packed column base

Published: March 24, 2011

r 2011 American Chemical Society 5680 dx.doi.org/10.1021/ie1020206 | Ind. Eng. Chem. Res. 2011, 50, 5680–5692

ARTICLE

pubs.acs.org/IECR

Designing a Packed Dividing Wall Column for an AromaticsProcessing PlantIgor Dejanovi�c,† Ljubica Matija�sevi�c,† Helmut Jansen,‡ and �Zarko Oluji�c*,§

†Faculty of Chemical Engineering and Technology, University of Zagreb, Savska cesta 16, HR-10000 Zagreb, Croatia‡Julius Montz GmbH, Postbox 530, 40705 Hilden, Germany§Process and Energy Laboratory, Delft University of Technology, Leegwaterstraat 44, 2628 CA Delft, The Netherlands

ABSTRACT: This paper introduces a comprehensive design method assembled using facilities of a commercial software packagethat complemented by Excel programs, which contain own column dimensioning and well established cost estimation procedures,enables proper assessment of the industrial viability of a dividing wall column (DWC) equipped with corrugated sheet structuredpackings. The heart of the performance simulation tool is a detailed four-column model that in conjunction with a simple,theoretically founded short-cut method providing reliable initial values for liquid and vapor splits and a simple but effective objectivefunction for design optimality indication allows determination of the adequate stage and reflux requirement of a DWC. Theproposed dimensioning method enables a close approach in accuracy to that required at the stage of conceptual design for purposesof making a bid by an equipment manufacturer. Compared to a two-columns-in-series configuration, as employed in an aromaticsprocessing complex within a refinery, a DWC equipped with state-of-the-art structured packing and auxiliary internals requiresapproximately 43% less energy to deliver three fractions at required product specifications. This, accompanied by savings of nearly51% based on total annualized costs, indicates that implementing a DWC could lead to a significant increase in profitability ofaromatics processing plants.

1. INTRODUCTION

Being the most widely used and most energy intensive amonglarge scale separation techniques, distillation became the maintarget of efforts oriented toward increasing the sustainability ofprocess industries.1 However, the implementation of energy-saving solutions is often capital intensive and process industriesare generally reluctant to implement them if this is not associatedwith significant improvement in the profitability of a plant. This isparticularly the case in the petroleum refining world. As elabo-rated in greater detail by Hartman et al.,2 catalytic reforming toprocess aromatics is an important economic factor for modernrefineries. Reactor effluent stream is rich in benzene, toluene,ethylbenzene, xylenes, and heavier aromatics, and an aromaticscomplex usually contains several distillation columns arranged intrains (sequences) to recover and separate the aromatic compo-nents into individual products and/or certain component-richfractions. A comprehensive review of the dividing wall column(DWC) state of the art, including a survey of application-relatedpatents, indicated that aromatics complexes offer various oppor-tunities for implementation of a DWC.3 The present paper isconcerned with design of a conventional, three-product DWCsuitable for a specific aromatics complex situation as encounteredin a Croatian refinery.

The energy-saving potential in this type as well as in manyother applications is significant and can be estimated withconfidence using different simulation methods. The papers byTiantafyllou and Smith4 and Segovia-Hernandez et al.5 providesome quantification in this respect, and the obtained results agreewell with the numbers reported for some real industrial applica-tions in a paper by Kaibel et al.6

However, as indicated by Dejanovi�c et al.,3 and in a recentstate-of-the-art paper by Asprion and Kaibel,7 the columns

dimensioning procedures still belong as proprietary knowledgeto a few equipment manufacturers active in this field. In order toarrive at total annualized costs, to enable comparisons of alter-natives, some dimensioning-related efforts have been undertakenin academic publications, e.g., refs 8 and 9. However, the natureof applied approximations/simplifications is such that it may leadto erroneous conclusions on both the process design side and theeconomics side.

One should realize that hydraulic design of the partitioned partof a DWC is a delicate activity, and that pressure drop on twosides of the wall must be equal. If not properly arranged in thedesign phase, by adjusting the necessary amount of flow resis-tance exhibited by internals used in conjunction with fixedspecific liquid flows, the equalization of the pressure will beimposed by nature, i.e., by spontaneous adaptation of vapor flowsin two sections. This will inevitably lead to establishing operatingliquid to vapor flow ratios that could differ from that required toachieve the desired degree of separation in beds on both sides ofthe partition wall. This most distinctive feature of hydraulicdesign of a DWC has not received adequate attention in the openliterature so far.

An objective of this paper is to fill this gap, and as will bedemonstrated later on, the method proposed in this paper allowsa close approach to actual design practices, as adopted by theequipment manufacturing company Julius Montz, Hilden,Germany, the pioneer in the field of design and construction ofpacked DWCs.

Received: October 5, 2010Accepted: February 24, 2011Revised: January 20, 2011

5681 dx.doi.org/10.1021/ie1020206 |Ind. Eng. Chem. Res. 2011, 50, 5680–5692

Industrial & Engineering Chemistry Research ARTICLE

With proper column shell dimensions and internal configura-tion, it is possible to obtain reliable installed equipment costestimates and consequently arrive at total annualized costs(TACs) with enough confidence to allow a realistic comparisonwith established, two-columns-in-series configurations, i.e., aproper assessment of potential benefits of implementing aDWC. As shown in what follows, the results of this study clearlyindicate that adopting a DWC as the standard design foraromatics plant applications could lead to increased profitabilityof complex refineries.

2. DESIGN CASE

Upon a recent decision to concentrate mainly on fuel produc-tion, the aromatics complex at the INA (presently MOL Group)refinery in Sisak, Croatia, has been reduced to a minimum, i.e., toa direct separation sequence containing two columns—theso-called “platformate splitter” and the “benzene recoverycolumn”—connected in series (see Figure 1) to separate theplatformer reactor effluent stream which contains some 40components into three fractions: (1) C5�C6 gasoline contain-ing no more than 1.5 mass % benzene, (2) a benzene-rich cut(BRC) containing 68mass % benzene, and (3) a heavy reformatestream (heavies) containing toluene, ethylbenzene, xylenes, andheavier components with no more than 0.5 mass % benzene.

