Assessment Submission Form

Student Name Gillian Ryan

Student Number 07366302

Assessment Title Literature Review

Module Title Fluid Mechanics and Mixing in Agitated Reactors

Module Co-ordinator Dr. Niall English, Dr Frank Mac Loughlin

Tutor (if applicable) Dr Frank Mac Loughlin

Date Submitted 05 January 2011OFFICE USE ONLY

Date Received

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Signed……Gillian Ryan………………. Date……05/01/2011…………

University College Dublin

School of Chemical and Bioprocess Engineering

Fourth Year Chemical Engineering

Literature Review

Date of Submission: 5th January 2011

to

Dr. Frank Mc Loughlin

Name: Gillian RyanStudent Number: 07366302

Fluid Mechanics and Mixing in Agitated Reactors

Introduction

This review gives a description of the performance of gas dispersion impellers. For many years the Rushton turbine was the standard impeller choice for gas dispersion applications. This report will discuss the benefits of the application of a gas dispersion impeller which consists of vertically asymmetric blades. Disc-style gas dispersion impellers, that consisted of blades which are symmetric with respect to the plane of the disc, was not necessarily the optimum selection, since the gas usually enters from the bottom of the vessel, causing a distinctly asymmetric flow pattern.

The new impeller design, which is discussed in this report, is to accommodate the different flow conditions which exist above and below the impeller disc, opposed to gas dispersion impellers which have been presented in literature that contained blades which are symmetric with respect to the plane of the disc.

This review presents a discussion of the characteristics which are important in the design of a suitable gas dispersion impeller such as; the use of Stirred Tank Reactors in industries, the Effect of Impeller design on mixing, the Energy Dissipation rates in Stirred Systems, the Effect of gas on Power draw of selected impellers and on the fluid mechanics of two-phase dispersion, the Heat and Mass Transfer, and finally, the Scale-up of vessels from lab-scale to industrial scale.

1. Stirred Tank Reactors;

widely used contacting geometry

Stirred Tank Reactors and stirred vessels are widely used in the Chemical, Pharmaceutical, food industries and also in waste water treatment plants (e.g. municipal and industrial WWTPs). STRs are also used in metallurgical process industries. Over this range of processes, there is a range of different mixing and quality requirements. These alternative mixing requirements include the blending of different viscosity liquids, heat transfer and dispersion applications based on the consistency of the mixing material (e.g. liquid-liquid, solid-liquid, gas-liquid etc) and the quality is monitored via the distribution of mean and turbulent kinetic energy T.Kumaresan, Jyestharaj B Joshi (2005).

The quality of flow generated by the impeller depends upon the impeller design in STRs. The power number plays an important role on the flow generated by the impellers of STRs. At a relatively low power number (0.1-0.5), impellers will generate a mean flow, while at higher power numbers (>3) a flow of higher turbulent kinetic energy is generated. The flow proceeds from the impeller and circulates within the vessel; the average kinetic energy is converted to turbulent kinetic energy.

STRs are very important pieces of equipment as the site of important factors such as agitation in chemical and biochemical reaction processes, like fermentation (and other oxygenation and hydrogenation processes) where a large gas handling capacity and an effective gas dispersion is required in order to generate a large interfacial area J M.T Vasconcelos, S. C.P Orvalho...(2000).

2. Effect of Impeller design on mixing;

multiple impellers verses mixed impellers

The Rushton turbine has been the standard impeller choice for gas dispersion applications in Stirred Tank Reactors for many years. More recently concave blade turbines feature blades which are symmetric with respect to the plane of the disc; this with the entry of gas through the bottom of the vessel creates a flow pattern which is distinctly asymmetric. The flow pattern and power number in a STR depend on many factors which relate to the impeller design, such as; the impeller bond angle, the number of blades, the blade width, blade twist, blade thickness, the pumping direction and the interaction of flow with the vessel wall. T.Kumaresan, Jyestharaj B Joshi (2005) carried out comparisons of the flow pattern in STRs, analysing factors such as; average velocity, turbulent kinetic energy, maximum energy dissipation rate, average shear rate and normal turbulent stress. These characteristics are presented on the basis of equal power consumption in order to characterize the flow generated by different geometrics.

