Design and Scale-up of Pressurized Fluid Extractors for Food and Bioproducts

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Review

Design and scale-up of pressurized ﬂuid extractors for food and bioproducts

C. Pronyk, G. Mazza

*

Paciﬁc Agri-Food Research Centre, Agriculture and Agri-Food Canada, 4200 Hwy 97, Summerland, BC, Canada V0H 1Z0

a r t i c l e

i n f o

Article history: Received 2 March 2009

Received in revised form 27 May 2009 Accepted 2 June 2009

Available online 6 June 2009

Keywords: Mass transfer Supercritical ﬂuids Subcritical water Pressurized solvent Fixed bed Extraction Leaching Counter-current Bioproducts

a b s t r a c t

This article provides an in-depth review of the literature on the design, scale-up, and effects of scale on the extraction of food and bioproducts in pressurized fluid extractors. The design of pressurized fluid extraction systems such as supercritical CO2, pressurized solvent, and pressurized low polarity water (subcritical water) are similar. Knowledge of phase equilibria, mass transfer rate, and solubility data are important first steps for the scale-up of extraction processes and equipment. The literature for the design, scale-up, and effects of scale on the extraction of bioproducts in pressurized fluid extractors is examined with particular attention to the mass transfer principal and important parameters for extrac-tion as they relate to the design and scale-up of fixed bed pressurized fluid extractors. Often when two scales of an extractor are examined, the scale-up has not been done uniformly, leaving the effects of the scale-up on extraction in doubt. There has been some success in design of a continuous pressurized fluid extractor by utilizing a battery of vessels in series to operate on a quasi-continuous basis, and with the use of screw conveyors to produce a gas-tight plug of material, which allows the extraction to operate at the necessary elevated pressures and temperatures.

Contents

1. Introduction . . . 216

2. Pressurized fluid extraction . . . 216

2.1. Types of pressurized fluid extraction systems. . . 216

2.2. Basic design of pressurized fluid extractors . . . 216

3. Mass transfer and phase equilibria . . . 218

3.1. Modeling of pressurized fluid extraction . . . 218

3.2. General model for the mass balance in a fixed bed extractor . . . 218

3.3. External mass transfer . . . 219

3.4. Internal mass transfer . . . 219

3.5. Effects of processing parameters on the extraction process . . . 219

3.5.1. Solubility . . . 220

3.5.2. Flow rate . . . 220

3.5.3. Material properties . . . 220

3.5.4. Length to diameter ratio . . . 220

4. Scale-up of pressurized fluid extraction . . . 221

5. Continuous pressurized fluid extraction . . . 222

5.1. Counter-current liquid–liquid pressurized extraction. . . 222

5.2. Quasi-continuous solid–liquid pressurized extraction . . . 223

5.3. Continuous counter-current solid–liquid pressurized extraction . . . 224

6. Summary . . . 225

References . . . 225

* Corresponding author. Tel.: +1 250 494 6376; fax: +1 250 494 0755. E-mail address:Giuseppe.Mazza@agr.gc.ca(G. Mazza).

Contents lists available atScienceDirect

Journal of Food Engineering

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1. Introduction

Given enough time and money almost any process or equip-ment can be made to work no matter how poor the design (Ritcey, 2004). The design of industrial-scale equipment is usually pre-ceded by laboratory (bench) and pilot-scale systems. Sometimes the pilot-scale system is skipped, and work goes straight from the laboratory to industrial production. It is thought that time and money from the pilot-scale work can be applied to making the industrial plant work. This is most useful when there is suffi-cient data and the process is similar to one already in existence (Ritcey, 2004). However, with quality data and determination of scale-up factors, the design of pilot or industrial sized equipment is much more efficient. There is a large body of work in the litera-ture dealing with the leaching (dissolution and removal of soluble components from a solid matrix)/extraction of bioproducts from solid organic materials with liquid solvents. Most of the work has been done in the laboratory with bench-scale equipment. There are general texts on the design of pressurized vessels or extraction systems (Bertucco and Vetter, 2001; King and Bott, 1993; Tzia and Liadakis, 2003), but there is little discussion on the aspects of scal-ing up a process or equipment. Such texts generally do not include any discussions on the performance of the equipment or on the appropriate scaling of the extraction process. In this paper the de-sign, scale-up, and effects of scale on the extraction of bioproducts in pressurized fluid extractors will be detailed by examining re-sults available from the literature and patents. The mass transfer principles and important parameters for extraction will be dis-cussed as they relate to the design and scale-up of fixed bed pres-surized fluid extractors.

2. Pressurized ﬂuid extraction

2.1. Types of pressurized ﬂuid extraction systems

Solid–liquid extraction is a separation process involving the transfer of solutes from a solid matrix to a solvent. Solvents are chosen based on solubility characteristics of the desired solute. Ideally to achieve as pure a substance as possible, the solute should have high solubility in the solvent while other components in the solid matrix should not. Economics and safety are always a consid-eration and indeed, safer and less harmful solvents that are easy to remove, or recover, are gaining in popularity. There has been much interest in the field of pressurized fluids, as seen with the growing work with supercritical fluid extraction (SFE) with CO2(

Díaz-Rein-oso et al., 2006; Herrero et al., 2006; Reverchon and De Marco,

2006; Shi et al., 2007) and the use of high pressure solvents includ-ing pressurized low polarity water (PLPW) (Kim and Mazza, 2009; Cacace and Mazza, 2007, 2006; Carabias-Martínez et al., 2005; Mazza and Cacace, 2005; Kaufmann and Christen, 2002; Smith, 2002). The versatility of pressurized solvents is excellent due to the physicochemical properties of the solvent, including density, diffusivity, viscosity, and dielectric constant, which can be con-trolled by varying the pressure and temperature of the extraction system. In this way the solvating power and selectivity of the sol-vent can be effectively controlled.

