A Novel Three-Phase Airlift Reactor WithoutCirculation of SolidsGiuseppe Olivieri, Antonio Marzocchella* and Piero Salatino

Dipartimento di Ingegneria Chimica, Universita degli Studi di Napoli “Federico II”, P.le V. Tecchio, 80, 80125 Napoli, Italy

A novel configuration of a three-phase internal loop airlift reactor is proposed. The draft is divided in two sections: (top) gas–liquid contact section;(bottom) liquid–solid contact section. Solids particles are fluidized in the bottom section by liquid circulation. The main advantage compared withclassic airlift or three-phase fluidized bed is the reduced stress on particles as interference with bubbles is prevented. Experiments with silica sand(325 �m diameter) were carried out to characterise the hydrodynamics of the airlift. The influence of superficial gas velocity, overall solids hold-upand sparger height were assessed. A theoretical analysis was proposed to derive simple design criteria.

Keywords: airlift, fluidized bed, gas–liquid–solids, hydrodynamic regimes, momentum balance

INTRODUCTION

The onset of exceedingly large particle stresses (e.g., impactor friction) may represent a critical issue in equipment forprocessing granular solids (bioreactors, catalytic reactors,

adsorbers, etc.). As an example, multiphase contactors associatedwith biological systems—for example, the fermentation of ani-mal cells or sensitive microorganisms, bioconversion by enzymesor microorganisms immobilised on carriers—require low shearstress as a key feature for successful application of the process(Chisti, 1999; Heltzer, 2000; Kieran et al., 2000; Yim and Shamlou,2000). Membrane reactors and inverse fluidized airlift reactors arefrequently used to overcome the problem of excessive shear in theparticle phase in conventional solid-fluid contactors.

Airlift reactors are characterised by effective mixing, excellentgas–liquid mass transfer and reduced shear rate, when comparedwith traditional mechanically stirred-tank fermenters (Chisti,1989; Merchuk and Gluz, 1999). However significant mechani-cal stresses may still be effect on the particle phase due to thepresence of a bubble phase, in addition to inherent turbulence inthe fluid phase. Bubble-induced stresses may be related to bub-ble formation at the gas sparger, coalescence and break-up, waketurbulence and bursting at the surface of the fluid (Heltzer, 2000;Kieran et al., 2000). The inverse fluidization airlift was success-fully proposed by Nikolov and Karamanev (1987); Garnier et al.(1990); and Lele and Joshi (1992) as a possible solution to fur-ther reduce stresses acting on the particles, but its application islimited to positively buoyant particles.

Negatively buoyant particles are frequently used as carriers forcell/enzyme immobilisation. In a three-phase internal loop airliftwith the sparger located in the draft tube Olivieri et al. (2003,2004) recognised the existence of three hydrodynamic regimes.These differed for what concerns the behaviour of the solid phaseat the bottom of the apparatus as the gas superficial velocity is var-ied: fixed bed, fluidized bed, circulating bed regimes. The range ofUG within which the fluidized bed regime establishes is typicallyupperly limited by the bed transport conditions, usually very closeto the incipient fluidization conditions. Guo et al. (1997) operatedan external loop configuration for different solid materials in thesame regime.

In the present study a novel configuration of a fluidized bedoperated as an internal loop airlift is presented. The main featureof the new configuration is the splitting of the draft in two regionsseparated by the gas sparger: (I) a lower section where the solidsare confined; (II) an upper section where gas is sparged. A netbeneath the gas sparger prevents solids entrainment to the uppersection. The liquid flow induced by the gas sparged in the upperdraft section promotes solids fluidization in the lower region. The

∗Author to whom correspondence may be addressed.E-mail address: antonio.marzocchella@unina.itCan. J. Chem. Eng. 88:574–578, 2010© 2010 Canadian Society for Chemical EngineeringDOI 10.1002/cjce.20314Published online 5 May 2010 in Wiley InterScience(www.interscience.wiley.com)

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Figure 1. The experimental apparatus. DPT, differential pressuretransducer; RPT, relative pressure transducer; DOT, dissolved oxygenprobe; Pt 100, thermoresistance. Inset: the gas sparger.

oxygen transport to the solids (or the conveying of other dis-solved gas) is indirectly actuated by the liquid, after saturationin the upper airlift section. Thanks to this configuration, bubble-induced mechanical stresses acting on the particles are entirelysuppressed.

The hydrodynamics of a three-phase airlift featuring this novelconfiguration has been characterised by monitoring pressure, liq-uid velocity and local hold-up of both gas and solid phases duringoperation. A theoretical analysis of the system has been proposedto support the interpretation of results and to provide designguidelines.