For purposes of this study the 40 components contained in theactual feed stream have been lumped together into a representa-tive 15-component mixture. The mass flow rates and composi-tions of the feed and product streams as considered in the presentsimulation study are shown in Table 1.

In order to provide an appropriate basis for the evaluation andcomparison of related costs, two conventional columns areconsidered here as new designs, and to provide completeinformation both options, i.e., tray and packed columns, areconsidered.

3. STAGE AND REFLUX REQUIREMENTS

The two columns of the base-case configuration (see Figure 1)have been simulated using detailed methods available in Chem-CAD. Regarding the simulation approach, a DWC, shownschematically in Figure 2, has significantly more degrees of

freedom than a conventional distillation column. For instance,the number of stages for six sections needs to be provided.Additional parameters indicated as circles in Figure 2, specific tothe internal configuration as encountered in a DWC, are liquidand vapor splits above and below the partition wall, respectively.A good set of initialization data is essential to ensure convergence

Figure 1. Conventional, two-columns-in-series configuration (indirectsequence) for separation of a three-component or multicomponent feedinto three specified products or fractions, respectively.

Table 1. Base Case Feed and Product Compositions

stream name

feed C5�C6 BRC heavies

total flow [t h�1] 31.74 6.94 3.70 21.10

component mass fractions

n-butane 0.019 0.088 0.000 0.000

isopentane 0.064 0.291 0.000 0.000

n-pentane 0.045 0.206 0.000 0.000

2-methylpentane 0.080 0.351 0.026 0.000

n-hexane 0.043 0.050 0.270 0.000

benzene 0.086 0.013 0.680 0.005

3-methylhexane 0.020 0.000 0.024 0.026

toluene 0.247 0.000 0.000 0.373

ethylbenzene 0.035 0.000 0.000 0.053

p-xylene 0.042 0.000 0.000 0.064

m-xylene 0.122 0.000 0.000 0.183

o-xylene 0.055 0.000 0.000 0.083

m-ethyltoluene 0.047 0.000 0.000 0.071

1,3,5-trimethylbenzene 0.077 0.000 0.000 0.116

1,4-diethylbenzene 0.017 0.000 0.000 0.025

Figure 2. Schematic representation of a DWCwith indication of designparameters.



of the detailed simulation method. Fortunately, these can beeasily obtained using a rather simple but in this respect veryeffective short-cut method, introduced recently by Halvorsen10

and Halvorsen and Skogestad.11,12

Following suggestions of Halvorsen and Skogestad,9 theso-called Vmin diagram method has been implemented in Chem-CAD. Since the original method requires dealing with purecomponents, the 15-component feed has been represented bythree representative key components: 2-methylpentane,benzene, and toluene. Then, the Vmin diagram was constructedby rigorous simulation of possible binary splits, with the stagenumber being at least 4 times the minimum number of stages,4Nmin, and mass recoveries of key components set to 0.999.From that diagram, appropriate estimates of required designparameters for initialization of a detailed simulation wereobtained. Details related to Vmin diagram application in conjunc-tion with a rigorous simulation model and correspondingnumbers can be found elsewhere, e.g., Dejanovi�c et al.13

Detailed simulations have been performed using the so-called“four-column model”, shown schematically in Figure 3. Thisconfiguration was the preferred choice in this case, because theprefractionator and main column sides of the partitioned part ofthe column are represented by separate column sections, whichmakes the dimensioning effort more straightforward. Anotherreason is that this is the only configuration that can be used fordynamic simulation with the purpose of the control system study.A detailed elaboration of relative advantages/disadvantages of

four different configurations that can be used to simulate theperformance of a three-product DWC can be found elsewhere,e.g., Dejanovi�c et al.3

The first step was to restore feed composition to the full 15components. Benzene mass fractions in the distillate and the sidedraw liquid flow rate were set to be the same as in the base-casesimulation, while the initial value for reboiler duty was set to givethe required minimum vapor flow. Liquid and vapor split ratioswere adjusted until the minimum possible mass fraction ofbenzene in the bottoms was achieved, which was the same asin the base case.

The next step was to determine the actual number of stages ineach section. This was done by reducing the number of stages insections keeping the mass fractions of benzene in the distillateand bottoms, as well as liquid side draw flow rate, constant. Forevery converged case, an optimization was performed using theoptimization tool built in ChemCAD, with the objective functionbeing minQr, and the independent variables the liquid and vaporsplit ratios.

To arrive at the optimum combination of reflux ratio and thenumber of stages, the empirical objective function, min N(Rþ 1),which approximates effectively the total annualized cost of a con-ventional distillation column, was used. According to Figure 4, theminimum value corresponds to 64 equilibrium stages, based on themain column stage count. The same result is obtained using allequilibrium stages contained in the DWC.

In the present case the number of stages at the prefractionatorside and the number of stages at the main column side are equal,i.e., 22. This means that the total number of stages contained inthe DWC is 86, which means eight stages more than required intwo sieve tray columns connected in series. The number ofequilibrium stages of compared configurations and the corre-sponding reflux ratios are given in Table 2. The simulation resultsfor conventional, two-columns-in-series sequence and a DWC

Figure 3. Four-columnmodel used to simulate the operation of a three-product DWC.

Figure 4. Plot indicating optimum number of stages for a DWC.

Table 2. Base Case and DWC Simulation Results

base case

C1 C2 C1 þ C2 DWC

number of stages 40 38 78 86

reflux ratio 1. 70 2.39 � 2.80

reboiler duty (MW) 3.55 2.63 6.18 3.50

condenser duty (MW) �3.16 �2.51 �5.67 �2.76



shown in Table 2 indicate that a DWC would require approxi-mately 43.3% less energy, based on reboiler duties.

Figure 5 shows the algorithm for stage and reflux requirementcalculations associated with a three-product DWC.

