0263±8762/97/$10.00+0.00
q Institution of Chemical Engineers
THE INFLUENCE OF SPARGER DESIGN AND LOCATION
ON GAS DISPERSION IN STIRRED VESSELS
D. BIRCH and N. AHMED
Department of Chemical Engineering, The University of Newcastle, Australia
A
well designed agitated gas dispersion system should generate a large interfacial area
between the gas and the liquid phases, exhibit minimal in¯ uence of gassing on the
power draw, and not be prone to ¯ ooding. These characteristics are shown to be
strongly affected by the location of the sparger. Signi® cant performance improvements, in
terms of improved power draw and delayed onset of ¯ ooding on aeration, are achieved
through the use of `larger than impeller’ ring spargers, positioned within the discharge stream
from the impeller. There is little or no penalty in terms of the gas holdup generated. Results
for the Rushton and upward and downward pumping pitched blade impellers are reported.
The observed behaviour may be explained in terms of the loading regime of the impeller and
the cavity forms observed behind the blades, which is governed by the interaction between
the sparger (gas discharge) position and the ¯ ow patterns generated by each impeller.
Keywords: aerated power consumption; gas holdup; stirred vessels; sparger design;
gas-liquid reactors
INTRODUCTION
pumping pitched blade turbine (PBD). They found that
`wall-near’ sparge rings were unable to ¯ ood the stirrer and
suggested that this would lead to `a higher safety for vessel
design’ . This arrangement also resulted in substantially
improved suspension behaviour, probably due to the
relatively higher aerated power draw and postponement of
¯ ooding. Rewatkar et al.12 found the effect of the sparger
location with downward pumping turbines to be the `most
important parameter for the gas dispersion process’ . They
recommend a large diameter (2D) ring sparger for
suspension of solid particles at higher super® cial gas
velocities (vs > 9.4 mm s- 1 ). Moreover, placement of the
sparger further towards the bottom of the vessel and thus
away from the impeller decreased the critical impeller speed
for solid suspension in a gas-liquid-solid system. In some
similar work Rewatkar and Joshi13 found that the critical
speed for gas dispersion was minimized when a large
diameter (2D) ring sparger was used and placed away from
the impeller. Rewatkar et al.14 observed the power
consumption and the gas holdup generated by the PBD,
using a variety of sparger designs and positions. Again,
large diameter ring spargers (2D) located below the turbine
plane were recommended. This arrangement resulted in the
holdup being maximized with a stable power draw by the
impeller.
The power consumption instability found with the PBD
under steady operating conditions has been the focus of
some speculation. It would seem that two mechanisms are
responsible; the opposing ¯ ows of gas from the sparger and
the impeller discharge, and the transition, with changing gas
rate, from the impeller being directly loaded with gas from
the sparger, to indirect loading (gas enters the impeller
region through recirculation). In an earlier study investigating gas-liquid mixing with downward pumping mixed ¯ ow
Attaining effective contact between a dispersed gas and a
liquid is of great importance to the biochemical, chemical
and pharmaceutical industries. Gas-liquid dispersion
operations, utilizing sparged agitated vessels, typically
rely on the Rushton impeller (FDT) in conjunction with a
sparger positioned below the stirrer. The design of the gas
sparger and its location relative to the FDT has not been
considered critical, provided a ring sparger of diameter
smaller than the impeller and positioned within one
impeller diameter (below) of the turbine plane is used1 .
However, Nienow et al.2,3 demonstrated that the impeller
hydrodynamics may be markedly modi® ed by changing
the sparger design. Using large diameter ring spargers
(1.2D), positioned T/25 below the bottom of the FDT
blades, they detected clinging cavities behind the impeller
blade right up to the ¯ ooding transition, rather than the
usual large cavities. Thus, the system was found to have
decreased ¯ ooding susceptibility, higher relative aerated
power draw and, as a consequence of these effects, a higher
speci® c mass transfer coef® cient. There has been recent
interest in other impellers and arrangements. Indications
are that mixed and axial ¯ ow impellers have advantages
over the FDT in aerated stirred reactors in terms of
increased operational stability, improved gas-liquid-solid
dispersion performance and greater power ef® ciency4- 10 .
The gas dispersion performance of axial and mixed ¯ ow
turbines however, has been shown to be sensitive to the
design of the sparging system and further work is
necessary to resolve the optimum geometry6,7 .
Breucker et al.11 investigated the in¯ uence of various
geometrical arrangements, including sparger design, on the
two-phase and three-phase operational behaviour of a
number of agitators, including the FDT and the downward
487
488
BIRCH and AHMED
Figure 1. Schematic of the experimental arrangement.
turbines, Warmoeskerken et al.15 found that the transition
from indirect to direct loading coincided with the formation
of large cavities behind the impeller blades. The authors
could not determine which of the two phenomena was the
cause and which was the effect. Nevertheless, this transition
coincided with a large change in the vessel and impeller
hydrodynamics as well as the aerated power consumption. A
system exhibiting power consumption independence from
the gassing rate is desirable for a number of reasons. This is
especially true in gas-liquid-solid systems where a drop of
impeller power draw with aeration may be coupled with a
disastrous (and costly) settling out of the solids particles.