Three methods of pressurized fluid extraction have gained pop-ularity in the research community as a means of extracting bio-products. They are: supercritical fluid extraction (SFE); pressurized solvent extraction (PSE, also known as pressurized li-quid extraction (PLE), subcritical solvent extraction (SSE), or accel-erated solvent extraction (ASE, Dionex trade name); and pressurized low polarity water (PLPW) extraction (also known as superheated water, subcritical water, pressurized hot water). All of these methods make use of vastly different pressurized solvents, each with a unique set of properties that influence the extraction of the desired compounds. Supercritical fluid extraction utilizes a sol-vent, usually CO2, although other substances such as water or

eth-ylene can be used, near its thermodynamic critical point, where it possesses the density of a fluid with the viscosity and diffusivity of a gas. By controlling the pressure of the supercritical fluid, the sol-ubility within the supercritical fluid can be changed. This permits a high degree of selectivity, which allows for the fractionation of the extract by changing the pressure in the system’s separators. Pres-surized solvent extraction utilizes solvents such as hexane, metha-nol, or ethanol at temperatures above their boiling point, under high pressures, in order to increase the efficiency of the extraction process with respect to extraction time, solvent consumption, and extraction yields. This is accomplished through improved solubility and mass transfer effects, in addition to a disruption of surface equilibria, such as reduced viscosity and increased solvent penetra-tion into the sample matrix (Richter et al., 1996). PLPW extraction is a specific case of PSE, where water is utilized as the solvent. Water is a polar solvent, but if heated from 25 °C to 200 °C under pressure to maintain a liquid state, its dielectric constant decreases from 79 to 35, reaching values similar to solvents such as ethanol (24) or methanol (33) (Cacace and Mazza, 2006).

2.2. Basic design of pressurized ﬂuid extractors

The solvents used for all three pressurized ﬂuid systems are vastly different, yet the basic design of the extraction equipment Nomenclature

ap speciﬁc surface area of a solid substrate particle (m1)

Cf concentration of solute in the solvent (kg/m3)

Cp concentration of dissolve solute in the solvent contained

in the particle pores (kg/m3₎

Cs concentration of the solute in the particle (kg/m3)

Cpo initial concentration of dissolved solute within particle

pores (kg/m3)

Cps concentration of solute on external surface of the

particle (kg/m3)

Cso initial concentration of solute in the particle or adsorbed

onto the particle surface (kg/m3₎

dE diameter of the extraction vessel (m)

De effective diffusion coefﬁcient of the solid particle (m2/s)

DL axial dispersion coefﬁcient (m2/s)

dp particle diameter (m)

kf mass transfer coefﬁcient in the ﬂuid phase (m/s)

L bed depth (m)

r radial position within the particle (m) R particle radius (m)

t extraction time (s)

z axial position along the bed (m) u interstitial velocity (m/s)

e

porosity of the bed (–)

e

p porosity of the solid particle (–)

s

residence time (s)

u U/

e

U F/AE

AE cross-sectional area of the extraction vessel (m2)

U superﬁcial velocity (m/s)

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is quite similar. At their most basic all three systems consist of a solvent supply, a pump for transporting the solvent, a heater for heating the solvent, a pressure vessel where the extraction occurs, a means to control the pressure in the system, and a collection ves-sel for the extract (Figs. 1 and 2). Individual units may have differ-ent conﬁgurations, instrumdiffer-entation, valves, by-passes, gas purge systems, and safety features not shown here.

The supercritical ﬂuid extractor differs slightly from the other pressurized extraction systems. Supercritical CO2, which is a

non-polar solvent, is sometimes modiﬁed with non-polar solvents such as ethanol to reduce the polarity of the system and allow for the extraction of a wider range of bioproducts. The co-solvent enters the system after the main pump (Fig. 1) via a secondary pump (not shown). Other differences for supercritical ﬂuid extractors in-clude a heat exchanger before the pump, which is necessary to cool the CO2to maintain it in a liquid state so that it may be pumped,

and there may be a secondary pressure regulator after the separa-tor (collection vessel) to control the pressure and allow for frac-tionation of the extract (Fig. 1). Without this pressure regulator, the CO2would go from a supercritical state to a gaseous one, and

all of the extract would precipitate from the solvent into the sepa-rator. By maintaining a small pressure in the separator, some com-pounds may remain in the solvent. By adding additional separators, it is possible to fractionate the extract into several dif-ferent compounds by altering the pressure within each separator (Díaz-Reinoso et al., 2006; Esquível et al., 1999; Wang et al., 2004).

Because PLPW extraction is a special case of PLE, the design of both systems is similar (Fig. 2). The PLPW extraction system adds a condenser after the extraction vessel to cool the water down so that it does not flash to steam. The heat capacity of water is very large, so considerable effort is necessary to remove the excess en-ergy. In comparison, most organic solvents have a much lower heat capacity (approximately 35–45% lower). The lower heat capacity, coupled with the fact that most PLE systems are small labora-tory-scale units using low volumes of solvent, means that the sol-vent will cool before exiting the extraction system provided that the length of tubing after the extractor is sufficiently long.Richter et al. (1996)found that when using extraction cells between 3.5 and 32 mL in volume, the temperature of exiting solvent was less than 35 °C, even if the extractor was at 100 °C, so long as a length of tubing of approximately 30 cm was used after the extractor. If bigger scale PLE units, using larger volumes of high temperature solvent were constructed, the use of a condenser would become necessary. For many laboratory-scale PLE and PLPW extractors, the solvent heater before the extraction vessel is omitted, and heating occurs by placing the extraction vessel within an oven. To allow the solvent to come to the proper operating temperature, long lengths of tubing are coiled within the oven before the extrac-tion vessel to increase residence time and surface area. The benefit of this system is that the extraction vessel is heated at the same time without the necessity of supplemental heaters for the vessel itself.

Due to the difﬁculties of designing a suitable device for charg-ing a solid material into a high pressure and potentially high tem-perature system, these extractors are usually constructed as a batch system, with extraction occurring in a vessel through a ﬁxed bed of material. Such equipment is most suitable for high value and low volume products such as pharmaceutical and nutraceutical products. As such, there has not been great accep-tance of this technology within the food processing industry where low value, high volume processing is the norm. Design considerations similar for all three extractors with the solvent, operating temperature, and pressure ranges being the system variables. With the choice of a suitable pump and extraction capacity, only the dimensioning of the extraction vessel and opti-mizing of the processing parameters are required. Work in the laboratory with small-scale equipment is invaluable for determin-ing the appropriate factors for scale-up, usually with the develop-ment of process models. Knowledge of phase equilibria, mass transfer rate, and solubility data are important for scale-up of the extraction process and equipment.

Fig. 1. Schematic diagram of a typical supercritical ﬂuid extractor with a single separator.

