(2000) - bubble-column and airlift photobioreactors for algal culture.pdf

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Bubble-Column and Airl ift Photobioreactors forAlgal CultureAsterio Sanchez Miron, FranciscoGarca Camacho, AntonioContrerasGomez, EmilioMolina Grima,

andYusufChistiD ept. of Chemical E ngineering, U niversity of Almera, E -04071 Almera, Spain

Bubble colum ns and airl ift photobioreactors can be useful for culturi ng phototrophic

organisms requirin g light as a nutri ent. L ight aailability in bubble columns and airlift(deices is influenced by aeration rate, gas holdup, and the liquid elocity mixing and

)turbulence . The photosynthetically generated oxygen also needs to be remoed, as ex-

cessie dissoled oxygen suppresses pho tosynt hesis. O xygen remoal capacity is goernedby the magnitude of the oerall gas liquid m ass-transfer coefficient, k a . T his workL Lcharacterizes th e releant hydrodynami c and mass-transfer parameters in three air-

agitated reactors: bubble column, split-cylinder airlift deice and concentric draft-tube

sparged airli ft essel. The reactors are then ealuated for culture of the mi croalga

Phaeodactylum tricornutum. All reactors were about 0.06 m3 in workingolume, and

the working aspect ratio was about 10. Data were obtained in tap water for a base-line

comparison and in M editerranean seawater, as a potential m edium for algal culture. A

theoretical relationship was deeloped and proed between k a and the aeration rate.L LI n addition, a m ethod based on mechanistic relationships was proed f or predicting the

liquid circulation elocity and k a in ai rlif t reactors. Existing correlations applied sat-L Lisfactorily to gas holdup and k a data obtained in the bubble column. Aqueous solu-L L

( )tion of sodium chloride 0.15 M closely resembled seawater in terms of its hydrody-namics and oxygen transfer behaior. Under the conditi ons tested, all three reactors

attained a biomass concentration of about 4 kg my 3 after260 h. The mean maxi-mum specific growth rate was 0.022 hy 1 in all cases at a power input of 109 W my 3.

Introduction

Airlift and bubble-column bioreactors are simple devicesthat have gained wide acceptance in gasliquid contactingapplications in bioprocessing, the chemical process industry,and treatment o f wa stewater. Substantial knowledge exists onga sl iquid hydrodynamics and mass transfer in bubblecolumns and airlift bioreactors, as comprehensively discussed

in major trea tise Ch isti and M oo-Young, 1987; C histi, 1989,1998, 1999a, b; Deckwer, 1992; Joshi et al., 1990; Merchuk

. and G luz, 1999 . With few exceptions Co ntrera s et al., 1998a;G arca Camacho et al., 1999; Matthijs et al., 1996; Sachez

.Miron et al., 1999; Silva et al., 1987; Suzuki et al., 1995 , ear-lier work with these reactors focused on nonphotot rophic ap-plications. Unlike in conventional bioreactors, light is an es-

Correspondence concerning this article should be addressed to Y. Chisti.

sential nutrient for phototrophic culture and the need forsufficient illumination significantly a ffects the design o f an

outdoor culture fa cility Tredici, 1999; Sa nchez Miron et a l., .1999 . For most commercial processing, outdoor illumination

.sunlight appears to be the o nly viable option.At present bubble columns and a irlift reactors are no t used

as photobioreactors, except for investigational purposes;however, because of the significant potential advantages of

.these systems Sanchez Miron et a l., 1999 relative to conven- .tiona l tubular loop solar ha rvesters Tredici, 1999 , there is a

need to further develop the airlift and bubble-column devicesas photobioreactors. Such systems have already shownpromising performance in outdoor culture of microalgae.D ata suggest that a single vertical tubular photobioreactor .bubble-column o r a irlift design cannot exceed about 0.2 min diameter or light availability will be reduced severely

September 2000 Vol. 46, No. 9 AIChE J ournal1872


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.Sanchez Miron et al., 1999 . In addition, the height of a single device is limited to about 4 m for structural reasons andto reduce mutual shading of reactors in a multicolumn facil-ity that would be necessary for any commercial-scale opera-

.tion Sanchez Miron et al., 1999 . Further restrictions on acceptable aeration rate are posed

by considerations o f shear sensitivity Chisti, 1999b; Contr-.eras et al., 1998a; Silva et al., 1987 and light penetration

.Sanchez Miron et a l., 1999 . A certain minimal aera tion ra te

is essential so that the cells do not stagnate for long in the .dimly lit interior of the reactor Sanchez Miron et al., 1999 .

At the same time, there is an upper limit on the acceptablelevel of turbulence, because hydrodynamic forces affect cer-

.tain algal cells, as reviewed recently Chisti, 1999b . Also, inseawater, excessively high aeration rates generate persistentmicrobubbles that accumulate over time, thus reducing light

penetration over time even at a fixed a eration rate Sanchez.Miron et al., 1999 . In addition to mixing the culture, aera-

tion aids in removing the photosynthetically produced oxygenfrom the broth. Accumulation of oxygen inhibits photosyn-thesis. Similarly, good gasliquid mass transfer is necessaryfor efficient transfer of carbon dioxide that is the carbonsource in photosynthetic cultures.

This article evaluates and compares airlift and bubble-col-umn devices, mainly in terms of hydrodynamics and transportphenomena, in anticipation of a more extensive use of thesesystems in producing microalgae. The focus is on fractionalgas holdup, l iquid circulation velocity, and the overallga sliquid oxygen mass-transfer coefficient and the interrela-tionships among these variables in regimes relevant to algalculture. Data are also reported on a culture of the microalgaPhaeodactylum tricornutum, which is a potential source of cer-tain omega-3 polyunsaturated fatty acids of therapeutic sig-nificance.