Frijlink et al.16 studied the gas-liquid-solid dispersion
performance of a PBD with a blade angle of 608 . A variety
of sparger-impeller arrangements were used in their work
and they linked the observed aerated power consumption
and impeller and vessel hydrodynamics to changes in the
geometry. They evince that large cavities ® rst form behind
the blades of these turbines from gas recirculated from
above, and that this is followed by a loss of pumping
capacity and the indirect to direct loading transition. This
results in a more radial out¯ ow from the impeller and
greatly altered vessel hydrodynamics. The operating conditions under which this transition was found to occur was
heavily in¯ uenced by the separation distance between the
sparger and impeller, with higher gassing rates being
necessary with larger separation distances.
Several distinct cavity forms have been reported for the
FDT with changes in ¯ ow number1 . Available literature
would indicate that similarities exist between the cavity
forms observed with the FDT and PBD, with some
important distinctions. At low ¯ ow numbers a single
`vortex’ cavity forms on each blade of the PBD1 . With
increasing ¯ ow number `clinging’ cavities result, progressing to `growing’ , and followed by `large’ cavities with
further increases in the gas rate1 . The `3-3’ structure has not
been reported for six bladed mixed ¯ ow turbines. Nienow et
al.6 described the cavity forms observed with a upward
pumping pitched blade turbine which were similar to those
reported for the PBD. Clinging cavities and large cavities
were observed but they state that the large cavities on each
blade were of roughly the same size.
Studies comparing the gas-liquid-solid dispersion performance of the FDT and upward and downward pumping
pitched blade turbines have shown the upward pumping
turbine to require less power to disperse gas and suspend
solids at high gassing rates (>1vvm). Further, little increase
in power is required up to 3.5 vvm, and the impeller has
comparatively stable power draw characteristics4- 9 . The
above investigations were conducted with the gas sparged
from either point spargers or small diameter ring spargers
(Ds < D) placed below the turbine. Parthasarathy et al.10
studied the gas holdup generated by the FDT and pitched
blade disc turbines pumping both upwards and downwards
(PDU and PDD respectively) in a non-coalescing air-water
system. The PDU was found to generate much higher gas
holdup values under similar operating conditions. However,
there seems to be no available information on the in¯ uence
of sparger location on the gas dispersion performance of
upward pumping mixed ¯ ow turbines.
In summary, it may be concluded that `larger than
impeller’ may lead to desirable hydrodynamics in gas-liquid
mixing operations, indicated by observations made with the
FDT and PDD. Available information however, is erratic,
and no effort has been made to explain the observations. The
aim of this study is to evaluate the in¯ uence of large
diameter ring spargers and their position on the gas
dispersion performance, especially the gas holdup, of the
PDU and PDD and compare their behaviour with the FDT.
EXPERIMENTAL
Figure 1 shows the experimental arrangement used in this
study. A ¯ at-bottomed 0.6 m diameter (T) cylindrical
baf¯ ed vessel, ® lled with tap water to a height equal to
the tank diameter (H = T ) was used. The FDT and mixed
¯ ow turbine (PDT) used in this study are shown in Figure 2.
The mixed ¯ ow turbine had six blades mounted on a central
disc at 458 from the horizontal. The pumping orientation of
this turbine was reversed by changing the direction of
Trans IChemE, Vol 75, Part A, July 1997
INFLUENCE OF SPARGER DESIGN AND LOCATION ON GAS DISPERSION
489
Figure 3. Schematic of the sparger size and location with respect to the
impeller position. Impeller diameter is 200 mm. (Note: the nomenclature a,
b, 1 in italics refer to positions above, below and level to the impeller plane,
respectively). The symbols shown are consistent throughout the text.
Figure 2. Details of the impeller geometry.
rotation of the stirrer. The impeller diameter (D) and the
impeller off-bottom clearance (C) were 0.2 m (T/3) for all
experiments. Figure 3 details the sparger sizes and locations
with respect to the impeller. It may be noted that the
diameter of Ring3 approximates to 1.4 times the impeller
diameter. A simple calculation will show that the volume
swept by the impeller, i.e. the volume enclosed between the
axis and the outer radius of the impeller, is numerically
equivalent to the volume enclosed in the annulus between
the outer impeller radius and 1.41 times the outer radius.
From photographic evidence of gas dispersion in the
impeller region reported by van’ t Riet and Smith17 and
Bruijn et al.18 , one may conclude that the bubble breakup
occurs well within a volume between the impeller tip and
1.3 to 1.4 times the impeller outer radius. This is
corroborated by the work of Parthasarathy et al.19 , who
demonstrated experimentally that this probably is the region
in which all gas breakage into bubbles take place. Thus, if
one is moving towards the impeller, Ring3 represents a
location, at and beyond which, the ¯ ow and turbulence
generated by the impeller are at the strongest.
All ring spargers had 163 ´ 0.0015 m diameter holes. The
spargers were positioned concentrically to the impeller at
three vertical locations, 0.10 m below, level with, and
0.05 m above the turbine plane. With the exception of the
spargers positioned level with the impeller, which released
gas upwards, all spargers discharged gas towards the turbine
plane (Figure 3). The point sparger discharged gas at a
location 0.10 m below the impeller plane, through a 0.019 m
diameter ori® ce. The super® cial gas velocity, v s was varied
from 0 to 0.027 m s- 1 (0 to 2.7 vvm).