EXPERIMENTAL APPARATUSThe experimental apparatus consists of a lab-scale internal loopairlift (ILA) equipped with a gas flow controller, a gas-humidifier,an internal heater, process diagnostics and a data acquisition unit.The ILA (Figure 1) has already been described by Olivieri et al.(2003). It consists of a 120 mm ID, 2.0 m high cylindrical col-umn made of Plexiglas equipped with a coaxial 70/80 mm ID/OD,0.90 m high (HD) Plexiglas draft tube. The main feature of the newconfiguration is the splitting of the draft into two sections:

(I) Section I, extending from the base of the draft to the gassparger located at a level HI above the bottom.

(II) Section II, extending from the sparger to the top of the draftand whose height is HII.

Gas was distributed in the draft by means of a cross-shapedsparger (inset of Figure 1) equipped with 40 horizontal holes(0.5 mm diameter) located along the four horizontal branches(4/6 mm ID/OD pipes). The sparger level was changed in thepresent study to investigate the effect of the parameter ˛ =HII/HD

(in the range 0.25–0.67) on the performance of the ILA. Thebottom of the draft was equipped with a 10 mm thick plate per-forated with 109 holes of 0.5 mm diameter on triangular pitch.

A 100 �m mesh net was located across the draft just underneaththe sparger, to prevent entrainment of solids from Section I toSection II.

The cylindrical column and the draft tube were equipped withseveral taps for the measurement of pressure, dissolved oxygenand temperature. The annulus was also equipped with a portto inject the liquid tracer when measuring the liquid circulationvelocity.

The plant was operated batchwise with respect to the liquidand solid phases and continuously with respect to the gas phase.Bidistilled water was used as liquid phase. The air stream fed tothe ILA was water-saturated prior to feeding in order to limit liquidevaporation from the equipment. Silica sand (325 �m diameter)was used as bed solids.

Pressure was measured along the draft and the annulusby means of broad-bandwidth relative pressure transducers(RPT—Druck). Differential pressure transducers (DPT—ABB)were used to measure local gas and solids hold-up and frictionloss along the loop. In particular, the pressure drop was mea-sured across the perforated plate distributor and the net. Pressuredrops associated with liquid flow reversal at the top and bottomof the airlift were also recorded. Signals from the pressure trans-ducers were simultaneously recorded for 300 s at 500 Hz on aPC equipped with a data acquisition card and low-pass analog-ical filters (150 Hz cut-off). A Labview® data acquisition/analysissoftware was purposely developed.

The liquid circulation velocity (UL) was measured as the dropvelocity of an immiscible fluid (diphenyl-methane, 1001 kg/m3)in the annulus. Gas–liquid mass transfer coefficient measurementwas carried out adopting the gassing-out technique (Olivieri et al.,2003).

Temperature was set at 30◦C by means of a heat exchangerlocated in the annulus. The bottom clearance was set at 15 mm.Water (10.5 L) were loaded into the reactor, corresponding to atop clearance of 20 mm—liquid height in the airlift 92 cm—undergas-free conditions. The mass of solids loaded in the lower draftsection was 520 g, corresponding to an overall draft volume frac-tion εD

S = 0.058 and, as settled bed, to a draft fraction HBed/HD of0.11.

Once εDS and ˛ were fixed, the gas superficial velocity (UG based

on the cross-sectional area of the column) was quasi-steadilyincreased from zero up to a maximum value that was limitedby one of the following constraints: (I) the maximum level HBed

of the bed of fluidized solids which could not exceed HI; (II) themaximum admissible over-pressure in the plant upstream of thegas sparger. Steadiness of operation was continuously monitoredby recording the pressure at the bottom of the draft with a chartrecorder.

Under each set of operating conditions the behaviour of the ILAwas characterised by the following techniques: (1) visual inspec-tion of the system, including continuous recording of the liquidfree-surface at the column top and fluidized bed levels; (2) timeresolved recording of the pressure profile along the loop. Time-series of measured variables were eventually worked out to yieldstatistical and spectral parameters. In particular, the average andthe variance of the time-series were selected in the present workto mark singularities associated with fluidization regime transi-tions. Axial pressure drop (�p) was rescaled with respect to thegas-free value (�p0).

�p = �p − �p0

�p0(1)

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Under fluidized conditions and assuming the contribution ofwall friction to the pressure drop negligible, the rescaled pressuredrop is related to the local solids (εS) and gas (εG) hold-up inagreement with the relationship (Fan, 1989):

�p = εS�S − εG (2)

where �S = (S − L)/L is the specific gravity of the solids. Thelocal solids hold-up was also characterised by visual inspection ofthe bed height (HBed). The overall gas-holdup in the draft (εD

G) wascharacterised by visual inspection of the expansion of the liquidfree-surface level at the column top.