Theoretical Developments

This section details the development of a novel theoreticalequation that links the overall volumetric gasliquid mass-transfer coefficient k a with gas holdup and the superficialL Laeration velocity, or the principal operational variable in air-lift and bubble-column reactors. The experimental data arediscussed later in terms of the t heoretically derived eq uation.

In a batch bubble column, the specific gasliquid interfa-cial area a , the overall gas holdup , and the mean bubbleL

.diameter d are related Ca lderbank, 1958; Chisti, 1989 byBthe equation:

6a s . 1 .L

d 1y .B

Equation 1 is based on fundamental principles, as discussed .in depth elsewhere Chisti, 1989 . M ultiplying both sides of

Eq . 1 by the ma ss-transfer coefficient k produces the equa-Ltion

6k Lk a s . 2 .L L

d 1y .B

Substantial experimental evidence affirms tha t

kLsconstant s z, 3 .

dB

irrespective of the flow regime and the type of fluid Chisti.and Moo-Young, 1987; Chisti, 1989, 1998 . In addition, based

.on theory, the gas holdup is necessarily related Chisti, 1989

with the superficial gas velocity U and the mean bubble riseGvelocity U , as follows:b

UGs . 4 .

Ub

Consequently, Eq. 2 can be expressed as

6 zU 6 zUG Gk a s s . 5 .L L U U yUG b G

U 1yb /UbEquation 5 is obtained by substitution of Eq. 3 and Eq. 4 inEq . 2. For a given fluid and flow regime, the mean velocity of

.bubble rise depends Clift et al., 1978 only on the diameterof the bubble,

U s f d . 6 . .b B

For otherwise fixed conditions, the bubble size is controlled .by the specific energy input E in a reactor Chisti, 1989 and,

for a bubble column, we have

kkd A E A gU . 7 . .B G

The exponent k is usually around y0.4 Bhavaraju et al .,.1978; Calderbank, 1958 . Thus, Eq. 5 can be written as fol-

lows:

6 zUGk a s , 8 .L L kcU yUG G

or

6 zk a s . 9 .L L ycU y1

G

The parameter c is close to unity in the bubble flow regime;thus,

k a s , 10 .L L yU y1

G

where s6 z. Eq uation 10 is dimensionally consistent whenthe product cU y is taken to be dimensionless; that is, c ha sGthe units my ysy. The parameters c and y may take othervalues, depending on the fluid and the flow regime.

September 2000 Vol. 46, No. 9AIChE J ournal 1873


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( ) ( ) ( )Figure 1. Reactors: a vessel dimensions and air sparger details; b location of dissolved oxygen DO and pHelectrodes.All dimensions in mm.



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Equation 10 is derived for bubble columns, but a similarrelationship can be shown to hold for airlift bioreactors. Thus,the overall volumetric gasliquid mass-transfer coefficientk a in an airlift device consists of contributions of the riserL L

.and the do wncomer zones Chisti, 1989, 1998 , a s follows:

A k a q A k a . .r L L d L Lr dk a s 11 .L L

A q Ar d

where A an d A are the cross-sectional a reas of the riserr dand the downcomer zones, respectively. The subscripts r an dd denote the riser and the downcomer zones. Even when d

. . .0, k a k a Chisti, 1989, 1998 and , generally,L L d L L r A F A . Consequently, in an airlift device,d r

A k a .r L L rk a f . 12 .L L

A q Ar d

Now, following the logic of Eqs. 29, we obtain

ak a s , 13 .

L L yU y1G

.where s6 zA r A q A . Note that in an airlift reactor,a r r d U is the bubble rise velocity relative to the liquid and notbrelative to wall of reactor. Equations 10 and 13 are used tointerpret the k a data obtained in this work.L L

Materials and Methods

Reactors and fluids

Measurements were made in a bubble column, a split-cyl-inder airlift device, and a concentric draft-tube airlift vesselsparged in the draft tube. All vessels were made of 3.3-mm-

.thick transparent poly methyl methyacrylate , except for thelower 0.25-m sections, which were made of stainless steel .Figure 1 . The vessels were 0.193 m in internal diameter.The riser-to-downcomer cross-sectional area ratio was unityfor the split cylinder and 1.24 for the draft-tube airlift vessel.The internal diameter of the draft tube was 0.144 m. Thedraft tube and the baffle were located 0.091 and 0.096 m fromthe bottoms o f the reactors, respectively. The gas-free liquidheight was about 2 m in all cases. Other geometric details,including those of the air spargers for the various reactors,are noted in Figure 1. Tap water and Mediterranean seawa-ter were the liquids used. The seawater contained a bout 36.6kg my3 total dissolved solids, almost all as inorganic salts.The ionic strength of seawater, estimated using the Langelier

.equation Snoeyink and Jenkins, 1980 , wa s 0.92.In all cases, the specific power input in the reactors was

.calculated Chisti, 1989; C histi and Moo -Yo ung, 1987, 1989using the equation:

PGs gU , 14 .L G

VL

where P is the power input due to aeration, V is the cul-G Lture volume, gis the gravitational a cceleration, a nd U is theG

superficial gas velocity based on the entire cross-sectional areaof the reactor tube. The specific energy input per unit masswas obtained with the equation:

PGEs s gU . 15 .G

VL L

All measurements were at 222C .