As pre-mentioned, the standard Rushton turbine (which consisted of six-bladed disc turbine) was once one of the most popular impellers found in industry. However, weaknesses of this type of turbine have been identified as an ideal gas disperser. Firstly, once gas in introduced to the vessel, there is an important fall in power demand, which reaches heights of ~50% of the gassed value. This results in a loss of the potential for heat and mass transfer. It also represents an economical loss with respect to the power prediction. Secondly, flooding, which occurs at respectively low gas flow numbers, handicaps the capacity of the Rushton turbine. These disadvantages are as a result of the formation of high-speed and low pressure trailing vortices which occur towards the rear of the impeller blades. These characteristics are associated with the boundary layer separations. Cavities at the rear of the impeller blades lead to an increase in the pressure, which leads to a fall in power with respect to the un-gassed situation. Cavities also control the turbine hydrodynamics and dispersion characteristics J M.T Vasconcelos, S. C.P Orvalho...(2000).The number of blades and the curvature of the blades influence the size and rotation of the trailing vortices. The blade shape has a major influence on the formation of the ventilated cavities; therefore the blade shape has the capacity to reduce the size of the cavities so that the gassed power may remain high. J M.T Vasconcelos, S. C.P Orvalho...(2000) describes alternative turbine designs other than the flat blade Rushton turbine, such as the arrowhead disperser, curved blade turbine and more recently, the divided, inclined blades turbine. The characteristics of concave and convex designs were one of the first alternatives to be proposed. The concave-blade turbine presented flattened power curves and improved gas handling capacity, which lead authors to have reason to propose this type of turbine as an alternative to the Rushton turbine. These new radial flow impellers are of preferred selection to be used as gas dispersers for the bottom of vessels, with axial flow impellers in the upper positions of vessels J M.T Vasconcelos, S. C.P Orvalho...(2000).

Factors which must be monitored in experimental methods to establish the most efficient turbine in STRs are; Power draw, Flooding regime, Inter-stage Exchange rate, Mixing Time, Gas Hold-up and Volumetric Oxygen –Transfer coefficient kLa.

This review will now discuss the performance of a new gas dispersion impeller, with vertically asymmetric blades on work done by K.J Myers., A.J Thomas, A Bakker...(1999). The performance of recent blade designs are found have a deeper concavity. With these deeper blades the gas is dispersed from within the blade rather than from the large cavities behind the blade. The performance of a new impeller, which is designed to accommodate the alternating flow conditions above and below the impeller disc, with the main results of a study conducted by K.J Myers., A.J Thomas, A Bakker...(1999), will be discussed here in a comparative of more than twenty different geometries. Results of the BT-6, PD-6, D-6 and CD-6 impellers are to be the main results discussed. The D-6 and the CD-6 impellers are of standard impeller design. The PD-6 impeller is not a standard design, nor is it commercially available. However, the BT-6 impeller is a more recently developed impeller design, with vertically asymmetric blades. Figure 1 shows the shape of this impeller. The results shown in Figure 1, show a nominal blade dept of 0.22 D and blade length of 0.3 D. In vessels of 0.44 m, 0.60 m and 1.52 m, the performance of the impeller was studied and based on gassed and un-gassed power draw calculated measurements. Flow visualisation experiments, visual gas hold-up measurements and dynamic mass transfer calculations were carried out. The effects of oxygen probe response time and gas residence time are example of methods which are used to calculate the dynamic mass transfer measurements

Figure 1: The Chemineer BT-6 impeller, with vertically asymmetric blades. When viewed from above, this impeller is designed to rotate clockwise.

Figure 2 below, shows a schematic of the Chemineer BT-6 and the blades have a concave shape. These curves consist of different radii and different length. The top leaf of the blade is longer than the bottom part. The back of each blade is rounded, as opposed to that of the Rushton turbine design. This prevents the formation of cavities at the rear of the blades. In operation, the gas is captured under the longer overleaf at the top part of the blade, and is then dispersed from a deep vortex at the inside of each blade. This design reduces the formation of large gas filled cavities at the rear of the blades of the impeller.

Figure 2: Shows the impeller power number versus the Reynolds number curves for the BT-6, CD-6 (semi-circular blades) and D-6 (flat blades) impellers

In the plot above, the un-gassed power number of the BT-6 is plot as a function of Reynolds number. The BT-6 is compared to the semi-circular blades of the CD-6, and the flat bladed turbine of the D-6, which is reprehensive of the Rushton type turbine. From the graph, it is clear that the D-6 has a fully turbulent power number of 5.5, CD-6 has power number of 2.8 and the BT-6 has power number of 2.3 K.J Myers., A.J Thomas, A Bakker...(1999). These results depend on scale and exact geometry.