Liquid velocity and gas–liquid mass transfer coefficient wereestimated under selected operating conditions.

THEORETICAL FRAMEWORKThe momentum balance extended to the modified ILA has beenwritten, based on the following assumptions:

• The momentum contribution of the bubble injection was neg-ligible as a consequence of the sparger design.

• Solids were not entrained in Section II and gas bubbles werenot entrained in the annulus section.

• A dissipative term �pf was introduced to cumulatively accountfor friction loss and for localised pressure drops associated withflow inversion, plate distributor and net resistance, respectively.

In agreement with the above assumptions, the pressure balancereads:

�pA = �pI + �pII + �pf (3)

where �pA is the pressure drop along the annulus. Under fluidiza-tion conditions the pressure drop can be expressed in terms of thedensity of the multiphase system (Fan, 1989):

Lg(HI + HII) = [L(1 − εI

S) + SεIS

]gHBed + Lg(HI − HBed)

+ [L(1 − εII

G) + GεIIG

]gHII + �pf (4)

Working out Equation (4) and neglecting the gas phasecontribution in Section II, the momentum balance yields thedimensionless relationship:

εIIG˛ = �SεI

S

HBed

HD+ �pf

LgHD(5)

Equation (5) highlights that the driving force of liquid circula-tion due to gas holdup balances the weight of the bed—correctedto account for buoyancy—and the friction loss. In the presentstudy, the dimensionless friction loss term has been expressedas a fraction of the driving force:

F = �pf /LgHD

εIIG˛

(6)

RESULTS AND DISCUSSIONFigure 2 reports a selection of data measured during a run car-ried out at εD

S = 0.04 and ˛ = 0.67. �p is the rescaled value of themeasured pressure drop (Equation 1) over a 45 mm high bed ofsolids. The analysis of Figure 2 and the phenomenology observed

Figure 2. Gas and solids holdup, rescaled pressure gradients andfrictional loss as a function of the superficial gas velocity. εD

S = 0.04,˛ = 0.67

during the run suggest the existence of two distinct hydrody-namic regimes for the solids bed, with a transition occurring atUG = 0.8 cm/s and εII

G = 0.14.

• UG < 0.8 cm/s: �p and εIIG increase linearly with UG. The solid

bed in Section I appeared not fluidized as indicated by the con-stant solids holdup εS (calculated fromHBed). The homogeneousgas bubbling regime was observed in Section II, as well asnegligible bubble coalescence.

• UG > 0.8 cm/s: �p and εS decrease with UG approaching valuesthat agree well with results of Equation (2) with a known valueof the ratio (S − L)/L. Under these conditions, the bed ofsolids was fluidized, as indicated by the finding that the bedexpanded and, correspondingly, εS decreased as UG increased.In Section II bubble coalescence was observed for UG > 1.5 cm/sup to 5.68 cm/s when turbulent-heterogeneous regime set in asindicated by the further increase of εII

G. At UG = 9 cm/s the bedexpands up to the net approaching a minimum εS = 0.23. Nobubble entrainment in the annulus was observed, whatever UG.

The liquid circulation velocity measured under the operatingconditions investigated was always smaller than 10 mm/s withan error of about 30%. At these very low values of UL gas bubbleswere not entrained in the annulus, as expected. At the onset ofsolids fluidization, corresponding to UG = 0.98 cm/s, the liquidvelocity in Section I was 1.0 ± 0.3 mm/s, in agreement with theminimum fluidization velocity estimated by Fan (1989) for theliquid–solids system investigated (1.2 mm/s).

Figure 2 reports also the dimensionless frictional loss F esti-mated in agreement with Equation (6) as a function of UG. Thevalue of F is always in the range 0.2–0.4 and increases as UG

increases in the fluidized state. It is interesting to note that thepresence of such a high value of �pf—likely dominated by thepressure drop at the liquid plate distributor—stabilises the flu-idization regime over a broad interval ofUG. In fact the fluidizationregime was observed also in other works (Petrovic et al., 1993;Olivieri et al., 2003, 2004) but over very limited ranges of UG

and without plate distributors at the bottom of the draft tube.The fluidized bed regime was typically upperly limited by thestabilisation of the circulation regime.