Gas holdup

The overall gas holdup in airlift reactors was measured bythe volume expansion method. Inverted manometers wereused to measure the separate holdup values for the riser andthe downcomer zones. The overall holdup in the bubble col-umn was measured manometrically. Both the volume expan-sion a nd t he manometric methods have been described in de-

.tail a nd used widely Chisti, 1989 . The holdup was calcu-lated using the equation:

hms 16 .

ht

where h is the vertical distance between the manometer taps,tand h is the manometer reading. The location o f themmanometer taps is shown in Figure 1 for the various reactors.

Liquid circulation elocity

A small amount of concentrated sulfuric acid was added tothe reactor to reduce the pH to around 4, and the reactor

y1.was bubbled with air U 0.02 m s f or a t le ast 30 m inGprior to measurements. This removed most of the bufferingdue to dissolved carbonates and bicarbonates. The pH was

.raised to around pH 5 by adding sodium hydroxide 12 M . Ameasured amount of concentrated acid tracer was now added

to the reactor instantaneously. Additions were made on thesurface of t he dispersion at the center o f the vessel cross sec-tion. The tra cer signal wa s followed by two identical pH elec-trodes positioned in the downcomer, as shown in Figure 1.The placement of electrodes a ttempted to minimize entranceeffects, but maintained a sufficient vertical distance betweenthe probes so that measurement inaccuracies were mini-mized. The signal from the pH electrodes was expressed as anormalized concentration of the Hq ion and plotted against

.time Figure 2 . The dimensionless normalized concentrat ionwas defined as

w q x w q xH y H oqw xH s , 17 .Normalized q qw x w xH y H o

w q xwhere H is the instantaneous molar concentration a ndsubscripts o an d denote initial and final equilibrium val-ues, respectively.

The tracer response signal displayed the dampened oscilla-tory pattern that is characteristic of a irlift reactors Chisti and

.Moo-Young, 1987; Chisti, 1989 . The mean linear flow veloc-ity in the riserdowncomer loop was calculated from the time

.interval between a djacent tracer peaks Figure 2 of a given



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Figure 2. Normalized tracer concentration profiles fromthe upper and lower pH electrodes.Time intervals used in estimating the mean loop velocity and

the linear velocity in the downcomer are shown.

pH electrode and the geometry-determined length of the cir-

culation path. The linear liquid velocity V in the down-L dcomer was calculated from the t ime interval between the cor-

.responding peaks such as the second peak of the two elec- .trodes Figure 2 and the known vertical distance between

them.The measured linear velocity V in the downcomer wasL d

.related Chisti, 1989 to the superficial velocity in the riser bythe continuity relationship:

U A sV A 1y sV A 1y , 18 . . .L r r L r r r L d d d

where V is the linear liquid velocity in the riser. In theL rsplit-cylinder reactor, because the cross-sectional areas of theriser and downcomer zones were identical, that is, A s A ,

r dEq. 18 simplified to

U sV 1y sV 1y . 19 . . .L r L r r L d d

Because the superficial liquid velocity U in the downcomerL d .is V 1y , U an d U values were identical, irrespectiveL d d L r L d

of the gas holdup; thus, the mean superficial velocity of theriserdowncomer loop was the same as U . For the draft-L rtube reactor, the equivalent of Eq. 19 was

A dU sV 1y sV 1y . 20 . . .L r L r r L d d

A r

For both airlift reactors, the mean linear velocity VLoopthrough the riserdowncomer loop was related to the super-ficial liquid velocity in the riser, as follows:

V qV 1 U UL r L d L r L d V s s q . 21 .Loop /2 2 1y 1y . .r d

For the split-cylinder vessel, because U an d U were iden-L r L d tical, Eq. 21 simplified to

V qV U 1 1L r L d L r V s s q . 22 .Loop /2 2 1y 1y . .r d

Gasli quid mass-tr ansfer coeffi cient

The well-known dynamic gassing-in and gassing-out meth- .ods Chisti, 1989, 1999a were used to measure the k a .L L

Two independent measurements were made simultaneously .using two dissolved oxygen D O electrodes, located as notedin Figure 1. Both probes provided identical k a values,L Lhence confirming the assumed well-mixed liquid phase. Mea-surements were made during absorption and desorption ofoxygen. For absorption, the fluid was deaerated by bubblingwith nitrogen until the DO concentration had declined to be-low 5% of air saturation. The nitrogen flow was then stoppedand the bubbles were allowed to leave the liquid. A presetflow of air was now established, and the increase in DO con-centration was followed with time until the concentrationreached almost 100% of the air saturation value. The k aL Lwas calculated as the slope of the linear equation:

C

yColn s k a ty t , 23 . .L L o /C yC

where C is the saturation concentration of DO, C is theoinitial concentration of DO at time t when a hydrodynamicosteady state has been reestablished upon commencement of

aeration, and Cis the DO concentration at any time t Chisti,.1989, 1999a . Both the absorption and desorption methods

gave identical values of k a under identical hydrodynamicL Lconditions. Only the k a values measured during oxygenL Labsorption are reported here.

Algal culture

Phaeodactylum tricornutum U TEX 640 was the microalgaused. The culture was obtained from the collection of theU niversity o f Texas, Austin. The inoculum for the photo-

bioreactors was grown indoors under artificial light 230y2 y1 .E m s light flux at the vessels surface in a 20-L bub-

ble column. The medium was prepared in seawater as previ- .ously detailed Sanchez Miron et a l., 1999 .