Figure 3: Shows the relative gassed power draw as a function of the gas flow number and the Froude number for the BT-6 impeller. Power draw curves at Fr = 0.9 are shown for the D-6 and CD-6 impellers

The gassed power draw of the BT-6 impeller is illustrated in Figure 3. The power draw curve is compared with that of CD-6 and D-6. From the graph, it is obvious that the relative gassed power draw of the asymmetric blade BT-6 impeller is much higher than that of the other impellers. Under operating conditions, the majority of the industrial agitators are designed to operate at approximately 80% of the rated motor power. The gassed power of the experimental studies conducted by K.J Myers., A.J Thomas, A Bakker...(1999) show

advantageous characteristics, since it does not drop below 80-90% of the un-gassed power draw, depending on the Froude number. No expensive dual speed or variable speed equipment is required since the motor will not overload when the gas flow is interrupted. .

In Figure 4, PD-6 have deep concave blades which are symmetric to the plane of the disc. This is plot and compared with the characteristics of the D-6, CD-6 impeller. The figure shows the average of many experiments. For a given geometry and impeller speed, the amount of air at the complete dispersion point is determined. The diameters are varied so that the power draw and torque are kept constant, with the impellers operating at the sane speeds. The BT-6 impeller is far more efficient than the other impeller types. This is illustrated since the BT-6 impeller can disperse approximately five times more gas than the D-6 impeller, approximately twice as much gas is dispersed, in comparison with the CD-6 and also approximately two thirds more gas is dispersed for the BT-6 impeller, than that produced by the PD-6 impeller, with deep symmetric blades K.J Myers., A.J Thomas, A Bakker...(1999).

Figure 4: Shows the Gas Dispersion capability which is relative to the D-6 impeller for the BT-6, PD-6 (deep concave, symmetric blades) and CD-6 impellers.

From this study conducted by K.J Myers., A.J Thomas, A Bakker...(1999), it may be concluded that the application of deep blades, which are vertically asymmetric, the gas dispersion performance of disc style impellers shows significant improvement of gas dispersion in STRs. The rising gas flow is captured by the overhanging blade on the top of the blade; the gas flow is then dispersed from the inside of the blade. This prevents the formation of gas filled cavities at the rear of the impeller blade. This impeller produces relatively flat power curves; it is suitable for applications of varying gas flow rate and varying liquid viscosity during the process.

3. Energy Dissipation rates in Stirred Systems

influence on liquid-liquid and gas-liquid contacting

M. Fujasova, V. Linek..(2004)

The energy consumption in STRs is an important parameter in the design of equipment for different impeller types. The impeller geometry has a significant influence on the mass transfer efficiency (kLa) of STRs. Steady state methods of volumetric mass transfer are suitable only for the measurements of kLa in lab scale conditions. Whereas dynamic volumetric mass transfer methods are suitable for application for industrial applications. However, there is a problem with regard to the driving force determination M. Fujasova, V. Linek..(2004). This paper highlights the work done by M. Fujasova, V. Linek..(2004), and provides the transport characteristics for two types of impellers. The first example is an impeller of 45o pitched blade impeller pumping upwards through the vessel and also the Narcissus impeller. The results of M. Fujasova, V. Linek..(2004)’s work was then compared to the transport characteristics of configurations of various impeller types, such as the Rushton Turbine (RT), the Techmix335 impeller pumping up (TXU) and down (TXD), the Pitched Blade Impeller pumping Down (PBD) and their combinations.

Significant attention is paid to the kLa measurement in this paper. Volumetric mass transfer, gas hold-up and the mixing intensity of dispersion are presented with respect to the power consumption and the superficial gas velocity for impeller configurations of single, double and triple impeller configurations.

The vessel of selection for experimental work which was selected by M. Fujasova, V. Linek..(2004), had an inner diameter of T = 0.29 m and equipped with one, two and three impellers on a single shaft which is set through the centre of the vessel. Four selected baffles were aligned symmetrically on the vessel wall, each of diameters. The liquid height was proportional to the vessel diameter, so increased at the same rate as the T value changed. This was dependant on the number of impellers used to have agitation for gas-liquid dispersion. The diameter of the impeller was one third of the diameter of the tank.

