thus modelled as a simple pressure jumpsurface of zero thickness and it is assumed that the liquid ishomogeneously distributed. This means that there is no radialdistribution of the pressure loss coefcients Kz. For gas inlet,the distributor is a straight tube with a single open orice at thecentre of the column oriented downward as sketched in Fig. 1a.The number of cells is about 400,000 cells for a column of 1min diameter containing 10 packing layers.

3. Results and discussion

3.1. Results at small scale

Hold-up in case of a corrugated smooth wall is evaluated onone periodic element. In the case of the rough wall, hold-up isevaluated as the same way as for the one-dimensional model(see Section 2.1). A mean thickness, eCFD, is rst calculatedon a periodic element. It is then corrected to take into accountthe inclination of the plate by

ecorrect = eCFD (sin )1/3,e = eCFD (sin )1/3. (6)The liquid hold up hL is then determined by Eq. (2).

Hold-ups obtained in the case of smooth corrugated wall andin the case of wall with texture are compared with the valueof the one-dimensional model and with experimental values

L. Raynal, A. Royon-Lebeaud / Chemical Engineering Science 62 (2007) 71967204 7201

Fig. 5. Ratio of liquid hold-up to the liquid hold-up given by theone-dimensional model versus liquid load. Comparison between experimentaland CFD results for two liquid viscosities.

of Alix and Raynal (2006) and of Spiegel and Meier (1992)in Fig. 5. First, one observes that the one-dimensional modeland the CFD calculations with corrugated smooth wall are inclose agreement and strongly underestimates the experimentalresults. This is observed for liquid supercial velocity rangingfrom 1 to 250m3 m2 h1 and for viscosity ranging from 1 to2.5 cP. This shows that the liquid lm/wall interaction is notproperly modellised. Second, it is observed that the simulationstaking into account the small scale roughness of the wall give amuch better agreement with the experimental data even if somediscrepancy is still observed.

Liquid ow over a one-dimensional periodic surface has beenlargely investigated by experimental and numerical studies (seee.g., Zhao and Cerro, 1992; Pozrikidis, 1988 for creeping ow,respectively). Such studies show that the relevant parametersare the liquid Reynolds number based on the liquid ow rateper unit width, ReL = 4qL/, the ratio of the amplitude to thewavelength of the roughness, the ratio of the liquid thicknessof the corresponding one-dimensional model to the amplitudeof the roughness and the Bond number which is the ratio ofgravity effects by surface tension effects. In the present study,only the liquid Reynolds has been varied; it ranges from 40 to1000. Depending on the liquid Reynolds number, one observesthat recirculation zones form in the cavities of the roughness,as shown in Fig. 6, and that the recirculation zone grows as theReynolds liquid increase. The presence of such recirculationzones explains why the liquid hold-up is greater in the case ofrough walls than in the case of smooth walls. The liquid holdup can therefore be considered as a sum of a static hold up anda dynamic hold up. The static hold up given by the volume ratiooccupied by the recirculation is limited for geometric reasonto 3.2% which represents the whole volume of the cavitiesfor the Sulzer M252.Y packing. This value is far from beingnegligible when compared to experimental values which rangesapproximately from 4% to 8% as measured by Spiegel andMeier (1992). The present calculations clearly show that thetraditional smooth wall approach cannot apply and that textureat walls has to be taken into account.

Fig. 6. Velocity elds within the liquid lm obtained with the VOF simulationsfor a liquid viscosity of 2.5 cP. (a) QL = 54m3 m2 h1 i.e., ReL = 96 and(b) QL = 161m3 m2 h1 i.e., ReL = 288.

Fig. 7. Wet lineic pressure drop, DP/L, versus F -factor.

The second important parameter that has to be determinedby these calculations at small scale is the velocity at gasliquidinterface. The present CFD calculations show that the velocityat the interface is not affected by the presence of recirculatingzones. In the case of a rough wall, one actually observes that theliquid velocity at the interface is equal to the interfacial velocitycalculated by the one-dimensional model of an innite smoothwall with a discrepancy less than 2%. It is thus reasonable totake the value of the one-dimensional model for the velocity asgiven in Eq. (3) for the calculations at larger scale.