Outdoor cultures were carried out batchwise, simultane-ously in all reactors, during August 516, 1999. The mea noutdoor irradiance during this period was 20069 E my2 sy1 at 8:00 h, rising to a mea n daily value of 1,056278E my2 sy1 at noon. The reactors were located in Almera .36 50 N, 2 27 W , Spain. The biomass concentration atinoculation was about 0.07 g Ly1. The aeration rate duringculture was constant at a U value of 0.011 m sy1, corre-Gsponding to a specific power input of 109 W my3. The tem-perature was controlled at 20C by circulating chilled waterthrough a jacket that surrounded the lower steel portion ofthe reactors. Seawater was added from time to time to makeup the losses. Other aspects o f the culture methodology have

.been detailed previously Sanchez Miron et a l., 1999 .



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Figure 3. Comparison of the measured gas holdup inthe bubble column with the correlations of

( ) ( ) ( )Chisti 1989 : a tapwater; b sea- or saltwa-

ter.

Results and Discussion

Gas holdup

Numerous gas holdup correlations are available for bubblecolumns Akita and Yoshida, 1973; Chisti, 1989; D eckwer,

. 1992 and airlift biorea ctors Akita et a l., 1994; Chisti, 1989,.1998, 1999a; Kawase et al., 1995; Miyahara et al., 1986 . While

there tends to be a general agreement among the differentcorrelations for bubble columns, this consistency is lacking

.for airlift reactors Chisti, 1989, 1998 . In the latter, theholdup is influenced by t he induced liquid circulation rate

that depends on the geometry of the flow path, the gasliquidseparating a bility of the head zone of the reactor Chisti and

.Moo-Young, 1993 , and also the height of the airlift column .Chisti, 1989, 1998 . As shown in Figure 3, the gas holdupdata in the bubble column agreed closely with equations pub-

.lished for tap water and salt solutions Chisti, 1989 , thusconfirming the accuracy of the measurements. In Figure 3,the slight deviation of the data from the correlations for spe-cific power input values greater than about 400 W my3 isbecause of the change in flow regime from bubble flow to

Figure 4. Comparison of the measured riser gas holdupin the draft-tube airlift vessel with published

( ) ( )data: a tap water; b sea- or saltwater.

churn turbulent regime. As expected for media containing alarge amount of dissolved salts, the flow transition occurs at

.slightly greater power input Chisti, 1989 in seawat er com- .pared to tap wa ter Figure 3 . D issolved ions inhibit bubble

coalescence Akita and Yoshida, 1973; Chisti, 1989: D eck-.wer, 1992; Hikita et al., 1981 , hence postponing flow transi-

tion to higher values of gas holdup. As shown in Figure 3, thecorrelation developed for 0.15 M sodium chloride solution .Chisti, 1989 is quite satisfactory f or seawater, because theeffect of dissolved ions on gas holdup is marginal once theionic strength is 0.2 or greater.

In the two airlift reactors, the gas holdup values agreed

less well with published data. Thus, as shown in Figures 4and 5, the riser holdup in tap water was consistently low incomparison with the equation:

1r121r82 3 2 gd gd U r r L r L L r r s U y ,G r4 2 / / / 1y gd'1y . rLrr

24 .



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Figure 5. Comparison of the measuredriser gas holdupin the split-cylinder airlift vessel with pub-

( ) ( )lished data: a tap water; b sea- or saltwa-

ter.

.established by Akita et a l. 1994 . In Eq . 24, the pa rameter is 0.20 for nonelectrolytes and 0.25 for electrolytes Akita and

.Yoshida, 1973 . The riser diameter d in Eq. 24 has no im-rpact on the value of gas holdup. For both a irlift devices,

agreement with Eq . 24 improved greatly in seawat er Figures.4 and 5 . In addition, the riser holdup in tap water did not

correlate well with the equation:

. ..nq2 r 2 nq1 U rn .r G r

s , .. . ..1r 2 nq1 3 nq2 r 4 nq1n1y

g K Ar

d3nq5.r nq1.2 1q

/ / AL r25 .

where, for Newtonian fluids, the flow index n is unity andthe consistency index K is replaced by viscosity. Equation 25

.was developed by Kawa se et al. 1995 using theoretical prin-ciples and assuming isotropic turbulence. Limited usefulnessof Eq. 25 is apparently due to its disregard for effects of liq-

uid circulation o n gas holdup. Also, E q. 25 does not considereffects of surface tension and dissolved ions. Moreover, thetheoretical foundation of the equation is questionable be-cause the assumption of isotropic turbulence in bubblecolumns and airlift reactors is not valid, even at high-energy

inputs in wa terlike media Chisti, 1998; Lubbert and Larson,.1990; Lubbert et al., 1990; Okada et al., 1993 . Although Eq.

25 is intended also for viscous non-Newtonian media, the as-sumption of isotropic turbulence in such systems is even less

realistic.Another equation for riser gas holdup in internal-loop air-

.lift reactors is that of M iyahara et al. 1986 :

'0.4 Fr s , 26 .r UL r'1q0.4 Fr 1q /UG r

where the F roude number Fr is based on the diameter of thesparger hole, that is,

U2G rFrs . 27 .

gdo

Equation 26 was developed for a variety of Newtonian fluids,but not for media containing large quantities of dissolvedsalts. However, a s shown in Figure 6 for tap water only, Eq.26 consistently and substantially overpredicts holdup values.The comparisons in Figures 36 clearly demonstrate that thegas holdup data in different airlift devices are not well corre-lated with equations developed without considering the un-derlying mechanics. Numerous correlations for gas holdup in

q airlift devices have taken the general form s pU ChistiGan d M oo -Young , 1987; Chisti, 1989, 1998; Merchuk a nd G luz,

.1999 , but these equations are suited only to specific combi-nations of fluid properties and reactor geometry, because theparameters p an d qare sensitive to these factors.