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3. Mass transfer and phase equilibria 3.1. Modeling of pressurized ﬂuid extraction

Modeling of the extraction process is a beneficial step in the effective scale-up from laboratory to pilot or industrial-scale sys-tems (Bulley et al., 1984; del Valle and de la Fuente, 2006).del Valle and de la Fuente (2006)present a review of several kinetic and equilibrium models of SFE of oilseeds. Some of the models in literature are empirical in nature and are simplistic, but they de-scribe the extraction process very well. However, these models do not provide insight into the underlying mass transfer phenom-ena that occur during the extraction process, with no correspon-dence to the materials being extracted, and therefore are unsuitable for scaling up purposes (del Valle and de la Fuente, 2006; Reverchon and De Marco, 2006). Kinetic (Anekpankul et al., 2007; Ho et al., 2008; Kubátová et al., 2002) and diffusion models (Anekpankul et al., 2007; Ho et al., 2008) have been applied to PLPW extraction, but no modeling seems to have been done for PLE. Some of the most interesting mass transfer models are those based on mass balance equations for thin sections of a packed bed (delValle and de la Fuente, 2006). Such models are able to uti-lize data from laboratory-scale units to determine the mass trans-fer within a larger unit.Bulleyet al. (1984) present a simplified model of mass balance in a single element of a bed for the extrac-tion of oil from flaked canola seed. The use of the fixed bed model for SFE for the extraction of oil-containing seeds by researchers has been reviewed bydel Valle and de la Fuente (2006). The model is also applicable for extraction of other substances from organic matter in packed beds such as the PLPW extraction of mannitol from olive leaves (Ghoreishi et al., 2008). However, PLPW extrac-tion and PLE are still in their infancy and larger units are still rare, so modeling from a system, or bed perspective is not common. 3.2. General model for the mass balance in a fixed bed extractor

The mass transfer within a ﬁxed bed, inside a cylindrical extrac-tor, may be described by a set of partial differential equations (Akgerman and Madras, 1994). The extraction vessel is divided into

ﬁnite difference volume elements of heightDz (Fig. 3). The result-ing equations may be solved usresult-ing numerical techniques. The assumptions of the model are: (i) pressure losses and temperature gradients are negligible within the bed, i.e. system is isothermal and solvent density remains constant along the bed; (ii) the solute is subjected to axial dispersion in the solvent due to a concentra-tion gradient along the bed; (iii) physical properties of the solvent remain constant; (iv) bed porosity (

e

) remains constant, hence interstitial velocity is also constant); (v) there are no concentration gradients over the cross-section of the extractor. The fluid phase mass balance for a finite volume element in a fixed bed is thus:

@Cf @t ¼ DL @2Cf @z2 u @Cf @z kfap

e

Cf Cps ð1Þ

If bed porosity is

e

, then the volume of particles is 1

e

. Assuming spherical particles, the surface area of the particles in a unit volume of bed (speciﬁc surface area) is deﬁned as:

ap¼ 3ð1

e

Þ

R ð2Þ

Combining Eqs.(1) and (2)gives:

@Cf @t ¼ DL @2Cf @z2 u @Cf @z 3 R 1

e

kfðCf CpsÞ ð3Þ

The associated boundary conditions are:

DL @2Cf

@z2 ¼ uCf ðat z ¼ 0; for all tÞ ð3aÞ @Cf

@z ¼ 0 ðat z ¼ L; for all tÞ ð3bÞ

Cf¼ 0 ðat t ¼ 0; for all zÞ ð3cÞ

The axial dispersion coefficient is a measure of back-mixing in the bed during fluid flow. Axial dispersion is often neglected (DL= 0)

and represents no back-mixing within the extraction vessel, or the case of plug ﬂow (del Valle and de la Fuente, 2006). It is desir-able to avoid axial dispersion of the solute along the bed in order to maintain the driving forces for extraction at their maximum level

z = 0 z + Δz z z = L U, Cf = 0 Bed void fraction (ε) Solid particle U, Cf (L) u, Cf (z + Δz) u, Cf (z) Particle void fraction (εp) R C r Cs Cp Cf Cs Cp

Fig. 3. Diagram of a fixed bed extractor showing a finite volume element, bed and particle characteristics, and the concentration profiles of solute within the particles and surrounding solvent (adapted fromdel Valle and de la Fuente, 2006).

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(del Valle and de la Fuente, 2006). Axial dispersion may be avoided and hence neglected in the modeling by: maintaining a larger inter-stitial solvent velocity in the bed; by using sintered metal plates to cause P10 bar of pressure loss in the solvent at the inlet; to avoid gradients in solvent viscosity due to radial gradients in temperature by operating at near ambient temperatures and/or using smaller diameter extraction vessels; and by maintaining a small enough particle diameter to extractor diameter ratio (dp/dE< 0.1) to

dimin-ish the radial variations in bed porosity (Brunner, 1994).

Fluid transport of the solute within or out of the particle is as-sumed to take place by diffusion through a network of pores (Fig. 3). The particle mass balance (of the ﬂuid within the pores) is thus:

e

p @Cp @t þ @Cs @t ¼ De

e

p r2 @ @r r 2@Cp @r ð4Þ

The associated boundary conditions are:

kfðCf CsÞ ¼

e

pDe @Cp

@r ðat r ¼ R; for all z and tÞ ð4aÞ @Cp

@r ¼ 0; @Cs

@r ¼ 0 ðat r ¼ 0; for all z and tÞ ð4bÞ Cp¼ Cpo;Cs¼ Cso ðat t ¼ 0; for all z and rÞ ð4cÞ

The concentrations, Cf, Cp, and Csin Eqs.(3) and (4), may be utilized

on a wt/wt basis rather than a wt/vol basis by incorporating the sol-vent and particle densities into Eqs.(3) and (4)(Bulley et al., 1984; Lee et al., 1986). Density of the solvent will vary with the processing parameters of temperature and pressure. Both are significant for supercritical fluids, however, for PLE and PLPW extraction, changes in density with pressure tend to be negligible due to the incom-pressible nature of liquids (no change in density with pressure). Experiments have shown that operating pressure is not a significant extraction parameter for PLE (Choi et al., 2003; Kawamura et al., 1999; Ong and Len, 2003) and PLPW extraction (Gogus et al., 2005; Rovio et al., 1999; Soto Ayala and Luque de Castro, 2001) of bioproducts. Yet, there may be some benefit of operating at higher pressures whereby the increased pressure aids in penetration of the solvent into the cells and pores of organic materials. This is weighed against the potential for collapsing of the bed, and a reduction in porosity leading to lower interstitial velocities. However, for PLPW extraction an increase in pressure from 10 to 600 MPa at 25 °C would result in an increase in the dielectric constant from 79 to 93 (Haar et al., 1984). This would yield a slight increase in the polar-ity of the PLPW, which would negatively affect the extraction of non-polar compounds. In both systems there is a critical constraint where it is necessary to maintain the pressure at a sufficiently high level in order to maintain the solvent in a liquid state, but there is little need to significantly increase pressure above that level. 3.3. External mass transfer

The extraction process takes place in two distinct stages gov-erned by two different modes of mass transport. The ﬁrst stage oc-curs early in the process, when the oil or solute is freely present on, or diffused to the surface of the material, and extraction is con-trolled by the external mass transfer (coefﬁcient kf in Eq. (3)).