3.2. Results at meso-scale

Calculations at meso-scales have been run rst without anyconsideration of the liquid phase, that is for dry packing, andsecond with the inuence of the liquid being taken into ac-count. The results obtained for the wet pressure drop per unitlength versus the F -factor are shown in Fig. 7. Results for drypressure drop are rst discussed. The closed symbols corre-spond to the different models which have been used for the gasow, laminar or standard k. turbulent model. Present resultsare compared with the experimental data obtained by Spiegeland Meier (1992). One rst observes that, while one would ex-pect a power law of 1 for laminar ow and a power law of 2

7202 L. Raynal, A. Royon-Lebeaud / Chemical Engineering Science 62 (2007) 71967204

for turbulent ow, both models give results with power lawsin between these two values, 1.4 and 1.8, respectively. Thismay be explained by the fact that strong changes in geometrythat occur at a similar scale as the hydraulic diameter inducea complex organisation of the ow differing signicantly fromcommon fully developed ow. Second, one observes that thebest agreement with experimental data is obtained with sim-ulation run assuming laminar ow. The straight line in Fig. 7indeed corresponds to the best t as suggested by Spiegel andMeier (1992) for their own experimental data. The agreementwith present CFD calculations assuming laminar is very goodwhile turbulent ow models give much higher pressure drops.This agreement with laminar ow model is obtained while gasReynolds numbers cover a wide range from 400 up to 2 104where the gas ow Reynolds number, ReG is given by

ReG = GVSG/ sin()4/aG(1 hL)G

, (7)

where 4/aG corresponds to the hydraulic diameter being fourtimes the hydraulic radius, the ratio of the wetted perimeter tothe surface area and , the packing porosity (here, = 0.95).With such high Reynolds numbers in a such complex geometry,one would expect a better agreement with turbulent models.It is thus believed that the turbulent models are not adaptedfor such a complex three-dimensional ow; the laminar model,which gives satisfactory results, is thus recommended.

Results for wet or irrigated pressure drops are also shownin Fig. 7. Since moderate effects are observed, only one liq-uid load is considered. The latter corresponds to the maximumvalue one would expect for CO2 absorption process, that is50m3 m2 h1. As obtained from calculations at rst step, suchliquid load corresponds to a velocity at gas/liquid interface, and,in the present strategy, to a wall velocity, Uw, of 0.5m s1.From Fig. 7, one observes that imposing a liquid velocity atwalls and taking into account the volume fraction occupied bythe liquid, lead to two modications on the pressure drop curvecompared to that obtained without any liquid. First, for low gasvelocity, or low F -factor, less than approximately 0.7, one ob-serves an important increase in the pressure drop. Second, athigher F -factor values, the curves at Uw = 0.5m s1 closelyfollow the curve obtained for dry packing, the offset betweenthe two curves being due to the correction of the F -factor, tak-ing into account the volume fraction occupied by the liquid asshown in Eq. (4). The present results are compared to values ob-tained with the Sulpak software developed by Sulzer and basedon the work of Spiegel and Meier (1992). It is seen that thepresence of liquid is well reproduced for low F -factor valuesand a close agreement for intermediate F -factor values. How-ever, when gas velocity becomes too large, the discrepancy be-tween experimental and CFD values becomes important. Thisdifference happens when the gasliquid interaction increases,inducing an increase in experimental pressure drop values. Thishappens above loading point and just below the point corre-sponding to 80% of ooding. The point corresponding to 80%of ooding is shown as a point with larger size. Note that thechoice for an industrial column diameter would be such thatthe ooding rate is less than 80%, which means pressure drops

Fig. 8. Velocity elds and velocity magnitude contours at the exit of the rstfour packing layers; colourmap in m s1 ranging at 5% around 1.48m s1.

less than 2mbar m1 for this precise case. From Fig. 7, oneobserves that it is precisely above that value that discrepancybetween experiments and calculations becomes important. Thisdifference is thus explained by the fact that present methodol-ogy is not able to determine strong gasliquid interactions as ithappens above loading. However, it is not of great importancesince this region corresponds to too high pressure drops for theCO2 capture industrial application.