Trans IChemE, Vol 75, Part A, July 1997
The main aim of studying the effect of sparger
con® guration on the gas dispersion performance of the
three impellers was met by measuring, at a particular
impeller speed and gassing rate, the power consumption and
the gas holdup. Experiments were performed at a constant
impeller speed with incremental changes in the gas ¯ ow
rate. At each setting the gas holdup was calculated from the
dispersion height which was measured by an ultrasonic level
sensor. The impeller speed was monitored by a magnetic
proximity sensor. Power consumption was calculated from
readings from a shaft mounted torque transducer. Torque,
level and impeller speed readings were monitored and
recorded on a personal computer (Figure 1). Observations
on the power draw stability, the bulk ¯ ows and the
uniformity of the dispersed gas in the vessel were also
made. As any change in the aerated power behaviour due to
sparger design was expected to be due to changes in the
impeller hydrodynamics, a high speed video recorder was
used to determine the type of cavity on the impeller blades.
Images of the impeller region could be obtained both
through the base and through the side walls of the vessel.
Particular attention was also paid to the detection of the
¯ ooding transition. The visual determination of the onset of
¯ ooding is, by nature, very subjective. In these experiments
the transition illustrated by Warmoeskerken20 , which is
characterized by a cessation of effective pumping by the
stirrer, is used.
RESULTS AND DISCUSSION
The relationship between the aerated power ratio (Pg / Pu )
and the ¯ ow number (Fl = Q/ ND3 ) is peculiar to a particular
tank and impeller geometry, and is commonly used to
490
BIRCH and AHMED
Figure 4. (a) Aerated power consumption and (b) the gas holdup as a
function of the ¯ ow number for sparging arrangements placed below the
impeller with the Rushton impeller (FDT), at a stirrer speed of 6.1 s- 1 .
Figure 5. (a) Aerated power consumption and (b) the gas holdup as a
function of the ¯ ow number for sparging arrangements placed level with the
impeller for the Rushton impeller (FDT), at a stirrer speed of 6.1 s- 1 .
interpret the changes in the impeller hydrodynamics with
aeration (cavity forms behind the blades). The correspondence between the aerated power and the cavity forms are
reasonably well established, at least from the FDT1 . The
above approach has been found to be justi® ed in interpreting
the aerated power draw behaviour with changes in the
sparger design. Results are presented separately for the three
impellers. The values for the gas holdup (gas void fraction)
are also reported to indicate the degree of gas dispersion at
each stage.
vortex, followed by clinging cavities, were detected for up
to the highest gas rate used. Similar behaviour is observed
for spargers positioned level to the turbine (Ring3l, Ring4l),
as shown in Figure 5. This would explain the better solids
suspension behaviour observed by Breucker et al.11 with
`wall-near’ spargers. None of the arrangements used were
observed to ¯ ood with the aeration rates used at the impeller
speed of 6.1 s- 1 . Flooding was observed however, at lower
stirrer speeds (for example, at 4.6 s- 1 ) with the Ring2b,
Ring1b and Point systems only, which agrees with the
® ndings of Nienow et al.2,3 .
The cavity forms generated depends on the amount of gas
which reaches the impeller for each arrangement, which in
turn depends on the path by which it does so. This becomes
apparent when the ¯ ow streams generated by the impeller is
studied in conjunction with the sparger arrangements.
Figure 6a shows the ¯ ow patterns due to the FDT,
determined visually by using neutral buoyancy ¯ ow
followers and ® ne bubbles. Thus, the arrangements which
can only load the impeller indirectly (Ring3l, Ring4l, and
Ring4b) exhibit clinging and vortex cavities, whereas gas
from the Point, Ring1b and Ring2b geometries rises directly
into the turbines, generating the larger cavities with
increasing gas ¯ ow rates. For the former group of spargers
(DS > D), an inspection of Figures 4 and 5 reveals another
interesting feature in that the aeration effect on the power
draw is dependent on the distance from impeller. Thus, the
aerated power becomes progressively independent of the
aeration as the sparger to impeller distance increases. For
these arrangements the gas is not able to reach the impeller
directly, and the quantity that does through recirculation
becomes progressively smaller as the sparger is moved
away from the impeller. For example, with Ring4b it was
noticed that very little gas reached the impeller, with a large
Flat Blade Disc (Rushton) Impeller, FDT
It will be noticed from Figure 3 that the sparger
con® gurations used in conjunction with this impeller are
either placed below or level with it. To avoid confusion the
sets are shown separately. Thus, Figure 4 shows the
con® gurations below while Figure 5 shows the arrangements level with the impeller, at an arbitrarily chosen stirrer
speed of 6.1 s- 1 . The point sparger (standard con® guration)
is included in each ® gure to highlight the deviation from the
most commonly used geometry. When the impeller is
sparged with gas from below, for the small diameter
(DS # D) ring spargers (Ring1b, Ring2b) or a point sparger
(Point), the aerated power draw drops to approximately
30% of the unaerated value (Figure 4), similar to those
reported in the literature. The cavities, determined by video
imaging, closely matched those described by Smith1 . At low
aeration numbers only vortex cavities could be observed,
followed by clinging and the 3-3 structure (clinging and
large cavities on alternate blades) with progressive increase
in the ¯ ow number. In contrast, with the larger than the
turbine sparger (Ring4b), the power drawn by the impeller is
seen to be remarkably independent of the gassing rate. Only
Trans IChemE, Vol 75, Part A, July 1997
INFLUENCE OF SPARGER DESIGN AND LOCATION ON GAS DISPERSION
491
Figure 6. The ¯ ow patterns generated by the impellers investigated: (a) Rushton impeller (FDT); (b) pitched blade disc impeller pumping upward (PDU);
(c) pitched blade disc impeller pumping downward (PDD). The sparger positions used are shown in the background.