4. COLUMN DIMENSIONS

Upon completion of the performance calculations, the num-ber of equilibrium stages (theoretical plates) in each columnsection was fixed as well as corresponding liquid and vapor flowrates. For the tray column option, common cross-flow sieve trayshave been chosen with a tray spacing of 0.6 m. Using trayefficiencies as experienced in similar applications, the platformatesplitter contains 61 trays and the benzene recovery columncontains 59 trays, with the feed stage in both cases in the middleof the column. For packed conventional columns and packedDWC, Montz-pak B1-350 MN, a state-of-the-art, high perfor-mance corrugated sheet structured packing was chosen. Bedheights in conventional and partitioned parts have been deter-mined in accordance with the number of contained stages,assuming an HETP value of 0.4 m. This number including the

common safety margin is based on total reflux test data obtainedwith Montz-pak B1-350MN using the chlorobenzene/ethylben-zene system at atmospheric pressure.14

The number of packed beds to be installed in rectification andstripping sections of a packed column has been determined basedon a rule of thumb indicating that the single bed height shouldnot exceed that equivalent to 20 equilibrium stages. If more than20 stages are required in a section, then two beds should beconsidered, which implies providing additional column height toaccommodate the necessary liquid redistribution section. Thelatter consist usually of a liquid collecting device, a chevron(vane) type or chimney tray type liquid collector (see Oluji�cet al.1 or Rix andOluji�c15) installed above a narrow trough gravityliquid distributor.

For conventional packed columns and similar sections in apacked DWCwe have assumed a constant spacing of 1.8 m, whichis appropriate for the size of columns considered in this study.Witha symmetrical distribution of stages in two conventional columns,each containing around 20 stages in rectification and strippingsections, it was chosen to have two shorter beds per section. Thismeans that both packed columns contain three liquid redistributionsections, with the middle one receiving also the feed.

Since the conventional columns and the DWC considered inthe present study operate at above atmospheric pressure and thefeed is a slightly subcooled liquid (q = 1.064), the critical loadswill be on the bottom stage of each column, which is thereforetaken as the basis for determining the diameter of the two con-ventional columns as well as the shell of a DWC. The character-istic values are shown in Table 3 together with the tangent-to-tangent length (height) of the column shells for two columnsin conventional configurations, equipped with sieve trays andB1-350MN packing, respectively.4.1. Layout of a DWC Containing Structured Packings.

Arranging beds and liquid collecting and distribution becomemore demanding when the partitioned part of a packed DWC isconsidered. Figure 6 shows schematically the internal configura-tion of the DWC, indicating the number of stages in each columnsection and the corresponding vapor and liquid loads at the topand the bottom of each bed. Since the upper part of the DWCcontains 26 stages, this section consists of two beds, eachcontaining 13 stages. The bottom section contains 16 stages inone bed. Due to a pronounced difference in vapor and parti-cularly liquid loads below the feed (F) and side product (S) draw-off points, the partition wall in the lower part is in an off-centerposition. That is, a much larger cross-sectional area is neededbelow the feed point on the prefractionator side to accommodatethe feed stream that enters as slightly subcooled liquid. Adding

Figure 5. Schematic illustration of stage and reflux requirement calcula-tion procedure.

Table 3. Dimensions and Components of the DWC and Two Conventional Configurations

configuration

C1/C2 (trays) C1/C2 (packings) DWC (packings)a

column top pressure (bar) 1.7/2.7 1.7/2.7 2.7

shell diameter (m) 2/2 1.6/1.8 1.7

shell height (m) 40.5/39.5 27.1/27.5 37.3

number of trays or packed beds 61/59 4/4 7 (4)

number of distributors � 4/4 7 (4)

number of liquid catchers � 4/4 7 (4)

number of support grids � 4/4 7 (4)aNumbers in parentheses indicate devices placed in partitioned part of the column.



the feed to the liquid coming from the bed above as well as acertain amount of liquid obtained due to direct condensation ofthe ascending vapor upon contact with subcooled feed liquidmakes the liquid load of the bed below the feed nearly 5 timeslarger than that on the main column side. Also, on the maincolumn side of the partition wall there is a side draw where sideproduct is taken out of the column as a liquid. Owing to a rathersmall vapor load this section requires a much lower cross-sectional area than the heavily liquid loaded stripping sectionon the prefractionator side. Although the vapor and liquid loadsin sections above the feed and draw-off differ to a certain extent,the partition wall is, as indicated in Figure 6, placed in the center.This appeared to be a convenient choice in the present case, andplacing a partition wall off-center, as required in the sectionsbelow the feed and draw-off stages, can easily be arranged byutilizing a nonwelded wall, which is a well-established practice inthe case of DWCs equipped with structured packings.16

One should note that welding a partition wall implies separat-ing the shell into two semicircular sections with equal cross-sectional areas. Owing to different loads in the prefreactionatorand themain column, this leads always to underdesign on the lesscritical side. This was one of the main obstacles for widerimplementation of DWC in the beginning years, which, pre-sently, can be circumvented by placing the partition wall off-center, in the most appropriate way. Therefore, the introductionof a nonwelded, self-fixing partition wall in 1996 is considered tobe a milestone in the development of DWC technology.1

Regarding the dimensioning procedure, the outgoing pointis the overall shell diameter determined using conventionalmethods for the stage with the largest vapor load. The sum ofcross-sectional areas of the partitioned part of the column, whichhowever may differ above and below the feed and side productdraw-off stages, should be equal to the overall one, i.e.

Ac ¼ Apc þ Amc ð1Þwhere Ac (m

2), Apc (m2), and Amc (m

2) are the cross-sectionalareas of the shell of the DWC, prefractionator, and main column,respectively. The cross-sectional area of the prefractionatorsegment of the overall cross-sectional area can be calculated from

Apc ¼ dc2

8ðΘ� sin ΘÞ ð2Þ

where dc (m) is the dividing wall column shell diameter and Θ(deg) is the angle subtended by the partition wall (see Figure 7).The latter is described as a function of the shell diameter and thedistance of the partition wall from the shell on prefractionatorside, lpcw (m):

Θ ¼ 2 arccosdc � 2lpcw

dcð3Þ

The cross-sectional area of the main column is obtainedsimply, by subtraction of prefractionator cross-sectional areafrom the total one, using eq 1.The obtained areas are then translated into diameters of an

equivalent cylindrical column. These are used in conjunctionwith local vapor and liquid loads to check the upper limit withrespect to flooding condition. This should not exceed 80% of thatwhich would cause flooding. A good indication for this is theloading point, i.e., the point of onset of loading, which can beestimated with some confidence using the appropriate equationof the Delft model, shown later on. Practically, this means that,the design point corresponds approximately with the ratio ofoperating and loading point vapor loads, i.e., the correspondingF-factors equal 1. However, the corresponding pressure dropshould not exceed 3 mbar/m. In other words, this pressure dropis taken as the upper limit value for design purposes.If the calculated pressure drop exceeds the limit in one of the

partitioned sections, a possibility is to use a coarser packing or thesame packing with a larger corrugation inclination angle. This is apractical option if the bed height is lower than the section height,

Figure 6. Internal configuration, indicating the number of stages in eachbed and the corresponding liquid and vapor flows at the top and bottomof each bed.