A superior approach to correlating gas holdup in airlift re-actors is the use of the drift-flux model-type equations incombination with a mechanistic relationship for liquid circu-

.lation velocity in two-phase flow Chisti, 1989, 1998 . Thedrift-flux relationship for gasliquid flow in vertical conduitstakes the form:

UG r s , 28 .r

U qU q .L r G r

where the parameters and have physical meanings .Chisti, 1989, 1998 . As shown in Figure 7 for two representa-

tive cases including both airlift reactors and the two fluids,the riser gas holdup correlates well with Eq. 28. For the two

. y1fluids Figure 7 , a mean value of 0.30 m s wa s in t heexpected range for bubble rise velocities. The parameter

.differed for the two reactors Figure 7 because the shapes ofthe flow channels were quite different for the two cases. Theriser of the draft-tube airlift had a circular cross section,whereas in the split-cylinder device the riser cross section wasa semicircle. In view of the fit in Figure 7, the induced liquidcirculation rate clearly needs to be taken into account for



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Figure 6. Comparison of the measuredriser gas holdupin the two airlift vessels with Eq. 26 of Miya-hara et al.

predicting gas holdup. In circular channels an -value of unityimplies a flat radial velocity profile. In view of the high -

.value Figure 7 in the split-cylinder device, the radial veloc-ity profile was apparently parabolic in the semicircular chan-nel.

U nfortunately, in an airlift device, the induced liquid circu-lation velocity is not usually known a priori, and therefore forpredictive purposes, the theoretical Eq. 28 needs to be used

.in combination with ot her equa tions Ch isti, 1989, 1998 . Theneed to predict gas holdup arises mainly because of the needto know the gasliquid mass-transfer coefficient that dependson holdup. As shown in the following subsection on the ef-fect of liquid circulation velocity on mass transfer, Eq. 28 in

combination with a well-known mechanistic model fo r the in- .duced liquid circulation velocity Chisti, 1989 , provided a re-

liable method for predicting the overall volumetric mass- .transfer coefficient k a values in airlift reactors.L L

With few exceptions G anzeveld et a l., 1995; Wenge et al.,.1996 , the relationship between the riser and the d owncomer

gas holdups has been generally expressed Contreras et al.,.1998b in the form:

s a , 29 .d r

Figure 7. Plots of Eq. 28 for two representative cases inthe airliftvessels.

.where the parameter a is a constant Chisti, 1989 . However, .as was recently pointed out Contreras et al., 1998b , Eq . 29

disregards the fact that gas holdup in the downcomer re-mains zero until a finite holdup value has been established in

.the riser Wenge et al., 1996 . C onsequently, a better correla-tion between riser and downcomer gas holdups has been pro-

.posed Contrera s et al., 1998b in the form:

s a y b. 30 .d r

The meanings of parameters a and b in Eq. 30 have been .discussed elsewhere Co ntrera s et al., 1998b . Representa tive

data in tap water and seawater are shown in Figure 8 plottedaccording to Eq s. 29 and 30. Clearly, E q. 30 provides a supe-rior fit of the data, all of which displays a distinct nonzerox-intercept. Similar behavior was seen for the draft-tube air-lift device.

In the two airlift reactors, the overall holdup was mea-sured by height displacement, where the holdup values in the

. .riser and downcomer were determined manometri-r dcally. In addition to the direct measurements, a value of theoverall holdup also could be calculated using the measured



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Figure 8. Relationship between riser and downcomergas holdups.Split-cylinder data are shown for tap wat er and seawater

plotted according to Eq. 29 and Eq. 30.

.riser a nd do wncomer gas holdups Chisti, 1989 , a s follows:

A q A r r d d s . 31 .

A q Ar d

Equation 31 is based on geometric reasoning and it is an ex-act relationship for the kinds of airlift reactors used here

Figure 9. Comparison of the measured overall holdupwith values calculated according to Eq. 31 fortwo cases in airlift reactors; diagonals repre-sentan exact agreement.

.Chisti, 1989 . As shown in Figure 9 for two representativecases, excellent agreement is observed between the measuredoverall holdup a nd the values calculated with E q. 31. Thisconfirms internal consistency of manometrically measured gasholdup values in the riser and downcomer while also validat-

ing Eq. 31.The various gas holdup values in the three reactors used

are compared in Figure 10 for the two fluids. G enerally, for agiven specific power input, the riser gas holdup in the twoairlift vessels is comparable to the overall holdup in the bub-ble column; however, the downcomer gas holdup is signifi-

.cantly less than the holdup in the riser Figure 10 . C onse-quently, the overall holdup of the airlift vessels is somewhatlower than in the bubble column. In comparison with the riserholdup, the holdup in the downcomer is much lower at the

.lower power input values than at higher ones Figure 10 .This is because under conditions of low-power input the meanbubble size in the riser is bigger, and bigger bubbles are lessreadily dragged into the downcomer with the flow.