The product is freely available to be removed so long as the max-imum solubility of the solute within the solvent has not been reached. Extraction of some substances and materials occur wholly within the ﬁrst stage of extraction. In this case the extraction would be controlled only by Eq.(3), and the transport within the particles (Eq.(4)), would be neglected. This is most often the case for SFE of oil bearing materials, when they are processed into thin ﬂakes (Bulley et al., 1984; Lee et al., 1986), although it has also

been applied to the extraction of other thin particles like androgra-pholide, a diterpenoid lactone, from Andrographis paniculata leaves (Kumoro and Hasan, 2006). Kim and Mazza (2006) found that extraction of phenolic compounds from flax shives (expressed as a total phenolic compounds value) in PLPW was controlled by external mass transfer at low flow rates.Anekpankul et al. (2007) found that at low flow rates (between 1.6 and 2.4 mL/min) the extraction of the anthraquinone damnacanthal from roots of Mor-inda citrifolia was controlled by external mass transfer. At high flow rates (>2.4 mL/min), results suggested that internal mass transfer was the controlling factor. Some extraction processes will be con-trolled by a combination of factors depending on the extraction conditions.

3.4. Internal mass transfer

The second stage occurs when the freely available solute pres-ent at the surface of the material is removed. The extraction is then controlled by the movement of the product within the material, or the internal mass transfer coefﬁcient (Dein Eq.(4)). The diffusion of

the dissolved solute within the solid matrix is usually the rate lim-iting step in the extraction of most botanicals (Schwartzberg and Chao, 1982) and for most SFE (Kubátová et al., 2002).Machmudah et al. (2008) found that the extraction of oil from ground seeds (rosehip, loquat, and physic nuts) with SFE was largely controlled by the internal mass transfer of oil, especially at high pressures. Ho et al. (2008)found that internal diffusion was the controlling factor in the extraction of lignans from ﬂaxseed meal with PLPW. Similar results were obtained for extraction of lignans from whole ﬂaxseed (Cacace and Mazza, 2006). Kubátová et al. (2002)found that the extraction of essential oil from savory was controlled by internal mass transfer of the dissolved solute within the solid ma-trix during SFE, but was controlled by the external mass transfer of solute to solvent in PLPW extraction. This is partially explained by the fact that PLPW is more likely to extract organic material than SFE (Hawthorne et al., 2000).Kubátová et al. (2002)hypothesized that PLPW is more effective at altering sample matrices and dis-placing analytes from their original binding sites than supercritical CO2. The extraction process may be controlled by a combination of

different processes and the exact mechanisms controlling the extraction may change depending on the extraction parameters (Anekpankul et al., 2007).

3.5. Effects of processing parameters on the extraction process Optimization of the extraction by altering the processing parameters is an important step in the scale-up procedure, often done with laboratory or pilot-scale units before ﬁnal scaling to pro-duction-scale systems. A great deal of work has been done on determining the operating parameters for SFE (Díaz-Reinoso et al., 2006; Reverchon and De Marco, 2006), PLE (Kaufmann and Christen, 2002), and PLPW extraction (Cacace and Mazza, 2007; Kim and Mazza, 2009), for the optimization of the extraction of bioproducts from organic material. A balance must be achieved be-tween maximum yield (ratio of amount extracted by total avail-able), and maximum concentration. For high value products, a maximum yield may be desirable, but as value of the product de-creases, the economics of the extraction would dictate an increase in the concentration of the extract. In many cases the extraction process in commercial operations is considered complete when 90% of the solute has been extracted (del Valle et al., 2005). Reduc-tion of the extract volume by removing the solvent can be an expensive process, and the added cost is not compensated from the additional product recovered. The ideal operating conditions for an extraction system occur when the system is operating in

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the ﬁrst stage of extraction, which is limited by the solubility of the solute dissolved in the solvent.

3.5.1. Solubility

Solubility of a solute in a particular solvent is fundamental to the design of an extraction process. It controls the ideal operating conditions and sets the minimum amount of solvent necessary to complete an extraction (Shi and Zhou, 2007). When the solubility of the solute in the solvent has been reached, under suitably low flow rates, the extraction will be operating in the first stage where-by the extraction is controlled where-by the external mass transfer. Solu-bility of a product can be determined from the slope of the linear (constant extraction rate, external mass transfer limiting stage) portion of the extraction curve (Bozan and Temelli, 2003; Güçlü-Üstündag˘ and Temelli, 2005; Özkal et al., 2005). Solubility is dependant on solvent density. With supercritical fluids this means that the solvating power of the fluid may be altered with changes in the system’s pressure (Fig. 4). Liquids are largely incompressible in the subcritical region, so pressure tends not to have a great influ-ence on solubility and extraction in PLE and PLPW systems (Smith, 2002). In this case, temperature has a more important effect on product solubility than pressure. As flow rate increases, there is the possibility in extraction systems that the free solute on the sur-face of the material will rapidly be removed and the mass transfer of the dissolved solute within the product will become the rate limiting step. The key is not to keep increasing the flow rate when there is no benefit. Doing so will only decrease the concentration of the extract and use more solvent than is necessary to complete the extraction.

3.5.2. Flow rate

Flow rate has a direct bearing on the mode of mass transfer for the system, and therefore the design of a fixed bed extractor. When the flow rate is sufficiently large there forms a concentration gra-dient between the solid’s surface and the solvent and the extrac-tion falls into the second extracextrac-tion stage whereby diffusion within the material becomes rate controlling. If the flow rate is lowered sufficiently, the removal of solute from the material’s sur-face decreases, and diffusion of the solute within the solid matrix is adequate to replace the solute removed from the material. The fluid phase mass balance (Eq.(3)) contains a term which incorpo-rates the flow rate of the extractor. As stated above, when an extraction is controlled only by external mass transfer, then the extraction is governed solely by Eq.(3), and the kinetics of the

sys-tem prevails. If an extraction is limited by the external mass trans-fer from the solid to the solvent, then an increase in flow rate would result in an increase of the extraction rate of the solute. Con-versely, if an extraction is limited by the internal mass transfer of the system, and the partitioning thermodynamics limit the extrac-tion rate, then it would be expected that an increase in the flow rate would have little effect on the extraction rate of the solute (Eq.(4)does not contain a flow rate term).Ho et al. (2008)found that internal mass transfer was the controlling factor in the extrac-tion of lignans from flaxseed meal with PLPW, as seen by the neg-ligible effects of flow rate on the process.Shalmashi et al. (2008) found that increasing the flow rate during PLPW extraction in-creased the rate of caffeine removal from black tea leaves, which indicated that extraction was mostly controlled by the transfer from the solid to the solvent (first stage). However, the most effi-cient flow rate was 2 g/min, which achieved the maximum yield of caffeine in the same amount of time, but with half the volume of solvent when compared to extraction with a flow rate of 4 g/ min. Knowledge of the mass transfer characteristics in the extrac-tion vessel can aid in the optimizaextrac-tion of the flow rate to minimize extraction time, while maximizing yield.