Figure 7. Geometry factors associated with placing the partition wall.



which is often the case if a part of the partition wall is placed off-center. Additional bed height is needed to accommodate therequired number of stages, because both a lower specific geo-metric area and an increased corrugation inclination angle areassociated with less efficiency. Also the lateral position of thepartition wall can be corrected accordingly, provided the otherside can have it. If not, then the shell diameter of the DWC needsto be enlarged accordingly. This means that a trial-and errorprocedure is necessary to arrange proper lateral positioning ofthe partition wall. The longitudinal positioning is related to thenumber of stages contained on the prefractionator and the maincolumn side, above and below the feed and side product draw-off

stage, respectively. If one side contains more stages and the samepacking is chosen, then the side with fewer stages will have aportion of empty space. This, however, allows installation of acoarser packing, to arrive at a lower section pressure drop, ifappropriate. One should also note that additional height of thecolumn shell may be required to accommodate inclined portionsof the partition wall. This is necessary if the lateral positions of thepartition wall above and bellow the feed differ (one or both ondifferent off-center positions), as encountered in this study.A reference design for the DWC was generated using the

Montz in-house method for establishing column dimensions.The diameter of the DWC shell is 1.7 m, based on the designpoint corresponding to 75% of the flooding vapor velocity on thebottom stage.A detailed drawing of this DWC is shown in Figure 8,

indicating seven packed beds and all auxiliary devices, i.e., columninternals of importance for cost estimation purposes, such asliquid collectors or catchers, an externally placed liquid splitter,the liquid distributors, the vapor distributor, and packing supportgrids. In all situations a narrow trough liquid distributor with driptubes is used while the type of the liquid collector depends on thespecific liquid load. For the specific liquid loads above 20 m3/m2 h, a chimney type collector is a preferred choice. This device isalso a common choice for side product draw-off location regard-less of the liquid load. For lower specific liquid loads, vane(chevron) type collectors are used. In present study a chimneytray collector, placed above the vapor inlet, is used to facilitateinitial vapor distribution.The proper distribution of vapor flow bellow the partition wall

is the key to success with the design of a DWC. This is somethingthat needs to be fixed in the design phase, by arranging theindividual flow resistances to ensure equal pressure drop on twosides of the partition wall. This is presently done by manufac-turers only, i.e., using proprietary design methods. However, asshown and demonstrated in what follows, it can be done withrequired accuracy using recently published methods for estimat-ing the pressure drop of corrugated sheet structured packings andstate-of-the-art packed column internals, respectively.4.2. Pressure Drop of a Packed DWC. The heights of

individual column sections are indicated in Figure 8. The bedheights and corresponding cross-sectional areas and equivalentdiameters for each section are shown in Table 4. The formerhave been determined by multiplying the required number oftheoretical plates with a constant HETP value as adopted byMontz for B1-350MN for this case (HETP = 0.4 m). Oneshould mention here that the Delft model arrives at conserva-tive enough values, similar to those used by Montz in thepresent case. However, a thorough verification effort is requiredbefore adopting this method for the determination of bedheight of packed DWCs.

Figure 8. Detailed drawing of the DWC with all major pieces ofequipment indicated.

Table 4. Basic Dimensions and Estimated Pressure Drop perPacked Bed Section as Indicated in Figure 8

prefractionator main column

Section 2.1a 2.1b 1.1 1.2 2.2 2.3 2.4

h (m) 5.2 5.2 3.6 5.2 4.8 4.0 6.4

A (m2) 2.27 2.27 1.135 1.599 1.135 0.671 2.27

d (m) 1.7 1.7 1.20 1.43 1.20 0.92 1.7

Δp (mbar) 6.95 6.91 5.75 7.95 4.25 9.76 13.51



The heights and cross-sectional areas of individual beds allowdetermination of the volume of packing and in conjunction withaverage liquid and vapor loads serve as a basis for estimation ofthe pressure drop, separately, for each bed and associated liquidcollecting and distributing equipment.The estimated pressure drop for each of the beds contained in

two packed columns is shown in the bottom row of Table 4.These numbers have been generated using the Delft model (seeFair et al.17 or Oluji�c et al.18), which, as shown by Oluji�c et al.,19

can easily be adapted to account for any change in geometry ofcorrugated sheet packings. Note that during the validation of thismodel using the total reflux data obtained with the chloro-benzene/ethylbenzene system,14 it appeared that the trend ofpressure drop curves is reproduced very well; however, thepredicted values appeared to be consistently on the lower side,and the most pronounced deviation was observed at the highesttest pressure. Therefore, in the present study concerned withabove atmospheric operating pressures, the estimated valueswere multiplied by the factor 2 to generate conservative predic-tions, resulting in numbers somewhat larger than those estimatedusing the Montz in-house method.4.2.1. Pressure Drop of Corrugated Sheet Structured Pack-

ings. In order to enable direct implementation of the proposedDWC dimensioning method, the present paper contains neces-sary working equations of the Delft model, accounting, whereappropriate, for macro geometry modifications as implementedin high performance structured packings, such as Montz-pakB1-350MN considered in this study. This packing contains abend in the bottom part of corrugations and a corrugation anglelower than the common 45�.A specific feature of the Delft model is that it makes a

distinction among three characteristic angles: the corrugationinclination angle,R (deg), that strongly influences the interactionof gas streams at the crossing planes formed between flowchannels oriented in opposite directions; the effective gas flowdirection change angle at the transition between packing elementor layers, RDC (deg); and the effective flow angle of liquid, RL