Tracer experiments for triple impeller configurations were used to measure the mixing intensity of the liquid phase. The responses of the upper stages were recorded with the application of conductivity probes. The volumetric mass transfer coefficient was obtained with the use of the dynamic pressure method (DPM). This method depends on the change of total pressure within the vessel M. Fujasova, V. Linek..(2004).

The thickness of the metal used to make the blades of the impeller; the diameter and height of the hub from which the impeller is fixed to the shaft are examples of subtle geometric changes which are likely to affect the power number of the impellers used. A Chemineer D-6 is the selected impeller type in this experiment, which was conducted by J.M Smith and Z.Gao (2001). Table 1 illustrates the dimensions of the impellers which were used.

Table 1: Shows the Un-gassed fully turbulent power numbers of the impeller types used.

Sketches obtained from the work carried out by J.M Smith and Z.Gao (2001) is shown in Figure 5.

This figure illustrates the design of the blade profiles of the various hollow blades. The Chemineer CD-6 impeller consists of semi round blades.

Figure 5: Shows the turbines and their representative hollow blade profiles.

The horizontal area projected by these blades is like that of the Rushton D-6 impeller of the same diameter. The longer upper overleaf of the blade surface of elliptical cross section, ensures effective gas capture and very good gas dispersion. The Ruston turbine has significantly higher Power numbers than the hollow blade turbines which have un-aerated power numbers. This highlights the reduced intensity of the vortices which may develop behind the blades of the impeller. The convective rotation of these vortices which are trailing, carry away the majority of the input shaft energy to the bulk of the fluid. At this point, the rotation will decay into random turbulence.

At ambient and elevated temperatures, the power draw of the impellers was obtained for aerated operation. High volumetric gas flow rates were available due to carrying out this work at high temperatures. It is suitable to establish the initial Power number for each impeller in an un-aerated liquid. The power draw was measured, at a given impeller speed, by flooding the impeller while it was rotating slowly or stationary. This was carried out with a high gas flow rate which was available and allowing the flow field to stabilize J.M Smith and Z.Gao (2001). The impeller speed was adjusted and brought to the fixed value which was required for a given set of measurements. Reducing the gas supply resulted in the lowering of

the gas flow number. This procedure is applicable in the majority of situations, unless otherwise stated.

Figure 6: Shows the Relative Power Demand (RPD) of a Rushton Turbine with alternating gas rates.

The open points in the figure were measured rising gas rates J.M Smith and Z.Gao (2001). The open points fall differentially from the closed points which were obtained with reducing gas rates. The turbine shows a slight hysteresis while regaining the RPD, once flooded, as the gas flow rate is lowered. A small percentage of gas flow, cause a delay in the recovery of the RPD to a loaded condition at 3rps.

Now the results from the CD-6 Hollow Blade Impeller will be investigated. More modern impeller types have a much greater ability to disperse gas in STRs in comparison with the conventional Rushton turbines. Modern impellers also exhibit much greater hysteresis in the change of the RPD once the flooding commences. The transition to buoyancy dominated the flooded conditions takes place at a gas flow number which is approximately three times the expected flooding point of a Rushton turbine J.M Smith and Z.Gao (2001). The loading-flooding transitions for a Chemineer CD-6 impeller with rising and falling gas rates in illustrated in Figure 7. Figure 7 was obtained in cold conditions. The maximum gas handling capacity was not included in these results. This is because the authors were unable to report it, since it exceeded the maximum gas flow number that was possible to be achieved using the available equipment.

Figure 7: Illustrates the loading-flooding transitions for a Chemineer CD-6 impeller with rising and falling gas rates.

Reducing the gas flow rate to the flooded CD-6 impeller causes the onset of significant hysteresis prior the loaded conditions are restored. From figure 7, the open points and lines were obtained with increasing gas rates and fall differentially from the closed points and lines which correspond to the reducing gas rates J.M Smith and Z.Gao (2001).A gas flow rate at impeller speed of 3 rps, during increasing gas rates operation, is twice that of the gas rate when the gas rate is falling.

Figure 8 below, illustrates data which is available while the CD-6 is operating in cool gassed and hot sparged operating conditions J.M Smith and Z.Gao (2001).