3.5.3. Material properties

Physical properties of the material being extracted are impor-tant, as they play a role in the phase equilibria and mass transfer within the extraction vessel. Size reduction can improve extraction because of mass transfer in the form of diffusion, which is inversely proportional to the square of the characteristic dimension (Crank, 1975; Cussler, 1984; Osburn and Katz, 1944). In addition, with lar-ger particles less cell walls will be undamaged, reducing solvent penetration and product extraction. Reducing the particle size will increase the surface area per unit mass between particle and sol-vent, but very small particles will produce smaller volume voids in the sample matrix, providing resistance to ﬂuid ﬂow within the extractor. There are limits to suitable size reductions before porosity of the bed (

e

, Eq.(3)) decreases to such an extent that flow becomes difficult, and the possibility of channelling and fines in the extract decrease quality. If particle size is reduced sufficiently, then internal diffusion ceases to be the controlling factor and external mass transfer from the particle to the solvent, becomes the controlling factor (Aguilera, 2003).

3.5.4. Length to diameter ratio

Bed length to diameter ratio is an important consideration in the design of extraction vessels. Vessel cost is partially inﬂuenced by volume, but mainly by the diameter (Laurent et al., 2001). In tall extractors back-mixing may occur, and large diameter vessels may result in heterogeneous extraction, resulting in radial effects with-in the extractor (Brunner, 1994; del Valle et al., 2004). For particle sizes between 0.4 and 0.8 mm the recommended length to diame-ter ratio is 6:1, while larger ratios of 9:1 have been used for the decaffeination of coffee (7 mm diameter particles) and smaller ra-tios of 3:1 for materials which may swell (Laurent et al., 2001). Eg-gers (1996)recommends a length to diameter ratio of 4–6 for the SFE of oilseeds with CO2. Many length to diameter ratios for

exper-imental SFE units vary between 2 and 29 (del Valle et al., 2005; Meireles, 2003) and for PLPW extraction and PLE they generally fall in a similar range of 2–25 (Cacace and Mazza, 2007; Ho et al., 2008; Richter et al., 1996). But, with respect to cost, this value should vary between 5 and 7 (Meireles, 2003). With knowledge of the sol-vent to solid ratio and mass transfer rate, a system may be de-signed to produce a certain yield, or throughput per day. Mass transfer will dictate efﬁcient ﬂow rates for the extractor. Compar-isons can be made between industrial units and experimental units if the solvent to solid ratio are the same, and scale-up is accom-plished by increasing the vessel diameter and length to maintain

0 5 10 15 20 25 30 35 0 10 20 30 40 50 60 70 Pressure (MPa) Solubility (mg/g) 40ºC 50ºC 60ºC

Fig. 4. Effect of pressure on the solubility of hazelnut oil in supercritical CO2 (adapted fromÖzkal et al., 2005).

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a constant ratio, and solvent is properly dispersed into the system (Meireles, 2003). However, this neglects to take into consideration the situations that may be encountered in a larger unit, such as, material agglomeration, increased pressure drop, fluid channelling, and the increased difficulty maintaining plug flow with increased vessel diameter.

4. Scale-up of pressurized ﬂuid extraction

An effective model of the extraction system is crucial for deter-mining the basic mass transfer data necessary for scale-up proce-dures. There have been many studies on the kinetics of extraction for high pressure extraction systems, but there has not been substantial analysis of the effect of process scale on the extraction rates (del Valle et al., 2004). When designing equipment certain inputs may be fixed, such as temperature, pressure, flow rate, or pH, in order achieve a desired extraction efficiency or rate. These are often determined in a laboratory using nothing more than glassware and chemicals, or other laboratory-scale equipment to produce product on the gram scale or smaller. The problem then is how to effectively scale-up to larger production runs, or produce equipment to do the same. Appropriate scale-up factors must be determined for the process being examined. It is not sufficient to scale everything up at the same rate, as the process is often con-strained by certain key parameters, and all others do not contrib-ute to the desired outcome. For instance, one of the key parameters in the operation of an extractor is the superficial veloc-ity of the solvent in the extractor itself. This is the velocveloc-ity across the cross-section of the extractor defined as the volumetric flow rate of the solvent divided by the cross-sectional area of the extrac-tor. If the diameter is increased by a factor of ten, the volumetric flow rate would have to be increase by 100 to maintain the same superficial velocity (area is a squared term). At the same time, to maintain the same bulk density and porosity in the extractor, the mass of material would have to increase by 1000 (volume is a cubed term) if the length is also increased by a factor of 10 to main-tain the same length to diameter ratio.

This process is illustrated byLagadec et al. (2000)who set out to scale-up a PLPW system to remove contaminates from soil. The capacity of a laboratory unit was scaled-up by a factor of 1000 to increase the amount of soil processed from 8 g to 8 kg. To maintain an equivalent bulk density, and hence bed porosity (

e

), the volume was increased by a factor of 1000. This was accomplished by scal-ing up the extraction vessel diameter and length by a factor of 10 (9.4 mm i.d. 100 mm long to 102 mm i.d. 1000 mm long). For removal of polycyclic hydrocarbons (PAHs) from contaminated soil the same solvent flow-to-solid ratio was used for the scale-up pro-cedure. The flow was set at 0.5 mL/min and 0.5 L/min for the labo-ratory- and pilot-scale extractor respectively. At 275 °C and the specified flow conditions, all PAHs were removed within 60–70 min. However, when the flow rate was increased to 600 mL/min in the pilot-scale extractor, the removal of PAHs was completed in 35–40 min, which indicates that the extraction was controlled by the external mass transfer of the solute to the solvent. When external mass transfer is the controlling factor, as it is in the first stage of extraction, it is common practice during scale-up to keep the solvent to solid ratio constant (del Valle et al., 2005). If the maximum solubility of the solute in the solvent is maintained, then the extraction may be completed using the least amount of solvent possible. Whereas, when internal mass transfer is the controlling factor, as in the second stage, the practice is to keep the solvent flow-to-solid ratio constant in order to extend the contact time between the solvent and material (del Valle et al., 2005).Lagadec et al. (2000) kept the solvent flow-to-solid constant even though the extraction was controlled by the external

mass transfer from the solid to the solvent. In removal of soil con-taminates only the efﬁcacy of the extraction is important. As such, Lagadec et al. (2000)did not examine the extraction kinetics and no information on the extraction rates between the different scale extractors was given.