(deg), which is steeper than the corrugation inclination angle dueto a more or less pronounced tendency of the liquid to flow overthe corrugation ridges driven by gravity. The latter two dependon the corrugation inclination angle and can be determined usingexpressions given later on, where appropriate. The corrugationinclination angle is a characteristic geometric feature of corru-gated sheet structured packings. The standard angle is 45�.Another commonly used (where appropriate) angle is 60�, withrespect to the horizontal. Experimental evidence on the effect ofthe corrugation inclination angle on the mass transfer efficiency,pressure drop, and capacity of corrugated sheet structuredpackings can be found elsewhere, e.g., Oluji�c et al.19,20

4.2.1.1. Basic Geometry and Flow Related Parameters. Inaddition to the corrugation inclination angle, R (deg), otherrelevant macro geometric features of a corrugated sheet struc-tured packing are the specific geometric area, ap (1/m), the voidfraction or porosity, ε (m3/m3), and the height of a packingelement, hpe (m). The requested bed height, hpb (m), is arrangedby placing the required number of packing elements or layers,npe, above each other:

hpb ¼ npehpe ð4Þ

Since the subsequent element/layers are usually rotated to eachother by 90�, to maximize lateral spreading and mixing of both

phases, the number of packing elements or layers correspondswith the number of flow direction changes the vapor phasemakeswhile ascending through the packed bed.For a packing with a smooth bend in the bottom part of the

corrugations, the vapor flow direction change angle is simply

RDC ¼ 90�þ R2

ð5ÞAssuming that only gravity and the corrugation shape affect

the liquid flow, the effective liquid flow angle, RL (deg), can bedescribed by

RL ¼ a tancosð90� RÞ

sinð90� RÞ cos a tanb2h

� �� 26664

37775 ð6Þ

where h (m) represents the height and b (m) is the width of thebase of corrugations. The corresponding length of the corruga-tion side, s (m), follows from

s ¼ffiffiffiffiffiffiffiffiffiffiffiffiffiffib2

4þ h2

rð7Þ

These three basic corrugation geometry dimensions dependon the specific geometric area of the packing, and, being highlyinfluential, should be measured upon delivery of a packing, aswell as the corrugation inclination angle, to avoid uncertaintieslater on. Assuming a standard configuration utilizing a crimpangle of 90�, which implies that b = 2h, the following expressioncan be used to determine the installed specific geometric area orto back calculate corrugation dimensions from the given area.

ap ¼ 4sbh

ð8ÞAnother important geometry-related parameter is the

V-shaped fraction of the cross section of a triangular gas flowchannel occupied by liquid film:

j ¼ 2sbþ 2s

ð9Þ

A geometry-related parameter of general importance is thehydraulic diameter of the triangular gas flow channels. Assuminguniform liquid distribution, i.e., constant film thickness, δ (m)

dhG ¼ðbh� 2δsÞ2

bh

bh� 2δs2h

� �2

þ bh� 2δsb

� �2" #0:5

þ bh� 2δs2h

ð10Þ

Since the film thickness encountered in practice and its variationsare rather small, this complex expression can be replaced withoutintroducing a significant error by a much simpler one, valid fordry channels.

dhGðdryÞ ¼ 2bhbþ 2s

ð11Þ

In addition to packing geometry related parameters, R, h, b,hpe, npe, and packing porosity, ε, the Delft model requires alsoinformation on mass flow rates, densities, and viscosities of twophases, i.e.,MG (kg/s),ML (kg/s), FG (kg/m3), FL (kg/m3), μG(Pa s), and μL (Pa s), respectively. These operating parameters



need to be provided for the top and bottom of each bed con-tained, and this information can easily be retrieved from theresults of detailed performance calculations.Corresponding superficial velocities are defined as

uGs ¼ 4MG

FGπdc2 ¼ FGffiffiffiffiffiffiFG

p ð12Þ

and

uLs ¼ 4ML

FLπdc2 ð13Þ

where dc (m) denotes the diameter of the column or the equivalentdiameter of a section in the partioned part of the column.FG (Pa0.5), which is generally known as the F-factor, representsthe vapor load of the column section. Per definition

FG ¼ uGsFG0:5 ð14Þ

The F-factor is a highly practical design and operating parameter(square root of the vapor flow kinetic energy), because it places alldistillation and similar column applications within a rather narrowrange of absolute values (roughly 0.5�4 Pa0.5) regardless of theoperating pressure.Due to the inclination of the gas flow channel within a packing

element or layer, the length of the vapor flow path is larger thanthe bed height, and the extent of this depends on the corrugationinclination angle. This also means that effective velocities ofvapor and liquid are larger than the corresponding superficialvelocity, and additional enhancement depends on the reductionin flow cross-sectional area due to packing porosity and the liquidholdup, hL (m3/m3). The working expressions for respectivelyeffective vapor and liquid velocity account for these effects:

uGe ¼ uGsðε� hLÞ sin R ð15Þ

and

uLe ¼ uLsεhL sin RL

ð16Þ

Assuming uniform wetting, the liquid holdup is simply definedas a product of film thickness, δ (m), and the specific geometricarea of the packing, ap (m

2):

hL ¼ δap ð17Þ

This simple basic expression appeared to hold well in practice, inconjunction with well-known Nusselt laminar falling film thick-ness expression adapted for inclined walls:

δ ¼ 3μLuLsFLgap sin RL

!1=3

ð18Þ

where g (m2/s) is gravity acceleration and RL (deg) is theeffective angle of liquid film flow, defined by eq 6.4.2.1.2. Working Pressure Drop Model Equations. The Delft

model makes a distinction between the preloading region, wherefilm flow conditions prevail, and the loading region, where amorecomplex fluid-dynamic situation is encountered.

ΔpΔz

¼ ΔpΔz

� �preload

Fload ð19Þ

Here, Fload is an empirical correction factor that describes theamount of pressure drop enhancement within the loading regionwith respect to the preloading region pressure drop.