amount of gas escaping the tank with little or no circulation, consistent with the interaction between this sparger
and ¯ ow generated by this impeller. From visual evidence
it also appeared that as the sparger is brought nearer to
the impeller, progressively smaller bubbles are generated,
leading to higher amounts of gas reaching the impeller
by recirculation. Thus, the Ring3l system being closest to
the impeller, generates the smallest bubbles, which results
in greater gas recirculation. Although the cavity form
remains in the clinging regime, more gas was observed
behind the blades (large clinging cavity). This incidentally,
is also the position which generates the best gap holdup.
It will be noticed that the drop off in holdup for Ring3l,
Ring4l and Ring4b is in the reverse order when compared
with their aerated power draw. As observed earlier, when
the sparger is placed away from the impeller the gas is
only able to reach the impeller through recirculation, the
amount decreasing with increasing sparger distance from
the impeller. Although this has bene® cial effects as far
as the aerated power and ¯ ooding are concerned, a
penalty is paid in terms of the gas holdup. Thus, the worst
holdup values is obtained with Ring4b. With the DS # D
arrangements (Point, Ring1b and Ring2b), although large
cavities are formed and power drop occurs with increasing
aeration, most of the gas come in contact with the impeller
zone, thus ensuring better dispersion around the tank. As
expected, however, the bubbles observed to emerge from
the larger cavities appeared to be larger than those generated
from vortex and clinging cavities.
Pitched Blade Disc Turbine Pumping Upward (PDU)
Figures 7 and 8 show the aerated power consumption and
gas holdup characteristics for the PDU, for spargers below
and above the impeller, respectively, at the same impeller
speed as the FDT. Figure 6b provides the ¯ ow patterns
determined for this impeller. The pattern of behaviour is
similar to that observed with the FDT. Thus, when the gas is
introduced directly below the impeller (Point, Ring1b and
Ring2b), the power consumed by the impeller drops
dramatically with aeration rate, ultimately levelling off at
less than half its unaerated value. Flooding of the impeller
takes place at relatively low ¯ ow rates, and the points are
indicated by a sudden drop in the power consumption
(shown by the dotted lines in Figure 7a). This leads to an
accompanying drastic drop in the gas holdups, as indicated
Trans IChemE, Vol 75, Part A, July 1997
in Figure 7b. These results are not unexpected as the
impeller is always directly loaded, and the gas and impeller
discharge is in the same direction (Figure 6b). Video
imaging of the PDU-Point sparger system shows clinging
cavities to be present at low ¯ ow numbers (Fl < 0.015).
With increasing gas ¯ ow rate, large cavities form and a
corresponding drop in the aerated power consumption is
seen. The `3-3’ cavity structure was not observed. Similar
cavity forms were found with the Ring1b and Ring2b
sparger systems. Interestingly, the larger Ring2b is observed
to lead to ¯ ooding before Ring1b. Thus, Ring1b is observed
to ¯ ood the impeller at Fl < 0.095, while the transition takes
place at a lower Fl of around 0.075 with Ring2b. This
seemingly anomalous behaviour, also observed by Bruijn
et al.18 , was explained by the authors in terms of the
mechanism for ¯ ow of gas from the sparger into the cavity.
With Ring1b, gas rises into the disc and is sucked into the
cavity from there. The authors include a photographic plate
which shows this mechanism clearly. Gas leaving Ring2b on
the other hand enters the cavity directly and it is likely that
this results in the ¯ ooding transition occurring with this
system before the Ring1b arrangement. Figure 6b, which
shows the ¯ ow patterns determined for this impeller, tends
to support this contention. For the sparger away from the
impeller (Ring4b), the in¯ uence of the aeration rate on the
aerated power consumption is small, and only vortex or
clinging cavities were detected. As seen from Figure 6b, this
is expected as the impeller may only be loaded indirectly,
with very little gas expected to reach the impeller in this
case.
When gas is introduced above the turbine plane (Ring1a±
4a), the in¯ uence on the aerated power is less than that
observed with the bottom spargers, as the impeller is not
loaded directly. Except for Ring1a, as the sparger moves
away from the impeller, the effect of aeration progressively
decreases due to smaller amounts of gas reaching the
impeller (through progressively larger circulation loops).
This is consistent with the bulk ¯ ow patterns, and no
¯ ooding is observed even though relatively high gas rates
are encountered (2.7 vvm). Thus, the Ring2a arrangement
shows a slightly larger dependency on aeration rate than
Ring3a, while as with the FDT, for the Ring4b system, very
little gas reached the impeller, and the gas was not well
dispersed. Obviously, the gas holdup shows the opposite
trend. Larger amounts of gas are progressively entrained as
the sparger arrangement moves towards the impeller due to
492
BIRCH and AHMED
Figure 7. (a) Aerated power consumption and (b) the gas holdup as a
function of the ¯ ow number for sparging arrangements placed below the
impeller with the pitched blade disc impeller pumping upward (PDU), at a
stirrer speed of 6.1 s- 1 . The dotted lines indicate the point of the ¯ ooding
transition.