Figure 3 reports selected variables reported as a function of ˛ fortests carried out at εD

S = 0.04. The selected variables are the super-ficial gas velocity (UG,mf) and the gas hold-up at the onset of solidsfluidization. Both UG,mf and εII

G,mf decrease with ˛. On the contrary

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Figure 3. Superficial gas velocity and gas-holdup at incipient fluidizationas a function of sparger height. εD

S = 0.04.

the overall draft gas holdup εDG,mf was approximately constant with

˛. To explain this result two features of the balance expressedby Equation (5) should be considered: (I) the dissipative termis constant, being the liquid circulation velocity always equal tothe value corresponding to the incipient fluidization of solids; (II)the term εI

SHBed is constant, as it expresses the—constant—solidsloading in Section I. Accordingly, the term εII

G˛ (or εDG), which rep-

resents the driving force for liquid circulation as the bed of solidsis incipiently fluidized, is constant. On the other hand, the valueof εII

G at the onset of the fluidized state of the solids was estimatedby means of Equation (5) (assuming F= 0.2, see Figure 2) andplotted in Figure 3 as a function of ˛. The agreement between themeasured and the estimated values εII

G is satisfactory.The gas–liquid mass transfer coefficient was measured in Sec-

tion II at incipient fluidization conditions. It was approximately80 h−1, regardless of the value of ˛ in the range investigated.

Equation (5) suggests a criterion to calculate the minimum gasholdup to be established in Section II for the incipient fluidizationof the maximum solids loading in Section I, under pre-set operat-ing conditions (˛, S, and εD

S ). Accordingly, theHI is the theoreticalmaximum bed height under incipient conditions (HBed

mf ).Figure 4 shows plots of the gas hold-up εII

G,mf as a function of theoverall draft solids holdup εD

S = εIS(1 − ˛) according to Equation

(5) for different values of the frictional factor F. The plot was

Figure 4. Gas-holdup at incipient fluidization as a function of the overallsolids loading, for different extent of frictional losses.

based on the assumption that the lower section of the draft tubeis entirely occupied by solids at incipient fluidization with a valueof εI

S ≈ 0.55 and �S = 1.6. For a fixed solids fraction, gas hold-upat the onset of fluidization increases with the frictional loss, asexpected.

The map εIIG,mf versus εD

S,mf reported in Figure 4 may be usedas a criterion to determine the maximum loading of bed solidswhich can effectively be operated in Section I in the fluidizedstate. At εII

G,mf = 0.35 and neglecting frictional losses (i.e., F= 0),the theoretical maximum εD

S is εDS,mf = 0.16, which corresponds to

HBed/HD = (1 − ˛) = 0.27.

CONCLUSIONSA novel configuration of a three phase internal loop airlift was pro-posed, characterised by a gas–liquid contact section at the top anda liquid–solid contact section at the bottom of the draft. Solids par-ticles can be fluidized by liquid circulation induced in the ILA bygas holdup established in the upper section. The hydrodynamicsof the ILA was characterised in terms of gas-holdup, bed voidage,pressure drop throughout the loop as a function of the superficialgas velocity. The influence of the level at which the sparger islocated along the draft was clarified. The theoretical analysis pro-vides a criterion to estimate the gas holdup which is required tofluidize the bed of solids as a function of the height of the uppersection and of the solids density. The overall solids holdup andgas–liquid mass transfer rate are comparable with those typicalof conventional airlift or three-phase fluidized bed designs.

The main advantage of this concept is that it prevents any con-tact between bubbles and solids. This feature may be desirablewhenever the reactor is operated with biological phases that maybe damaged by bubble-induced shear or interfacial stresses. Oxy-gen for the biological phase uptake in the lower section is providedanyway by liquid aeration in the upper section, where gas–liquidcontact is nearly as effective as in classical airlifts or three-phasefluidized beds. The main drawback of the proposed concept isrepresented by the longer liquid circulation times, a feature thatshould not be relevant for applications involving animal or veg-etable cells, that is, biological systems that are mostly affected bybubble-induced stresses, in view of the relatively long metabolictime-scales.

NOMENCLATURE˛ volumetric ratio between Section II and draftε phase holdupF friction loss to driving force ratiog gravity (m/s2)� specific gravityH height (m)p pressure (kPa)p pressure reduced with respect to the gas-free hydrostatic

value density (kg/m3)U superficial velocity (cm/s)

Subscriptsf frictionG gas phaseL liquid phaseS solids phasemf minimum fluidization

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Superscripts0 gas-free conditionA annulusD draftI Section III Section II

ACKNOWLEDGEMENTSThe support of Mr. Antonio De Simone in experimental investiga-tion is gratefully acknowledged.

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Manuscript received November 23, 2009; revised manuscriptreceived January 18, 2010; accepted for publication January 26,2010.

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