Fload ¼ 3:8FGFG, lp

!2=ðsin RÞu2Ls

ε2gdhG

!0:13

ð20Þ

This correction is activated when FG/FG,lp > 1, i.e., when theoperating vapor load exceeds that corresponding to the point ofthe onset of loading, FG,lp, or, in other words, the point of depar-ture from preloading conditions. The loading point F-factor isdescribed reasonably well by the following empirical expressionintroduced by Verschoff et al.,21 which accounts explicitly for apronounced effect of flow direction change angle.

FG, lp ¼

0:053ε2gdhGFL � FG

FG

!uLs

ffiffiffiffiffiffiFLFG

r !�0:25

ðsin RDCÞ1:240@

1A

0:57 ffiffiffiffiffiffiFG

p

ð21ÞThis critical load corresponds roughly with the design point,which is usually set to 0.7�0.8 of the flood point. However, asmentioned before, if the pressure drop corresponding to theloading point condition FG/FG,lp = 1 exceeds 3 mbar/m, thelatter should be taken as the design limit.According to the Delft model, the preloading region pressure

drop consists of three major contributions: the pressure dropassociated with the vapor�liquid interaction at the interface,ΔpGL; the pressure drop associated with the vapor�vaporinteraction at crossings of open sides of triangular gas flowchannels, ΔpGG; and the pressure drop associated with thechange in flow direction at the transitions between subsequentpacking elements or layers, ΔpDC.

Δppreload ¼ ΔpGL þΔpGG þΔpDC

¼ ðζGL þ ζGG þ ζDCÞFGu

2Ge

2ð22Þ

Three major sources of flow resistance are expressed in termsof characteristic overall interaction coefficients. The overallgas�liquid interaction coefficient is given by

ζGL ¼ jξGLhpb

dhG sin Rð23Þ

The friction factor, ξGL, described by the well-known Coleb-rook and White expression,22 accounts for the effect of theroughness of the interface, which is here assumed to be equalto film thickness.

ξGL ¼ �2 logδ=dhG3:7

� 5:02ReGrv

logδ=dhG3:7

þ 14:5ReGrv

� �� 2

ð24Þ

Here, ReGrv represents the relative phase velocity Reynoldsnumber:

ReGrv ¼ FGðuGe þ uLeÞdhGμG

ð25Þ



The gas�gas interaction coefficient appeared to be a ratherstrong function of the corrugation inclination angle.

ζGG ¼ ð1� jÞξGGhpb

dhG sin R

¼ ð1� jÞð0:722Þðcos RÞ3:14 hpbdhG sin R

ð26Þ

The direction change loss coefficient expression makes adistinction between bulk and periphery of the packing. That is,in the core of a packed bed the gas flow changes direction attransitions between packing elements or layers only. However, inthe wall zone this happens also with channels ending at the wall,which means that here flow direction changes also within theheight of a packing element. This appeared to be a reason forincreased pressure drop observed with the same packing incolumns with small diameters.23

ζDC ¼ hpbhpe

ðξbulk þψξwallÞ ð27Þ

with

ψ ¼ 2hpeπdc

2 tan Rdc

2 � h2petan2 R

!0:5

þ 2πarcsin

hpedc tan R

� �

ð28ÞThe pressure drop due to direction changes in the bulk

appeared to be a strong function of corrugation inclination angleonly:

ξbulk ¼ 1:76ðcos RDCÞ1:63 ð29ÞIn the wall zone also the liquid and vapor load play an influentialrole:

ξwall ¼4092u0:31Ls þ 4715ðcos RDCÞ0:445

ReGeþ 34:19u0:44Ls ðcos RDCÞ0:779 ð30Þ

where ReGe represents the Reynolds number based on theeffective gas velocity.

ReGe ¼ FGuGedhGμG

ð31Þ

The relative magnitudes of the three different pressure dropcomponents depend on the corrugation inclination and flowdirection change angles involved. A detailed numerical illustrationof exhibited effects can be found elsewhere, e.g., Oluji�c et al.19

4.2.2. Pressure Drop of Liquid Distributors and Collectors.The pressure drop caused by liquid distributors and collectorscan easily be determined using the recently proposed model byRix and Oluji�c:15

Δpint ¼ζcc, ct, ntj2cc, ct, nt

!F2G2

ð32Þ

where ζ is the characteristic flow resistance coefficient and j isthe free (void) area, while indices cc, ct, and nt denote respec-tively the chevron type liquid collector, the chimney tray liquidcollector, and the narrow trough liquid distributor.The characteristic flow resistance coefficient is expressed as a

function of free area using the following expressions.

narrow trough liquid distributor:

ζnt ¼ 1:2½1:5� jð2:5� jÞ� ð33Þ

chevron type liquid collector:

ζcc ¼ 1:5ð2:5� 2:5jÞ ð34Þ

chimney type liquid collector:

ζct ¼ 1:2½1þ 2:5ð1� jÞ� ð35ÞThe free areas of all (narrow trough) liquid distributors and

corresponding vapor loads expressed as characteristic F-factorsare shown in Table 5, while the same can be found for differenttypes of liquid collectors in Table 6. The estimated pressure dropfor each device is shown in the bottom row of Tables 5 and 6,respectively.One should mention here that the above expressions account

for the observed liquid load effect and according to givenconstant (representative) values of coefficients the extent of thiseffect is most pronounced in the case of the chevron collector.The chosen value (1.5), however, may appear conservative in thecase of rather low liquid loads (below 3 m3/m2 h). Theexperimental data that have served as the basis for developmentand validation of the above expressions can be found elsewhere,e.g., Rix and Oluji�c.15

The results are summarized in Table 7, indicating the relativemagnitudes of the pressure drops of internals and the partitionedpart of the column with respect to the total pressure drop. Oneshould note that in the present case a constant free area (25%) for

Table 5. Pressure Drop Caused by Narrow Trough Distri-butors (nt) Installed in Different Positions (1, 2, 4, 5, 8, 9, 12)along the DWC, with Indication of Free Area and Corre-sponding Vapor Load