Berna et al. (2000)modeled the extraction of essential oil from orange peels with supercritical CO2 using Lack’s extended plug

flow model as adapted bySovová (1994). Experiments were con-ducted at two different scales (feedstock scaled-up by a factor of 20), while maintaining the same solvent flow-to-solid ratio and solvent-to-solid ratio. By studying the parameters from Sovová’s (1994)model,Berna et al. (2000)were able to incorporate the ef-fect of bed height into the model. Their results showed that there was similar behaviour in all experiments and that as long as there was homogeneous flow within the extractor, the height of the bed had little effect on the extraction at the same scale of operation. Berna et al. (2000)found that the addition of diatomaceous earth at the inlet of the extraction vessel promoted homogeneous flow, possibly by dispersing the solvent evenly across the diameter of the vessel and ensuring plug flow. In addition to issues of flow het-erogeneity, there is a greater possibility for the agglomeration of particles with high oil contents at larger scales, which would in-crease the resistance to internal mass transfer of particles due to a larger effective diameter (Berna et al., 2000).

Perrut et al. (1997) developed a model of supercritical CO2

extraction of sunﬂower seed oil based on the results of a laboratory and pilot-scale system with an extraction vessel that was scaled-up by a factor of 10 (from 0.15 103_{to 1.5 10}3_m3_{). However,}

none of the other extractor dimensions were scaled proportionally, resulting in an altering of the length to diameter ratio from 4.4:1 in the laboratory-scale unit to 3.5:1 in the pilot-scale unit. Literature suggests that comparisons can be made between units if the scale-up is accomplished by increasing the vessel length and diameter to maintain a constant ratio (Meireles, 2003). There was no discussion byPerrut et al. (1997)about the effects of scale-up on the extrac-tion kinetics.

Simple scale-up by a set factor has been questioned bydel Valle et al. (2004) who studied the effects of process scale on the oil extraction kinetics of supercritical CO2extraction of rosehip seed.

The ﬁrst series of experiments were carried out on a laboratory-scale system with a vessel volume of 5.0 105_m3_{using 26 g of}

sample. The extraction conditions for the laboratory-scale system were 40 °C and 30 MPa, or 50 °C and 40 MPa at superﬁcial CO2

velocities of 0.23–1.40 mm/s. Results were used to develop an one-dimensional, unsteady state model with axial dispersion of solute in the supercritical phase for a packed bed. It was a two-stage model with extraction rate initially controlled by oil solubil-ity and mass transfer from the material to the solvent, and by inter-nal mass transfer within the particles during the latter stages of extraction. The model was used to simulate the extraction of a pi-lot-scale unit, which was scaled-up by a factor of 30 for the vessel volume, with recycling of the solvent. However, equivalent length to diameter ratio of the two vessels was not maintained as the in-ner diameter was scaled by a factor of 3 while the length was only scaled by a factor of 1.5. The extraction conditions for the pilot-scale system were 40 °C and 30 MPa, or 50 °C and 40 MPa at a superﬁcial CO2 velocity of 0.57–0.58 mm/s. Results showed that

the extraction in the pilot-scale system was slower than those pre-dicted from the model which utilized parameters from the kinetic data from the laboratory-scale unit.del Valle et al. (2004) attrib-uted the discrepancy to ﬂow heterogeneity in the extraction vessel, increased dispersion of solute between the extraction and separa-tion vessels, entrainment of oil droplets in recycled gaseous stream, which would decrease the concentration gradient between the solute and solvent in the extraction vessel, or a combination of the three. Flow heterogeneity may occur due to radial variations of

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bed porosity and solvent viscosity, which could result from packing the bed with relatively large particles (dp/ue< 10, where dpis the

particle diameter and ueis the internal diameter of the extraction

vessel), or attempting heat transfer by conduction through the wall of the extraction vessel, respectively (Brunner, 1994). These would result in higher interstitial velocities near the wall, which would result in slower extraction in the centre of the extraction vessel (del Valle et al., 2004). This was manifested as a reduction in the period of external mass transfer in the pilot-scale system caused by earlier depletion of solute near the wall of the extraction vessel, causing an earlier drop in solute concentration in the exiting sol-vent. del Valle et al. (2004) concluded that radial variations in porosity could be more important with small diameter vessels on the laboratory-scale, whereas variations in viscosity could be more important when using large diameter vessels and high tempera-tures at the pilot-scale. Hence the caution in scaling up of extrac-tion vessels by a constant factor because of changes in mass transfer phenomena related to extraction conditions that do not al-ways depend on the mass transfer coefﬁcient.

Eggers and Sievers (1989)hypothesized that scale-up should be accomplished by altering the vessel height while maintaining the same superﬁcial velocity in both extractors to ensure equivalent mass transfer properties. Results for the extraction of evening primrose seed oil (EPO) in a pilot-scale (200 L extractor volume) and production-scale unit (2000 L extractor volume) with super-critical CO2differed. Extraction in the production-scale unit took

longer but required a lower solvent to solid ration. Loading of EPO in the solvent increased from 0.63% in the pilot-scale, to 0.83% in the production-scale unit. The theoretical phase equilib-rium for EPO in supercritical CO2under the extraction conditions

of the experiment was 0.9% (Eggers and Sievers, 1989). To account for this discrepancy,Eggers and Sievers (1989)proposed the use of higher velocities in larger scale extraction vessels compared to smaller scale vessels in order to keep the residence time (calcu-lated as a function of bed depth and superﬁcial velocity,

s

= L/U) equal in the vessels at the two scales.Ho et al. (2008)found that the yield of lignans from ﬂaxseed meal increased with increasing bed depth, and was independent of the solvent to solid ratio. This was attributed to the residence time. It was suggested that longer residence times allowed sufﬁcient time for the solvent to penetrate through the solid pores and allowed for increased yields.

King et al. (1997) extracted EPO on a laboratory-scale unit (72 mL extraction vessel) and compared their results with those

of Eggers and Sievers (1989). The results from the laboratory-and production-scale units were in agreement. Only Eggers and Sievers (1989)results for the pilot-scale unit were lower. No expla-nation was given for these results. However, the bulk density was higher and the superficial velocity was lower in the laboratory-scale unit compared to the levels used by Eggers and Sievers (1989)for the pilot- and production-scale units. This would result in lower bed porosities in the laboratory-scale unit that would lead to a lower interstitial velocity that would be offset to a degree by the decrease in superficial velocity. It is hard to make concrete con-clusions about the effect of scale when parameters that influence the extraction are not held constant, or scaled in the appropriate fashion.