Figure 8. (a) Aerated power consumption and (b) the gas holdup as a
function of the ¯ ow number for sparging arrangements placed above the
impeller with the pitched blade disc impeller pumping upward (PDU), at a
stirrer speed of 6.1 s- 1 . The dotted lines indicate the point of the ¯ ooding
transition.
the shortening circulation loops. In addition, due to the high
turbulence and strong ¯ ows progressively experienced, the
bubble size generated is also affected, a factor which needs
further investigation. Overall therefore, as indicated by
Figures 7 and 8, the Ring2a sparger generates the highest
holdup with the PDU. Figure 6a shows that this is also the
position where the bulk ¯ ow streams from the impeller
emanate. The behaviour of Ring1a is at odds with the other
above impeller spargers in that it leads to ¯ ooding at
relatively low Fl, in spite of being indirectly loaded.
Flooding in this case may be explained in terms of a gas
driven bulk ¯ ow sucking suf® cient liquid upward through
the turbine due to its close proximity to the sparger, as to
in¯ uence its pumping ef® ciency. Moreover, it was observed
that the bulk ¯ ow generated by the impeller is such that gas
is easily entrained in substantial amounts, albeit indirectly,
from smaller (secondary) ¯ ow loops, into the main stream
going into the impeller (Figure 6b). This leads to rapid
growth in cavity size and drop in aerated power with
increasing Fl. The formation of larger clinging cavities were
evident prior to ¯ ooding.
presence of the `3-3’ cavity formation prior to ¯ ooding21 . A
careful scrutiny of the high speed video recordings clearly
indicated such a cavity formation, although the `3-3’
structure has not been reported for the PDD before.
Interestingly, they were not observed with the PDU. The
Ring1b and Ring2b arrangements exhibit a signi® cant
dependency on aeration rate, although, due to their location,
they would be expected to load the PDU indirectly. With
these spargers however, it was observed that the ¯ ow pattern
was such that a larger volume of gas entered the impeller
from the side and above the impeller (small circulation
loops formed (Figure 6c), as with the PDU), leading to the
formation of larger cavities (pseudo-direct-loading). As
indicated in Figure 6c, the intensity of agitation in the zone
below the impeller is signi® cantly higher than that obtained
with the PDU. This leads to a comparatively complicated
aerated power curve for the DS # D spargers. With Ring2b,
the power drawn by the PDD initially drops with increasing
aeration rate until levelling off to some extent at Fl < 0.015.
The turbine was indirectly loaded with gas at this point and
clinging cavities were observed. With further gassing,
another distinct drop in power draw is seen followed by
another levelling off period (Fl < 0.06). This event coincides with the indirect-direct loading transition and the
formation of large cavities (Figure 9). This system is then
found to ¯ ood at Fl < 0.09 , with perceptible drop in the gas
holdup. Frijlink et al.16 reported similar trends for a PBD
sparged with gas from below by a ring sparger at a large
vertical impeller-sparger separation (s/ T = 0.25, as compared to s/ T of 1/6 used in this work). They related the initial
levelling off in power, at low gassing rates, to the
postponement of the growth of clinging cavities, and
Pitched Blade Disc Turbine Pumping Downward (PDD)
When the gas is introduced directly below the PDD
(Point, Ring1b and Ring2b), as shown in Figure 9, there is a
large drop in power draw at low ¯ ow numbers (Fl < 0.01).
The indirect to direct loading transition being visually
con® rmed for the Ring1b system at this stage. A positive
step change in the aerated power consumption accompanies
the ¯ ooding transition (Figure 9). In the case of the FDT a
positive step change at ¯ ooding would be indicative of the
Trans IChemE, Vol 75, Part A, July 1997
INFLUENCE OF SPARGER DESIGN AND LOCATION ON GAS DISPERSION
493
Figure 9. (a) Aerated power consumption and (b) the gas holdup as a
function of the ¯ ow number for sparging arrangements placed below the
impeller with the pitched blade disc impeller pumping downward (PDD), at
a stirrer speed of 6.1 s- 1 . The dotted lines indicate the point of the ¯ ooding
transition.
Figure 10. (a) Aerated power consumption and (b) the gas holdup as a
function of the ¯ ow number for sparging arrangements placed above the
impeller with the pitched blade disc impeller pumping downward (PDD), at
a stirrer speed of 6.1 s- 1 . The dotted lines indicate the point of the ¯ ooding
transition.
hence the direct loading transition. It was assumed that the
cavity growth was inhibited as a result of a lack of bubbles
in the ¯ ow loops close to the stirrer.
Direct loading was found to occur only after the holdup in
the sparger zone had increased, leading to increased cavity
size and a reduction in impeller power draw. The ¯ ow
patterns for this impeller does provide support for that
argument with the Point and Ring1b arrangements, but does
not offer an explanation for non-¯ ooding with the Ring1b
con® guration. With DS > D spargers (Ring3b and Ring4b),
the effect of aeration on the power is reduced, and more gas
is dispersed to above the impeller region. This happens more
effectively with Ring3b which lies in the discharge stream
of the impeller. In addition, due to the ¯ ow pattern,
reasonable amount of gas is entrained into the impeller to be
dispersed, but not enough to form large cavities. Thus, this
sparging con® guration generates the maximum gas holdup
with this impeller. As would be indicated by the follow
streams, little gas reaches the impeller with the Ring4b
design, with most of the gas leaving the tank with some
circulation in the top half of the vessel. This results in poorer
gas holdup.