Gasliquid mass transfer

Much more information exists on gasliquid mass transferin bubble columns than exists in a irlift devices Akita and

.Y oshida , 1973; C histi, 1989, 1998, 1999a; D eckwer, 1992 . I naddition, compared to airlift reactors, fewer factors influencek a values in bubble columns a nd, for a given fluid, theL Lk a data obtained in different columns generally compareL L

.well irrespective of the column aspect ratio Chisti, 1989 , so



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Figure 10. Comparison of various gas holdup values inthe three reactors for various values of spe-

( ) ( )cific power input: a tap water; b seawater.

long as the column diameter exceeds about 0.1 m, as was thecase here. Because of this consistency, accuracy of the mass-transfer measurements may be demonstrated by comparingdata obtained in a bubble column with o ther well-establishedreference data before applying the same measurementmethod to airlift reactors, or before new data are interpretedin a novel way. As shown in Figure 11 for the bubble column,the measured k a data agreed remarkably well with someL Lof the well-known correlations in both fluids. The correla-tions used in the comparison were as follows.

.1. That of H ikita et al. 1981 :

y0.2481.76 0.243414.9 g U g G L L G k a sL L 3 / / /U G LL

y0.604L, 32 . / DL L

whereas was 1.0 for tap water. For seawater the -value .was 1.2, as recommended by H ikita et al. Ch isti, 1999a . The

other variables in Eq. 32 are gravitational acceleration g, the

Figure 11. Comparison of the measuredk a in bubbleL L

( )column with the published correlations: a( )tap water; b sea- or saltwater.

.surface tension , the viscosities of the gas and theG .liquid phases, the density of the liquid phase, andL L

the diffusivity D of oxygen in the liquid.L .2. The one recommended by Heijnen and Va nt Riet 1984 :

k a s0.32U0. 7. 33 .L L G

.3. That established by Ch isti 1989 :

0.86PGy4k a s2.3910 . 34 .L L /VL

For the comparison in Figure 11, E q. 34 was expressed interms of the superficial gas velocity.

Having validated the k a data, let us see how the theo-L Lretically developed Eq. 10 and Eq. 13 fare in correlating themeasurements. As shown in Figure 12, for all six combina-tions of reactors and fluids, the k a data correlated excep-L Ltionally well with equations of the same general form as Eq.10. Use of Eq. 10 is preferred to purely empirical relation-ships of the type k a s aUb that have been used commonlyL L G



11/16

Figure 12. Correlation of the measured k a with theL L

superficial aeration velocityU according toG

Eq. 10 or Eq. 13; data are shown for all reac-tor-fluid combinations.

Chisti, 1989, 1998, 1999a; Heijnen and Vant R iet, 1984;.Merchuk and G luz, 1999 . This is particularly so for the air-

lift reactors, because there is no general agreement on the

values of a and b for these reactors, even for a given fluid .and reacto r geomet ry Chisti, 1989, 1998, 1999a . .The parameters or an d y in Eq. 13 depend on thea

reactor a nd the fluid, a s can be seen in Figure 12. For a given .type of reactor, the or value is always greater fora

.seawater than for tap wat er Figure 12 , whereas yis close tounity for all cases. Note that these values apply to bubble-flowregime, which prevailed over most of the operational rangetested. A comparison of the data in Figure 12 revealed that

in the bubble column, the k a in seawater was alwaysL Lw y0.979 .greater than in tap water by a factor of 2.5 U y1 rG

y1.171 .x y1U y1 , or 1.3 at U s0.03 m s , so long as the su-G Gperficial gas velocity exceeded 0.01 m sy1. Similarly, in thedraft-tube airlift device, the sewater k a was always 1520%L Lgreater than in tap water. However, in the split-cylinder de-

. .vice, k a r k a was about 0.85 at a gasL L s e a w a t e r L L t a p w a t e rvelocity of0.01 m sy1, and increased to 0.94 when the ve-locity approached 0.03 m sy1. For tap water, the k a in theL Lsplit cylinder was a marginal 2% less tha n in the bubble col-umn. In contrast, the k a values in the draf t-tube deviceL Lwere typically 1215% lower than in the bubble column. Insea water, the k a values in the split-cylinder device wereL L1530% reduced relative to the bubble column, whereas in

the draft-tube airlift the reduction was 020%, depending onthe gas flow rate. These changes in k a values relative toL Lthose in the bubble column were largely explained by a re-duced gas holdup in the airlift units.

In obtaining the theoretical Eq. 10, the k rd ratio wasL Bassumed to be constant. Although a constancy of the k rdL B

ratio has been validated previously Chisti and Moo -Young,.1987; Chisti, 1989 , direct evidence from the present study

further supported the assumption, as discussed next.

Relationship B etween G as H oldup and M ass-Transfer Coeffi-cient. As expected from E q. 2, plots of k a against 6rL L r . .1y were linear Figure 13 , confirming that the k rdr L B

.ratio that is, the slope was constant for a given fluid. Thus,for seawater, a mean k rd value of 0.042 sy1 was withinL B8% for the two airlift reactors, whereas the value for tap

y1.water 0.056 s was within 11% of the mean for bothreactors. In the bubble column the k rd values were 0.055L Bsy1 and 0.062 sy1 for seawater and tap water, respectively.Clearly, in a given fluid, the k rd ratio was little affected byL Bthe type of gas-agitated reactor used, and this was consistent

.with ear lier observatio ns Ch isti, 1989 . The magnitude o f thek rd values obtained compared well with earlier reportedL B

y1 .ones: for example, a value of 0.05 s Chisti, 1989 was re-ported in 0.15 M sodium chloride, and it was within 20% ofthat for seawater in the present study. The observed k rdL Bvalues agreed with the equation within 10%:

0.52k gD 2L Ly5 y0.131Css5.6310 e . 35 .3 /d B L

I n E q . 3 5 C is the concentration of solids in suspensions .wtrvol % , D is the diffusivity of gas in liquid, a nd is theLinterfacial tension. Equation 35 was developed for air waterdispersions and for suspensions with a waterlike continuous

.phase Chisti a nd M oo-Young, 1987; C histi, 1989 .