5. Continuous pressurized ﬂuid extraction

Most applications of pressurized fluid extraction in the litera-ture operate as a batch process, whereby a material is extracted by immersion or percolation of solvent through a fixed bed. This method of extraction is largely undesirable in industry due to nec-essary interruptions for loading and unloading of the extraction vessel, and the large amounts of solvent required to complete an extraction (Eggers and Jaeger, 2003). To be truly economical, and therefore desirable to industry, most processing must operate on a continual basis so that all equipment in a facility is being utilized to the fullest. There is great difficulty to convert the technologies of pressurized solvent extraction to continuous operation. Design of a continuous system would be best done using information gained from a corresponding continuous laboratory-scale unit. However, it is more convenient to keep the product stationary in a fixed bed. Design of a continuous extractor is still possible from informa-tion gained in experimentainforma-tion with fixed beds by studying the equilibrium distribution of oil between the material and solvent, and determination of the mass transfer rate of solute from the material to the solvent (Bulley et al., 1984). One area where contin-uous pressurized solvent extraction has achieved success is with counter-current liquid–liquid extraction.

5.1. Counter-current liquid–liquid pressurized extraction

Counter-current contact is achieved when solvent ﬁrst comes into contact with the material that has been the most extracted, and exits the extractor in contact with the least extracted material. In this way the differences in concentration between the solute and solvent are greatest, providing the greatest driving force in the extraction process. Counter-current operation of an extractor is thus able to reduce the amount of solvent used, increases through-put, enables higher extract concentrations in the solvent, and low-er residual concentrations in the matlow-erial (Brunner, 2005). True counter-current contact is easily achieved when a liquid–liquid extraction is conducted, or when a slurry of the material is created that may be pumped (Fig. 5). Systems may be constructed to in-clude multiple separators, solvent regeneration, or reﬂuxing of the extract (Singh and Rizvi, 1994; Brunner, 2005). Continuous counter-current pressurized solvent liquid–liquid extractors have been studied by several researchers. Brunner (1998) reviewed and discussed the development process for supercritical counter-current liquid–liquid extractors.Singh and Rizvi (1994)have com-pleted a design and economic analysis of a system for processing milk fat with supercritical carbon dioxide.Ibáñez et al. (2002) con-structed a system for the concentration of sterols and tocopherols from olive oil using supercritical carbon dioxide.Ooi et al. (1996) used a continuous counter-current supercritical CO2extractor to

reﬁne palm oil. The system was successful in removing the free fatty acids while retaining the carotenes in the oil which are

Fig. 5. Schematic diagram of a continuous counter-current system used for liquid– liquid extraction with supercritical carbon dioxide without solvent regeneration and only one separator (adapted fromOoi et al., 1996).

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normally destroyed in traditional physical or chemical refining processes. Transfer of material into and out of counter-current li-quid–liquid extractors is not as difficult as for counter-current so-lid–liquid extractors because the product can be pumped into the extractor in liquid form with the use of a piston or diaphragm pump. The biggest problem with continuous pressurized counter-current solid–liquid extraction systems is how to feed solid mate-rial from atmospheric pressure into a vessel at high pressure, then feed the material back to the atmosphere after extraction. There is a lack of reliable sluice systems for high pressure vessels for con-tinuous charging and discharging of solids into the extraction ves-sel. To gain the benefits of a continuous counter-current extraction without a suitable sluice system it is possible to conduct a quasi-continuous extraction system with multiple stages.

5.2. Quasi-continuous solid–liquid pressurized extraction

To create a quasi-continuous extraction system, several batch extractors are connected and operated in series to form a battery (Fig. 6). Fresh solvent is loaded into the vessel which contains

the most exhausted solid (Vessel 2). The solvent and extract from this vessel then passes through each vessel in succession, until it is discharged from the last one, which contains the most recently loaded material (Vessel 1). One vessel is bypassed to unload spent material and recharge with fresh material (Vessel 3). With this set-up extraction occurs on a continuous basis, although loading and unloading is discontinuous to allow for depressurization/pressuri-zation and opening of the pressure vessels. On the laboratory-scale the ASE extractor operates on a semi-continuous basis by setting up a series of vessels and batch extracting them one at a time (Richter et al., 1996). Most industrial-scale units are proprietary, and therefore detailed information of their design and operation are not available in the literature. However, several quasi-continu-ous pressurized extraction system designs have been proposed and discussed. Leyers et al. (1991) presented a detailed design and study of a system that used supercritical CO2to decaffeinate coffee.

The system consisted of a battery of four vessels operating in a con-tinuous mode with three vessels operating at one time with one being loaded or unloaded with fresh coffee beans. The system was operated at 14–35 MPa at temperatures of 70–130 °C with a

Fig. 6. Schematic diagram of the setup and ﬂow through a battery of extraction vessels. Each vessel can be bypassed for loading and unloading as indicated by the discharge bypass shown only for the middle extractor (adapted fromEggers and Jaeger, 2003).

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residence time of 6–12 h.Lack and Seidlitz (1993)proposed a sim-ilar design of a quasi-continuous system to decaffeinate coffee using a battery of four extraction vessels for while working for Schoeller–Bleckmann before moving on to form NATex.

5.3. Continuous counter-current solid–liquid pressurized extraction The design of a continuous counter-current extraction system without a suitable sluice system is possible by using screw convey-ors or extruders to pack the solid material into a gas-tight plug, preventing pressure losses along the moving bed (Eggers, 1988, 1994). Eggers et al. (1985) proposed a design for a continuous counter-current extraction system using supercritical CO2for the

extraction of oil bearing materials, which was later patented (Uni-ted States Patent 4,675,133;Eggers and Schade, 1987). The system consists of a high-pressure extraction vessel placed between two screw presses (Fig. 7). The charging press serves to first partially de-oil the material, then to build up the pressure to form a gas-tight plug. The plug of material that is formed within the press can reach temperatures of 60–90 °C due to the work of compres-sion and wall friction (Eggers and Sievers, 1989). At the same time, the gas-tightness of the plug has been demonstrated for pressure differences up to 40 MPa for periods as long as 5 min, which is lar-ger than the residence time within the screw press (Eggers and Sie-vers, 1989). The pressure in the extraction vessel influences the pressure profile of the screw press. When the pressure in the extraction vessel increases, there is a corresponding pressure in-crease in the screw press which helps ensure the gas-tightness of the plug (Eggers, 1996). The discharging press acts in the same manner to form a gas-tight plug for the extraction vessel while allowing for the residue to be continuously removed. Extraction of the material occurs in the high-pressure vessel between the two screw presses. Supercritical CO2 enters the bottom of the

extraction vessel and leaves at the top, moving opposite the ﬂow of the material and thus extracting under counter-current condi-tions. Inside the extraction vessel there is a slowly revolving mechanical stirring mechanism, which aids in the transport of the material through the extraction vessel and prevents clogging. A pilot plant was produced that could process 2–25 kg/h of mate-rial (Eggers and Sievers, 1989; Eggers, 1996).