The experimental results for the spargers located above
the impeller with the PDD (Ring1a, Ring2a and Ring4a,
Figure 10) are subject to easy interpretation with the use of
Figure 6c. The ¯ ow patterns dictate that the gas may enter
the impeller region only by entrainment from ¯ ow streams
moving into the impeller from above with these arrangements. Figure 6c indicates that such ¯ ow is concentrated in a
region surrounding the impeller blades, or at an impeller
diameter distance away from the centre. Thus, Ring2a draws
in the maximum amount of gas followed by the Ring1a
arrangement. Ring4a due to its position is able to draw little
gas into the impeller, generates poor gas holdup and only
clinging cavities were observed behind the impeller blades.
Large cavities were detected with the Ring1a and Ring2a
con® gurations. The cavities were all observed to be of
similar size, in contrast to the 3-3 structure exhibited by the
spargers placed below with this impeller con® guration. The
observation is puzzling and is probably related to the role of
the central disc in the distribution of gas to the blades of the
impeller.
In general, observing the correspondence between the
aerated power draw and the gas holdup, it seems that the
best outcome in terms of both is obtained when the impeller
is indirectly loaded, and the amount of gas ¯ ow into the
impeller should be such that formation of large cavities is
avoided. Although vortex and clinging cavities may be
maintained at large gas rates by placing the sparger near the
walls, there is progressive deterioration in terms of gas
holdup as more and more gas leaves the system without
dispersion. A balance seems to be reached by placing the
sparger in the output stream of the impeller, within the
sphere of in¯ uence of the impeller generated turbulence.
With the FDT, the discharge from the impeller is in the
radial direction and thus the best performance in terms of
the gas holdup is obtained with Ring3l (level with impeller)
arrangement. As mentioned in the experimental section, this
position represents the outer edge of the zone in which the
impeller generated turbulence has been shown to be the
strongest. It would be dif® cult to place the sparger any
nearer to the impeller due to physical limitations. For the
PDD, location 3b satisfy the criteria of being in the output
stream of the impeller and within the in¯ uence of the
Trans IChemE, Vol 75, Part A, July 1997
494
BIRCH and AHMED
Table 1. Percent torque ¯ uctuation; deviation from the time averaged value.
Sparger: Point
FDT
Gas rate (vvm)
0.12
0.30
0.61
PDU
PDD
0.3 W/kg
1.4 W/kg
0.3 W/kg
1.4 W/kg
0.3 W/kg
1.4 W/kg
5.365
4.939
7.333
2.145
2.952
4.371
5.044
5.543
8.005
2.372
2.833
3.135
7.532
7.031
12.828
4.078
3.824
9.721
Sparger: Ring2b
FDT
Gas rate (vvm)
0.12
0.30
0.61
PDU
PDD
0.3 W/kg
1.4 W/kg
0.3 W/kg
1.4 W/kg
0.3 W/kg
1.4 W/kg
4.862
4.142
5.388
1.981
3.898
3.101
5.116
4.903
4.745
1.745
2.394
2.134
9.470
18.434
9.144
3.062
8.353
9.165
Sparger: Ring3
FDT, Ring3l
Gas rate (vvm)
0.12
0.30
0.61
PDU, Ring3a
PDD, Ring3b
0.3 W/kg
1.4 W/kg
0.3 W/kg
1.4 W/kg
0.3 W/kg
1.4 W/kg
5.078
4.645
5.130
2.472
2.902
4.155
4.670
4.368
4.966
2.224
2.341
2.551
5.510
8.806
6.276
2.631
2.905
16.674
Sparger: Ring4
FDT, Ring4l
Gas rate (vvm)
0.12
0.30
0.61
PDU, Ring4a
PDD, Ring4b
0.3 W/kg
1.4 W/kg
0.3 W/kg
1.4 W/kg
0.3 W/kg
1.4 W/kg
3.373
3.088
4.100
1.761
2.500
3.187
4.198
5.114
4.351
1.903
1.920
2.149
5.530
5.733
8.066
2.744
3.078
7.808
impeller generated turbulence. The same rationale is valid
for the PDU with the Ring2a arrangement. In addition to
generating large holdups, placing the sparger in the
discharge stream of the impellers may have the added
advantage of producing smaller bubbles, and it may be
speculated that the gas holdup thus generated would have a
higher surface area for the same volume, leading to greater
gas-liquid mass transfer. This aspect needs to be studied
further.
Power Draw Stability
A gas dispersion system in which power draw ¯ uctuations
are observed under otherwise steady operating conditions is
undesirable. This is true for several reasons, in particular,
increased equipment were and process performance ¯ uctuations. No gross power draw ¯ uctuations or ¯ ow instabilities
were observed with any of the sparging arrangements used
in combination with the PDU or FDT. This is not
unexpected as the FDT is noted for its stable gas dispersion
operation, and any gas entering the impeller region of the
PDU has the tendency to rise, thus complementing the
upward ¯ ow from the impeller.
Power draw ¯ uctuations were evident with the PDD. The
instability associated with the indirect to direct loading
transition was observed with the Point, Ring1b and Ring2b
systems. The large diameter spargers portioned below the
turbine plane (Ring3b and Ring4b) had perhaps the best
operating characteristics of the arrangements tested with the
PDD. Both exhibited good aerated power draw characteristics with increasing ¯ ow number. However, some power
draw and ¯ ow instabilities were observed, despite the
avoidance of the direct loading state. Indications are that
this is due to the opposing ¯ ows of agitator discharge and
recirculated gas rising into the impeller region from below.