12/16

Figure 13. Correlation of the measured k a and theL L

measured riser holdup in airlift reactorsr

according to Eq. 2; vertical bars denote stan-dard deviation of selected mean k a val-L L

ues.

The k rd ratio in seawater was significantly less than inL B .tap wa ter Figure 13 ; thus, for a given bubble diameter the

k value was lower in seawater. This made sense, as k isL Lproportional to D or D , depending on the situation'L L .Chisti, 1989 and, relative t o pure water, dissolved ions re-

duce diffusivity o f o xygen. As previously noted the section.on gasliquid mass transfer; Figure 12 , the k a values inL L

seawater were always greater than in tap water. B ecause thepresence of salts reduced k , an enhancement of k a wa sL L Lexplained by an increase in the specific interfacial a rea a inLthe presence of salts. The increase in a more than compen-Lsated for a decline in the k value.L

Effect of L iquid Circulation Velocity on Mass Transfer.

Measurements of liquid circulation velocity are important in

themselves, but here the discussion is restricted only to theimpact of liquid circulation on gas holdup and the mass-transfer coefficient values. A well-known a nd theoreticallybased model for predicting the induced liquid circulation ve-

.locity in an airlift device has been published Chisti, 1989 as:

2 g y h .r d DU s . 36 .L r 2

K A 1T rq K) B2 2 /A1y 1y . .dr d

In Eq. 36, h is the height of dispersion, and K and K areD T Bthe frictional loss coefficients for the top and the bottom

zones of the airlift loop. Eq uation 36 is based on principles ofenergy conservation, and it has been repeatedly validated fora broad range of scales and configurations of airlift devices .Abashar et al., 1998; Chisti, 1989, 1998 . When K and KT Bare approximately equal, as expected for the reactors in Fig-ure 1, Eq. 36 simplifies to:

2 g y h .r d DU s . 37 .L r 2

1 A 1rq KB) 2 2 /A1y 1y . .dr d

The measured values of the riser a nd downcomer gas holdup

were used in Eq. 37 for predicting the U value. The best fitL rK value, that is, one that produced the closest agreementBbetween predicted and measured U values, was 4.5. AsL rshown in Figure 14, the predicted and the measured valuesof U agreed within 15% for both tap and seawater. A KL r B

.value of 4.6 was calculated Chisti, 1989 using the geometricparameters A and A for the reactor a nd the publishedd bequation:

0.79A d

K s11.4 . 38 .B /A bThe K determined with Eq . 38 and that determined by fit-Bting the U data agreed within 3% of the best fit value ofL r4.5.

As shown in Figure 15 for the bubble-flow regime, the driv-ing force for liquid circulation, that is, the difference betweenthe riser and the downcomer gas holdup values, remainedfairly constant once the specific power input had exceededabout 25 W my3 in the split-cylinder airlift device. Corre-spondingly, the liquid circulation velocity rose sharply withthe gas flow rate, attaining an almost constant value that wasnot particularly sensitive to further increases in the aeration



13/16

Figure 14. Predicted vs. measured superficial liquid ve-locities in the riser of the split-cylinder airliftreactor.

rate. Compared to the split-cylinder airlift device, in the .draft-tube reactor, the y value showed a slight in-r d

.crease as the gas flow rate increased Figure 15 , hence theplateau region of the liquid circulation velocity vs. power in-put curve had a slight positive slope over the entire range of

.aeration rates tested Figure 15 . Compared to the split-cylin-der device, a nd despite a greater circulation driving force,the U value was lower in the draft-tube reactor, a nd thisL rreduced the ability of the liquid to drag bubbles into the an-nular downcomer. The relatively lower U value in theL rdraft-tube reactor was explained by a bigger A rA ratio forr dthat reactor and the greater resistance of its circulatory chan-

.nel that is, a higher K relative to the split cylinder , asBexpected from Eq. 37.

Having validated Eq. 37, let us see how it can be used topredict the mass-transfer coefficient, k a . For predicting theL Lk a values, Eqs. 37 and 28 are solved simultan eously to cal-L Lculate the riser holdup and the U value. Equation 30 isL rthen used to calculate the downcomer gas holdup. The ear-

.lier determined k rd values Figure 13 are then used toL Bcalculate the k a . Figure 16 compares the predicted andL Lthe directly measured values of k a , revealing a remarkablyL Lgood agreement. The agreement in Figure 16 lends addi-tional support to the various mechanistic equations used inthe predicting and confirms the k a prediction methodol-L Logy used.

The k a values in all reactors were such that a ccumula-L Ltion of photosynthetically generated oxygen did not occur

during culture, even at a relatively low aeration velocity ofy1 y3 .0.011 m s 110-W m -specific power input , as shown

in Figure 17. The DO concentration in Figure 17 followed acyclic pattern in all reactors because the oxygen generationrate increased from dawn t o solar noon a s the irradiance levelpeaked. Oxygen generation rate then declined through theafternoon and night. As shown in Figure 17, the DO concen-tration remained at F100% of air saturation, except on twooccasions when the concentration rose to 110% of a ir satu-ration value in the bubble column. In contrast, in conven-

Figure 15. Variation of the riser and downcomer gasholdup values with the specific power inputin the airliftreactors.

tional tubular loop photobioreactors for algal culture, the

concentration of DO commonly reaches G400% of air satu- .ration value Sanchez M iron et al., 1999; Tredici, 1999 . D O

concentrations exceeding a bout 120% of a ir saturation areknown to inhibit photosynthesis and otherwise damage theculture.