Wingerson and Lehrburger (2004)designed and tested a contin-uous counter-current screw extractor to operate at elevated

pres-sures and temperatures for which a patent has been filed (United States Patent Application# US 2006/0283995 A1; Wingerson (2006)). The goal was to design and validate a system to produce a purified cellulose product on a continuous basis by utilizing a process which minimizes the use of enzymes by fractionating bio-mass with aqueous media (water and alkali), at elevated tempera-tures up to 235 °C. The extractor utilises a single threaded shaft that has a plurality of reaction zone segments along the length of the shaft that are separated from each other by dynamic plug seg-ments. The shaft has a different thread pitch in the reaction zone segments than in the dynamic plug segments, which help to create the plugs.Wingerson and Lehrburger (2004)successfully achieved the formation of two plugs within the screw extractor using ground corn stover capable of retaining pressures of 1.4 MPa to over 6.9 MPa. They also achieved some success producing coun-ter-current extraction and discharging the liquid and solid stream separately. However, they were as yet unable to operate on a com-pletely continuous basis at the desired temperatures and pres-sures, or to produce a purified cellulose product.

There has been some success in designing a sluice system for charging and discharging solids from a continuous counter-current pressurized ﬂuid extraction system. A group from Denmark has created a system that utilizes a so called ‘‘particle pump”, which is capable of transferring solid particles between two areas of dif-ferent pressure (Christensen and Christensen, 2004; Thomsen et al., 2006, 2008). The particle pump, for which a patent applica-tion has been submitted (United States Patent Applicaapplica-tion# US 2004/0184900 A1; Christensen and Christensen, 2004), utilizes two sluice chambers, one of which is charged with a piston screw while the other is being discharged. Pressure tightness is achieved by the use of pistons or pressure locks in the pump. The system is capable of processing low density biomass such as straw on a con-tinuous basis for the extraction of hemicellulose with PLPW for the purpose of ethanol production. The pilot plant system, for which a patent has been granted (WIPO Patent# WO/2007/009463; Chris-tensen and ChrisChris-tensen, 2007), has a potential production capacity of 1000 kg/h although it has only been tested in the 50–150 kg/h range (Thomsen et al., 2006, 2008). The system has been designed to operate as a three stage extraction/reaction process (Fig. 8). In the ﬁrst stage the biomass is pre-soaked in a 10 m long screw con-veyer and soaking tank (1) at temperatures up to 100 °C for resi-dence times up to 30 min. Water for the pre-soaking is obtained

Fig. 8. Continuous counter-current PLPW extraction plant for the fractionation of biomass: (1) soaking tank; (2) belt conveyer; (3) particle pump; (4) screw conveyer extractor; (5) particle pump; (6) screw conveyer extractor/reactor; (7) exhaust particle pump (adapted fromThomsen et al., 2006).

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from the water outlet of the extraction system. During this stage all of the air is driven from the biomass as it is saturated with the pro-cess water, in addition there is some transfer of thermal energy which increases the temperature of the biomass. After the pre-soak the biomass is transferred to a hopper on a belt conveyer (2) where it is fed into the second stage by means of the particle pump (3) mentioned previously. The second stage occurs in the extractor (4), which is a screw conveyer operated at elevated pressures and temperature up to 200 °C (Thomsen et al., 2006). The biomass moves upwards through the conveyer while the PLPW enters at the top of the conveyer and moves downward, counter-currently by gravity. The purpose of the second stage is the removal of most of the hemicellulose sugars (xylose and arabinose), which may be utilized for ethanol production by C5-sugar fermenting micro-organisms (Thomsen et al., 2008). Biomass exits the second stage into the third stage through another particle pump (5). In the third stage, which is a screw conveyer that acts as an extractor/reactor (6), the remaining biomass is treated at 195 °C at a residence time of 3 min with the addition of steam and water in the upper portion of the reactor, to improve the enzymatic convertibility of the remaining cellulose fraction for ethanol production. The remaining biomass exits the system through a ﬁnal particle pump. Water for the system enters at the end of the screw conveyer in the third stage and leaches through the system counter-currently to the ﬂow of biomass. Hemicellulose recovery in the system ranged from 33% to 83% depending on the amount of water added to the extractors (Thomsen et al., 2008). Yields of cellulose and hemicellulose could not be maximized concurrently as highly digestible cellulose was only obtained under severe conditions, which resulted in larger amounts of hemicellulose degradation and formation of inhibitory fermentation compounds. However, the system has shown the po-tential under such conditions to still have an estimated ethanol production capability of 184–192 kg ethanol/ton straw compared to the theoretical maximum if all cellulose and hemicellulose in the straw was hydrolysed to monomeric sugars of 318 kg etha-nol/ton straw.

6. Summary

There is little information in the literature about the effect of scale on the extraction process in pressurized fluid extractors. Of-ten when two scales of an extractor are examined there is no expli-cit discussion of the effects or the scale-up has not been done uniformly, leaving the effects of the scale-up on extraction in doubt. In general laboratory-scale extractions may be utilized to determine the mass transfer and phase equilibria data for a partic-ular extraction vessel and material. With appropriate models, this data may be used to examine the effects of scale. Caution should however be taken as the effects of scale-up of extractors may not be accounted for in most models. Homogeneous flow is a crucial assumption in most models of extraction but in larger extraction vessels radial variations in bed porosity and viscosity could pro-duce flow heterogeneity. This would result in higher interstitial velocities near the vessel wall, resulting in slower extraction in the centre of the extraction vessel. The goal in the design of pres-surized fluid extractors is not just to scale the process up, but to produce a system able to operate on a continual basis. There has been some success in this area by utilizing a battery of vessels in series to operate on a quasi-continuous basis, and with the use of screw conveyors to produce a gas-tight plug of material, which al-lows the extraction to operate at the necessary elevated pressures. Others have utilized the material itself to create a plug that can hold the pressure in the system to operate on a continual basis. De-sign of a suitable sluice system for charging solid materials be-tween zones of different pressure is another method that has

been used to achieve true continuous ﬂow. Most scale-up has oc-curred with the relatively older pressurized ﬂuid extraction tech-nology of SFE. The younger technologies of PLPW extraction and PLE are just now gaining popularity in the research community, and modeling and scale-up is sure to be forthcoming.

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