In an effort to quantify the torque instability in operating
systems, the percentage ¯ uctuations in torque from the
mean were recorded at steady state operating conditions for
various con® gurations, and the tabulated ® gures are shown
in Table 1. Results are shown for the more conventional
sparging arrangements (Point and Ring2b) as well as for
systems with superior gas dispersion characteristics
(Ring3a and Ring4a with the PDU; Ring3b and Ring4b
with the PDD; Ring3l and Ring4l with the FDT). Each data
point represents the normalized measure of spread (%
coef® cient of variation) of the impeller torque calculated
from more than ninety consecutive readings. In every case
the ¯ uctuations for the PDD are greater than those obtained
with the PDU or FDT for the corresponding conditions,
which are in agreement with the qualitative observations. As
Trans IChemE, Vol 75, Part A, July 1997
INFLUENCE OF SPARGER DESIGN AND LOCATION ON GAS DISPERSION
expected the usage of large diameter ring spargers (Ring3b
and Ring4b) with the PDD minimizes the power draw
¯ uctuations.
CONCLUSIONS
It has been demonstrated that the sparger dimensions and
location with respect to the impeller has signi® cant effect on
the gas dispersion characteristics in an aerated stirred vessel.
In broad terms, larger than impeller spargers which lead to
indirect loading of the impeller offer superior operational
alternatives for gas-liquid systems, with implication for
three phase operations. This is because indirect loading, in
turn, hinders the formation of large cavities behind the
impeller blades, thus ensuring relatively low power loss
with aeration. Clinging cavities may be preserved at
progressively higher ¯ ow numbers by moving the sparger
nearer to the vessel wall. However, a balance needs to be
established, considering that beyond a certain point the
in¯ uence of the impeller is insuf® cient to maintain good
overall dispersion of the gas, thus reducing the gas holdup.
Experimental results show that the most suitable location
for introducing the gas appears to be in the output stream
from the impeller, a ® nding supported by the observed ¯ ow
patterns for the impellers studied. With the PDU and PDD,
larger than impeller ring spargers placed within the
in¯ uence of the impeller discharge meet the above
conditions, and thus provide the best gas dispersion
characteristics in terms of improved gas holdup and good
aerated power draw. The direction of ¯ ow from these
impellers dictates that the sparger be placed above the
impeller for the PDU, and below for the PDD. With the
FDT, a sparger placed level to the impeller, and close
enough to be in¯ uenced by the impeller discharge stream
provide the best gas holdup. As expected, this arrangement
also results in good aerated power consumption characteristics when measured against the more conventional
arrangements of sparging from below. Of all the sparger/
impeller arrangements examined, the FDT generates the
maximum holdup, most probably as it also draws the
maximum power.
In practical terms, the investigation has shown that
sparger design deserves to be given greater attention in the
design of aerated stirred vessels than hitherto granted.
Although qualitative in nature, the general trends observed
from this study give valuable clues towards making
reasoned guesses. This would be especially true for three
phase operations, where power loss on aeration may lead to
potentially costly settling out of the suspended solids. It is
imperative that the current work be extended to consider
associated factors like the bubble size, gas-liquid interfacial
area, mass transfer rates and liquid phase mixing time to
allow a quantitative evaluation of the sparger con® gurations. Studies involving simultaneous solid suspension and
gas dispersing capabilities of these systems would be
valuable from an industrial perspective, and are in progress.
NOMENCLATUR E
PBD
PDD
PDT
PDU
downward pumping pitched blade turbine
downward pumping 458 six bladed pitched blade disc turbine
458 six bladed pitched blade disc turbine
upward pumping 458 six bladed pitched blade disc turbine
Trans IChemE, Vol 75, Part A, July 1997
C
CS
D
DS
FDT
Fl
H
N
Pg
Pu
Q
q
r
s
T
vs
vvm
w
495
impeller off-bottom clearance, m
sparger off-bottom clearance, m
impeller diameter, m
sparger diameter, m
Rushton turbine
¯ ow number (Q/ ND3 ), dimensionless
liquid height in vessel, m
impeller speed, s- 1
power draw, gassed liquid, W
power draw, ungassed liquid, W
volumetric gas ¯ ow rate, m3 s- 1
blade width (see Figure 1), m
blade length extending past the disc (see Figure 1), m
vertical separation between impeller and sparger, m
tank diameter, m
super® cial gas velocity, m s- 1
volume gas/volume liquid/min
blade length (see Figure 1), m
Greek letters
a
blade angle, degrees
eg
gas holdup, %
q
liquid phase density, kg m- 3
REFERENCES
1. Smith, J. M., 1985, Dispersion of gases in liquids: The hydrodynamics
of gas dispersion in low viscosity liquids, in Mixing of Liquids by
Mechanical Agitation, (J. J. Ulbrecht and Patterson, G. K., eds.)
(Gordon and Breach Science Publishers, New York).
2. Nienow, A. W., Haozhung, W., Huoxing, L. and Allsford, K. V., 1986,
The advantage of using large ring spargers with standard Rushton
turbines in gassed reactors, in Proc World Congress III of Chemical
Engineering, Volume 2, pp. 354±357.