1 2 4 8 5 9 12

type nt nt nt nt nt nt nt

jnt 0.40 0.40 0.40 0.40 0.40 0.40 0.4

FG (Pa0.5) 1.16 1.23 1.41 1.14 1.00 1.65 1.40

Δp (mbar) 0.03 0.04 0.05 0.03 0.02 0.07 0.05

Table 6. Pressure Drop Caused by Chevron (cc) and Chim-ney Tray (ct) Collectors Installed in Different Positions (2, 3,6, 7, 10, 11, 13) along the DWC, with Indication of Free Areaand Corresponding Vapor and Liquid Loads


Position 2 3 6 10 7 11 13

uLe m3/m2h 16.5 15.5 8.9 39.3 15.8 17.9 34.7

type - cc cc cc ct ct cc ct

jcc/ct - 0.25 0.25 0.19 0.25 0.25 0.25 0.25

FG Pa0.5 1.27 1.20 1.25 1.34 0.93 1.63 1.57

Δp mbar 0.36 0.32 0.68 0.49 0.24 0.60 0.54



all liquid collectors was chosen, except for the chimney collectoron the main column side, which was reduced to 19% to ensureequal pressure drop on both sides of the partition wall.Being rather low, the pressure drop of support grids is

neglected. Also, the numbers shown in Tables 5 and 6 indicatethat the pressure drop of narrow trough liquid distributors ismuch lower than that of liquid collectors. This is mainly due to amuch larger free area available for ascending vapor than in thecase of liquid collectors that require relatively more cross-sectional area for liquid handling. In any case, as demonstratedhere, the free area of liquid collectors can serve as a means forfine-tuning the pressure drop of sections in the partitioned part ofa DWC. This can be done automatically using Excel solver, whichwill indicate the free area bound between 5 and 30%, generatingthe requested pressure drop. The dimensioning procedure for athree-product DWC is shown schematically in Figure 9.

5. ECONOMIC EVALUATION AND COST ESTIMATIONPROCEDURES

Total annualized costs (TACs) are taken as the basis forevaluation of the economic feasibility of a DWC. Since different

configurations are compared on the same basis, TACs arerepresented simply as a sum of the annual utility cost and 10%of the installed equipment cost. The latter is based on theassumption of a plant financial lifetime of 10 years. The yearlyoperation time is taken to be 8322 h, as encountered at INA. Tobe closer to the common European and United States situation,the following utility prices were considered in the present study:US $0.03/t for cooling water, US $13/t for steam, and US $130/tfor fuel oil. The latter is used because a fired heater is necessary inconjunction with a rather high temperature of the bottoms of theplatformate splitter (first column in the conventional sequence)and the DWC.

The capital costs in the present cases include those associatedwith column shell, trays, and/or packings and, for packedcolumns, also liquid collectors, liquid distributors, and supportgrids. Important external equipment components of each dis-tillation column are the reboiler, condenser, and reflux accumu-lator, respectively. The latter has not been considered in thepresent study. In order to bring the installed equipment costsestimated using the SI unit version of correlations from theDouglas textbook24 to the price level corresponding to the year2009, the corresponding annual value (1468.6) of theMarshall &Swift Equipment Cost, published in the March 2010 issue ofChemical Engineering,25 has been used.

The installed cost of a column shell (in US $) made of carbonsteel is estimated using the correlation:

Cshell ¼ fp1468:6280

� �dc

1:066h0:802c, t-t ð36Þ

where dc (m) is the column diameter and hc,t-t (m) is the tangent-to-tangent column height, while the cost factor fp depends on theoperating pressure. In the present case, for p e 3.5 bar, fp =2981.68.

Table 7. Overall Pressure Drop (Δp) of Packings and Col-umn Internals asWell as That of a DWCwith Indication of theContribution of Either of Two Sections in Partitioned Part ofthe Column

Δp (mbar)


packings mbar 41.07 41.38

internals mbar 2.73 2.41

total mbar 43.80 43.80

partitioned section mbar 14.95 14.95

Figure 9. Schematic illustration of the DWC dimensioning procedure.



For the purchased cost for sieve trays (carbon steel) as well asstructured packings and internals (both made of stainless steel),we have used the values approved by Montz as reasonableestimates for the purposes of this comparative study. These aresummarized in Table 8. For structured packings, the price isproportional to the specific geometric area of the packing. Thismeans that in the present case the value given in Table 8 needs tobe multiplied by the factor 1.4, which is the ratio of specificgeometric area of packing used in this study, i.e., B1-350MN, andthe basic type B1-250. That is, the given numbers represent thenominal specific geometric area of a packing in m2/m3.

Given values are valid for conventional tray and packedcolumns. In order to account properly for complexities asso-ciated with building a packed DWC, the cost of both packingsand related internals installed in the partitioned part is increasedby 20%, i.e., the value for standard equipment multiplied by thefactor 1.2.

The number of beds and their volumes and the number ofdistributors, collectors, and support grids contained in conven-tional and partitioned parts of the column can be retrieved bycareful inspection of the detailed drawing shown in Figure 8.Finally, to translate the purchased cost into installed cost for sievetrays and for structured packings and related internals factors 3and 2 are used, respectively.

For shell and tube condensers and reboilers, the followingexpression allows estimation of the installed cost in US dollars:

Ccond ¼ 1468:6280

� �ctypeA

0:65 ð37Þ

where A (m2) is the required heat transfer area and ctype is thecoefficient depending on the type of heat exchanger. For con-densers, ctype = 1609.13, and for kettle reboilers, ctype = 1775.26.These numbers are valid for common construction materials andthe operating pressure as encountered in the present case.

The installed cost (in US dollars) of a fired heater is estimatedfrom

Creb ¼ 1468:6280

� �½14390:71Q 0:82ð1:23þ ft þ fm þ fpÞ� ð38Þ

where Q (MW) is the reboiler duty, ft is the correction factoraccounting for the type of heater (for a cylindrical heater, ft = 0),fm is the correction factor that accounts for the constructionmaterial (for carbon steel, fm = 0, and for stainless steel, fm = 0.5),and fp is the correction factor for the effect of the operatingpressure (for pressures below 34 bar, fp = 0).