One can thus conclude that, for the case of CO2 industrialcapture plants which would be designed for liquid load be-tween 10 and 50m3 m2 h1 and with maximum pressure dropper unit length of about 2mbar m1, the agreement betweenpresent calculations and experiments is satisfactory. The rela-tionship between pressure drop and local values of gas and liq-uid velocities may thus be considered for use at larger scale.

3.3. Results at large scale

Calculations at large scale have been run for different owconditions and for different column diameters. However, only

L. Raynal, A. Royon-Lebeaud / Chemical Engineering Science 62 (2007) 71967204 7203

Fig. 9. Velocity magnitude contours at the exit of the rst four packing layers; colourmap in m s1. Two assembly procedures are compared (a) rst packinglayer plates are normal to the x direction and (b) rst packing layer plates are parallel to the x direction.

one case is presented here, the purpose of the present pa-per being to demonstrate the interest of the approach. Simu-lations have been carried out for an equivalent liquid load of50m3 m2 h1 and with a gas velocity of 1.47m s1 whichcorresponds to a F-factor of 1.62 Pa1/2 that is almost at 80%of ooding as previously discussed. Fig. 8 shows the velocityeld in the y = 0 plane and the velocity magnitude contours insections corresponding to exits of the rst four packing layersof the packed bed. At the exit of the rst layer, the ow is ob-served to be strongly heterogeneous, and an important portionof the section corresponds to local velocities more than 5% dif-ferent from the average velocity. Three layers of packings areneeded to ensure a uniform ow. One also observes that thevelocity contours are oriented along the x direction at the exitof the rst packing layer, and along the y direction at the exitof the second packing layer. This is due to the fact that theequivalent porous media is turned by 90 between each layeras previously discussed. This type of simulations consideringthe structured packing as an anisotropic porous media is foundto be able to reproduce the impossibility for the gas ow toredistribute itself across the section in the direction normal tothe plates which is characteristic of structured packings. Fig. 9shows the impact of packing layers orientation. In case (a) therst packing layer is oriented such that packing walls are nor-mal to the inlet tube direction; in case (b) the rst packing layeris oriented such that packing walls are parallel to the inlet tubedirection. The above packing layers are turned by 90 to eachother as usual. The colours correspond to the velocity contoursranging from 1.35 to 1.7m s1. One observes that in the rstcase strong heterogeneities are observed along the x direction,while they are along the y direction in the second case. More

importantly, the heterogeneities are observed to be stronger forthe case (a) than for the case (b). From this type of calcula-tions, it thus possible to choose the most appropriate assemblyto avoid local ooding of the column. Or it is possible to testdifferent gas distributors design (not shown here) to obtain themore homogeneous ow as possible.

Note that the total pressure drop is 22.2mbar. This shows thatthe pressure drop may be important even for a moderate sizecolumn (4m in total height). Singular pressure drops at inletand outlet are not negligible in such columns where the packinginduces pressure drop as low as 2mbar m1. This shows thatsuch calculations make optimised design of distributing devicespossible. A good design would correspond to cases where theow is as homogeneous as possible within the shortest distancein the packed bed for minimum distributor pressure drop. Todo so, it is of course very important to take the packing intoaccount as showed in the present study and not the distributoronly as it is commonly done.

4. Conclusions

This study proposes a methodology that enables represen-tative CFD calculations at large scales. It is based on dif-ferent types of calculations run at three different scales, theresult obtained at one scale being used at the larger scale.The methodology gives satisfactory results in the case of lowgasliquid interaction, that is below the loading point and forlow and intermediate liquid ows. Since these two conditionsare those that would be encountered in CO2 capture plants, thepresent strategy is well adapted for calculations of gasliquidow in column containing structured packings. It will allow

7204 L. Raynal, A. Royon-Lebeaud / Chemical Engineering Science 62 (2007) 71967204

for the development and design of appropriate packings anddistributors.

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A multi-scale approach for CFD calculations of gas--liquid flow within large size column equipped with structured packingIntroductionCalculation strategyCalculations at small scaleCalculations at meso-scaleCalculations at large scale

Results and discussionResults at small scaleResults at meso-scaleResults at large scale

ConclusionsReferences