3. Nienow, A. W., Huoxing, L., Haozhung, W., Allsford, K. V., Cronin,
D. and Hudcova, V., 1988, The use of large ring spargers to improve
the performance of ferments agitated by single and multiple standard
Rushton turbines, Proc 2nd Int Conf on Bioreactor Fluid Dynamics,
Cambridge, England, (BHRA Fluid Eng., Cran® eld, England) pp.
159±177.
4. Chapman, C. M., Nienow, A. W., Cooke, M. and Middleton, J. C.,
1983, Particle-gas-liquid mixing in stirred vessels. Part II: Gas-liquid
mixing, Chem Eng Res Des, 61: 82±95.
5. Chapman, C. M., Nienow, A. W., Cooke, M. and Middleton, J. C.,
1983, Particle-gas-liquid mixing in stirred vessels. Part III: Three phase
mixing, Chem Eng Res Des, 61: 167±181.
6. Nienow, A. W., Konno, M. and Bujalski, W., 1985, Studies on threephase mixing: a review of recent results, in Proc 5th Eur Conf on
Mixing, Wurzburg, West Germany, (BHRA Fluid Eng., Cran® eld,
England) pp. 1±13.
7. Nienow, A. W., Konno, M. and Bujalski, W., 1986, Studies on threephase mixing: a review and recent results, Chem Eng Res Des, 64: 35±
42.
8. Bujalski, W., Konno, M. and Nienow, A. W., 1988, Scale-up of 458
pitch blade agitators for gas dispersion and solid suspension, in Proc
6th Eur Conf on Mixing, Pavia, Italy, (BHRA Fluid Eng., Cran® eld,
England) pp. 389±398.
9. Bujalski, W., Nienow, A. W. and Huoxing, L., 1990, The use of upward
pumping 458 pitched blade turbines impellers in three-phase reactors,
Chem Eng Sci, 42(2): 415±421.
10. Parthasarathy, R., Ahmed, N. and Jameson, G. J., 1992, The effect of
impeller type on gas holdup in stirred vessels, Trans IChemE, Part C
70(C1): 37±43.
11. Breucker, Ch., Steiff, A. and Weinspach, P.-M., 1988, Interactions
between stirrer, sparger and baf¯ es concerning different mixing
problems, in Proc 6th Eur Conf on Mixing, Pavia, Italy, (BHRA
Fluid Eng., Cran® eld, England) pp. 399±406.
12. Rewatkar, V. B., Raghava Rao, K. S. M. S. and Joshi, J. B., 1991,
Critical impeller speed for solid suspension in mechanically agitated
three-phase reactors. 1. Experimental part, Ind Eng Chem Res, 30:
1770±1784.
13. Rewatkar, V. B. and Joshi, J. B., 1993, Role of sparger design on gas
dispersion in mechanically agitated gas-liquid contactors, Canad J
Chem Eng, 71: 278±291.
14. Rewatkar, V. B., Deshpande, A. J., Pandit, A. B. and Joshi, J. B., 1993,
Gas hold-up behaviour of mechanically agitated gas-liquid reactors
496
15.
16.
17.
18.
19.
20.
21.
BIRCH and AHMED
using pitched blade down¯ ow turbines, Canad J Chem Eng, 71: 226±
237.
Warmoeskerken, M. M. C. G., Speur, J. and Smith, J. M., 1984, Gasliquid dispersion with pitched blade turbines, Chem Eng Commun, 25:
11±29.
Frijlink, J. J., Kolijn, M. and Smith, J. M., 1984, Suspension of solids
with aerated pitched blade turbines, in Fluid Mixing II; IChemE Symp
Series No. 89, 49±58.
van’ t Riet, K. and Smith, J. M., 1973, The behaviour of gas-liquid
mixtures near Rushton turbine blades, Chem Eng Sci, 28: 1031.
Bruijn, W., van’ t Riet, K. and Smith, J. M., 1974, Power consumption
with aerated Rushton turbines, Trans IChemE, 52: 88±104.
Parthasarathy, R., Jameson, G. J. and Ahmed, N., 1991, Bubble
breakup in stirred vessels Ð predicting the Sauter mean diameter, Trans
IChemE, Part A, 69(A5): 295±301.
Warmoeskerken, M. M. C. G., 1986, Gas-liquid dispersion characteristics of turbine agitators, PhD Thesis, (Delft Technical University,The
Netherlands).
Nienow, A. W., Warmoeskerken, M. M. C. G., Smith, J. M. and Konno,
M., 1985, On the ¯ ooding/loading transition and the complete dispersal
condition in aerated vessels agitated by a Rushton-turbine, in Proc 5th
Eur Conf on Mixing, Wurzburg, West Germany, (BHRA Fluid Eng.,
Cran® eld, England) pp. 143±154.
ACKNOWLEDGEMENTS
The ® nancial contribution and technical support provided by Metquip
Pty. Ltd., Sydney, Australia is gratefully acknowledged. DB would also like
to express his sincere appreciation for support through the DEET APA(I)
scholarship program.
ADDRESS
Correspondence concerning this paper should be addressed to Dr N.
Ahmed, Department of Chemical Engineering, University of Newcastle,
NSW 2308, Australia. Fax: + 61 4921 6920.
The manuscript was received 31 May 1996 and accepted for publication
after revision 4 April 1997.
Trans IChemE, Vol 75, Part A, July 1997