A summary of the capital, operating, and total costs for twoconventional configurations and a DWC is given in Table 9.Although the conventional configuration with two columnsequipped with structured packings appears more cost-effectivethan the conventional configuration employing tray columns, a

compact DWC is by far the most attractive option. Compared tothe conventional tray column configuration, a DWC wouldrequire 46.8% less capital and 51.7% less utilities, which com-bined results in a 50.7% lower total annualized cost (TAC). Thisis really appealing, indicating that implementation of a DWCcould be highly rewarding, i.e., could lead to a substantial increasein profitability of aromatics processing plants.

6. CONCLUDING REMARKS

In this paper we have demonstrated that a commercialsimulator can be used in conjunction with initial guesses forgoverning variables obtained from a simple but theoreticallyfounded short-cut method to generate without computationdifficulties an optimized internal configuration of a DWC.Compared to the conventional two-column-in-series configura-tion for obtaining benzene- and toluene-rich fractions from a 15-component feed, a DWC requires 43.3% less energy to get thesame product specifications.

The Delft model for structured packings in combination withthe Rix andOluji�c method for packed column internals proved tobe an effective and reliable tool for preliminary dimensioning ofDWCs equipped with both conventional and high performancecorrugated sheet structured packings. The free area of liquidcatchers appeared to be a suitable variable for tuning the pressuredrop in the partitioned part of a DWC.

The compact dimensions of a DWC translate into a consider-ably lower installed equipment cost. Expressed in total annual-ized costs (TACs), a DWC enables a 50.7% savings with respectto the conventional two-columns-in-series configuration em-ploying tray columns, and 47.0% savings with respect to thoseemploying the same type and size of structured packings.

The fact that much less energy, capital, and space is neededmakes a DWC a highly interesting option for implementation inaromatics processing plants.

Table 8. Unit Prices of Structured Packing, Packed ColumnInternals, and Sieve Trays

device unit price

sieve trays 600 US $/m2

sructured packing (250 m2/m3) 2000 US $/m3

liquid distributor 4000 US $/m2

liquid collector 2000 US $/m2

support grid 800 US $/m2

Table 9. Equipment Costs, Utility Costs, and Total Annual-ized Cost (TAC) for Conventional Sequence with Respec-tively Trayed and Packed Columns and a DWC (PriceReference December 2009)

configuration

C1�C2 (tray) C1�C2 (packed) DWC (packed)

Installed Equipment Costs (US $)

column shell 1,261,781 781,468 501,621

column internals 678,240 611,793 516,332

reboiler 443,109 401,809 259,461

condenser 399,610 386,898 203,090

total capital 2,782,740 2,181,967 1,480,504

savings (%) � 21.6 46.8

Operating Costs (US $/year)

cooling water 121,834 119,337 59,169

fuel oil/4 bar steam 938,071 907,096 452,491

total utilities 1,059,905 1,026,433 511,660

savings (%) � 3.2 51.7

TAC (US $/year)

1,338,179 1,244,630 659,710

savings (%) � 7.0 50.7



’AUTHOR INFORMATION

Corresponding Author*E-mail: [email protected].

’ACKNOWLEDGMENT

The authors thank J. Montz for providing the referencedesigns for conventional columns and for the correspondingDWC as well as the information on purchased and installed costsof trays, packed column internals, and structured packings. Wewould also like to thank INA Refinery Sisak, Croatia (the MOLgroup), for providing actual plant data.

’NOMENCLATUREA = cross-sectional area (m2)ap = specific geometric surface area of packing, m2/m3

ae = effective (interfacial) area, m2/m3

b = corrugation base length (m)C = installed cost of equipment (US $)ctype = heat exchanger type cost coefficientd = diameter (m)dhG = hydraulic diameter for the gas phase (m)FG = uGs(FG)

0.5 = gas load factor (Pa0.5 or m/s (kg/m3)0.5)FG,lp = loading point gas load factor (m/s (kg/m3)0.5)Fload = loading effect factorf = cost-related correction factorg = gravity acceleration (m/s2)HETP = height equivalent to a theoretical plate (m)h = corrugation height (m)hc,t-t = tangent-to-tangent column height (m)hL = operating liquid holdup (m3 of liquid/m3 of bed)hpb = height of the packed bed (m)hpe = height of the packing element (m)htray = height between top and bottom tray (m)lG,pb = total length of gas flow channel in a packed bed (m)lG,pe = length of gas flow channel in a packing element (m)lpcw = distance from column shell to partition wall on prefrac-

tionator side (m)MG = mass flow rate of gas/vapor (kg/s)ML = mass flow rate of liquid (kg/s)npe = number of packing elements (layers) in a bedp = operating pressure (bar)ΔP = pressure drop (Pa or mbar)Q = reboiler duty (MW)ReGe = effective gas phase Reynolds numberReGrv = relative velocity Reynolds numberReL = Reynolds number for the liquids = corrugation side length (m)TAC = total annualized cost (US $)uGe = effective gas velocity (m/s)uGs = superficial gas velocity (m/s)uLe = effective liquid velocity (m/s)uLs = superficial liquid velocity (m/s)Δz = unit bed height (m)

Greek SymbolsR = corrugation inclination angle (deg)RL = effective liquid flow angle (deg)RDC = flow direction change angle (deg)δ = liquid film thickness (m)ε = packing porosity (m3 of voids/m3 of bed)

ζ = column internals flow resistance coefficientζDC = overall coefficient for direction change lossesζGG = overall coefficient for gas�gas friction lossesζGL = overall coefficient for gas�liquid friction lossesΘ = angle subtended by partition wall (deg)μG = viscosity of gas (Pa s)μL = viscosity of liquid (Pa s)ξbulk = direction change factor for bulk zoneξGG = gas�gas friction factorξGL = gas�liquid friction factorξwall = direction change factor for wall zoneFG = density of gas (kg/m3)FL = density of liquid (kg/m3)j = fraction of the triangular flow channel occupied by liquidj = free or void area of column internalsΨ = fraction of gas flow channels ending at column walls

Subscriptscond = condenserc = columncc = chevron collectorct = chimney tray collectorDC = direction changeGG = gas�gas interactionGL = gas�liquid interactionL = liquidlam = laminar flowm = related to construction materialmc = main columnnt = narrow trough distributoro = overallp = related to operating pressurepc = prefractionator columnreb = reboilershell = related to column shellt = related to fired heater typetrays = related to traysturb = turbulent flow

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