The data in Figure 17 confirm the existence of oxygenmass-transfer limitation, at the a eration rate used, for all

.those occasions that is, 4 of 12 days in the bubble columnwhen the D O concentration exceeded 100%of the air satura-tion value. In airlift and bubble-column photobioreactors, low



14/16

Figure 16. Predicted vs. measured k a values in theL L

split-cylinder airlift reactor.

aeration rates corresponding to a power input of 120 Wmy3 are necessary to reduce the formation of stable mi-

crobubbles that reduce light penetration Sanchez Miron et .al., 1999 . However, low aeration rates reduce radial mixing.

This effect can be countered by using larger sparger holes to .increase the size of bubbles Sanchez Miron et al., 1999 .

Larger bubbles tend to improve mixing. Also, cyclic variationof aeration rate may be necessary for attaining the right bal-

ance between the conflicting demands of low gas holdup light.penetration , good mass transfer, and mixing.

Cultur e perf ormance

The oxygen evolution and removal behavior during culturehas already been discussed in the previous section Figure

.17 . The biomass growth profiles are shown in Figure 18 forthe three reactors. In batch culture, at least at the low aera-

Figure 17. Changes in DO concentration during cultureofP . t r i c o r n u t u m in the three bioreactors.

Figure 18. Outdoor batch culture profiles of P . t r i c o r n u -tu m in the three bioreactors during August516, 1999.

.tion rate used Figure 18 , there was no significant d ifference

in culture performance in the three bioreactors. This obser-vation was consistent with the fairly similar values of gasholdup and the k a measured earlier in the reactors. In allL Lcases, the mean value of the maximum specific growth ratewas 0.022 hy1, which is high for P. tricornutum. Note that thisvalue is a mean for the entire culture duration. Within anyone daylight period, the biomass concentration increasedrapidly, as shown in F igure 18; however, t here wa s some lossof biomass during the night because a portion of the intracel-lular stored carbohydrate was consumed by respiration. Thebiomass concentration rose again during the next light pe-riod, attaining a higher value than in the previous light pe-riod. The nighttime decline could be avoided if sufficientlyintense artificial illumination was provided, but this option is

not practicable.

Concluding Remarks

In view of the findings discussed, the principal conclusionsare as follows:

1. The gas holdup and k a in bubble column are sat isfac-L Ltorily predicted with the available correlations. In contrast,existing nonmechanistic correlations perform poorly in pre-dicting the behavior of airlift bioreactors. Hydrodynamic andoxygen-transfer characteristics of seawater are essentially

.equivalent to those of a queous sodium chloride 0.15 M .2. Equations 10 and 13, developed through an analysis of

the underlying fundamentals, describe exceptionally well therelationship between k a and t he superficial gas velocity inL Lall reactors. The two parameters in these equations dependon the properties of the fluid and the reactor.

3. In airlift reactors, the gas holdup, and phase velocitydata are consistent with the drift-flux model equation. Therelationship among the induced liquid circulation velocity, thereactor geometry, and the gas holdup difference d riving forcefor liquid circulation is well described by the theoretically de-veloped Eq. 37.



15/16

4. In airlift vessels, the interdependence of riser and down-comer gas holdup values is best expressed as Eq. 30, ratherthan the often used Eq. 29.

5. The ratio of true mass-transfer coefficient to bubble di-ameter, that is, k rd , is constant for a given fluid reactorL Bcombination. C onstancy of the k rd ratio, in combinationL Bwith E qs. 28, 30, a nd 37, allow a good prediction of the k aL Lvalue.

6. Culture studies confirm that sufficient oxygen is re-moved and a significant accumulation is prevented even whenthe aeration power input is around 110 W my3.

7. Performance of the three types of reactors was equiva-lent under the conditions tested. In all cases, a maximumspecific growth rate of 0.022 hy1 was achieved in batch cul-ture and the final P. tricornutum biomass concentration at-tained was high at 4kg my3.

Pneumatically agitated bubble columns and airlift devicesclearly attain the requisite value of k a and the inducedL Lliquid circulation velocity at a relatively low power input y3.F120 W m for practicable culture of microalgae. So far,the data do not suggest a clear preference for one type ofreactor over another of the three kinds evaluated. A more

exhaustive assessment is necessary o ver a range of cultureconditions. Further studies are underway with P. tricornutumin batch and continuous culture a t various dilution rates, a er-

.ation conditions, and levels of outdoo r illumination seasons .

Acknowledgments

.This work was supported by C ICY T BI O-98-0522 , Spain, a nd the .European C ommission Project BR PR CT97-0537 .

Notation

A scross-sectional area for flow under the baffle or draft tube,bm2

d ssparger hole diameter, moqsexponent

t sinitial or start time, soU ssuperficial gas velocity in the riser zone, m sy1G r

sfractional gas holdup in the riserr sviscosity of gas, P a sG sviscosity of liquid, Pa sL sdensity of the liquid, kg my3Lsinterfacial tension, N my1

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M anuscript receied No. 8, 1999, and reision receied Apr. 13, 2000.


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