Figure 8: Illustrates the RPD data of a sparged CD-6 impeller in hot and cool conditions.

Here flooding operation was not maintained at a rotational speed of 5 rps. The Power demand in hot operational conditions is approximately 10-15% higher than similar measurements in cool conditions. The cavities behind the blades of the flooded impeller are thrown upwards as a result of the buoyancy J.M Smith and Z.Gao (2001). Therefore a larger surface of the impeller blades is now in contact with the liquid than there was prior to flooding.

4. Effect of gas on Power draw of selected impellers and on the fluid mechanics of two-phase dispersion

more effective gas dispersion research

Gas Hold-Up in Stirred tank Reactors

Gas hold up in Stirred Tank Reactors (εg), is a widely studied parameter in the Literature of Stirred Tank Reactors (STRs). Gas hold up affects the gas-liquid mass transfer in STRs. Yawalkar, Pangarkar and Beenackers (2001) give an overview of the experimental work on the gas hold up in STRs, which has been carried out by various experimenters. These results are divided into two main categories: Correlations with an approach based on dimension (such as Froude Number, D/T ratio, Flow number etc) and correlations which are based on the Kolmogoroff’s theory approach, which has the power dissipation as the basis. The majority of the early work was carried out based on small diameter reactors. However the results obtained from the smaller tank reactors had a notable divergence from the actual gas hold up in large tank reactors Smith (1991), this also attributed to a different degree of dispersion of gas in a STR to the alternative flow regimes.

Yawalkar, Pangarkar and Beenackers (2001) illustrates comparison made in the study of gas hold up, in large STRs (T=1 to 2.7 m) and equipped with six blade Disc Turbines (DT), in comparison with large STRs (T= 0.57 to 1.5m) equipped with six blade pitched down flow turbines (PTD). A wide divergence in predictions based on different correlations for the configuration (tank size, type of impeller, D/T, C/T etc .) of the system in question and operating parameters ( impeller speed and gas flow rate). The disagreement is attributed to the exclusion of dependence of gas hold up on the flow regime for specific operating conditions and geometric configuration. The gas hold up data was correlated based on the clinging cavity regime and the large cavity regime. Smith (1991) proposed that it is possible to predict gas hold up in STRs with the use of dimensionless groups. With an exception for

small smaller tanks (T 0.44 m), Smith (1991) observed that gas hold up can be predicted

adequately with the use of the following term: (Re.Fl.Fr) 0.35(D/T) 1.25. Comparing this work to the work carried out by Greaves and Barigou (1990), Smith (1991) observed that a higher result of gas hold up (εg>0.08) contributed to a substantial increase in the difference between the two data sets. Table 2 shows the comparison of the gas hold up predictions for both of the mentioned correlations, available in literature for large tank STRs which are equipped with six blade disc turbines with alternating geometric configurations and operating parameters of T=1 m. From the table, it can be seen that the correlation of Greaves and Barigoue (1990) predicts large values of εg in comparison with the correlation of Smith (1991) as the gas flow

rate is increased. From Table 2, it is noted that at a gas flow rate Qg 1.20*10-2m3/s the gas

hold up predicted by Greaves and Barigou (1990) is almost twice the predicted correlation of Smith (1991).

Table 2 also illustrates a comparison of work carried out by Rewatkar et al (1993). Rewatkar et al (1993) carried out work on large diameter STRs which were equipped with six bladed

pitched down flow turbines. They studied the effect of alternative types of sparger (pipe, ring, concentric ring and conical spargers) and its location in association with the impeller on the systematically gas hold up in addition to D/T and gas flow rate.

Table 2 gives a summary of the results of studies and experimental details of work carried out by Greaves and Barigou (1990), Smith (1991) and Rewatkar et al (1993). With the works carried out by the pre-mentioned, on large diameter STRs (0.57 m to 2.7 m) and using the Kolmogorff’s theory approach with power dissipation as the basis, data of gas hold up generated with tank diameter of 0.57 m and the STR equipped with either DT and PTD, can be regressed in the form εg = f(εm,Vg), with εm representing the power input per unit mass of the aerated liquid.

Table 3: illustrates the power input data obtained

This relationship is illustrated in Table 3. Yawalkar, Pangarkar and Beenackers (2001) illustrate that the power input data obtained in their work for DTs with T=0.57 m were approximately 15% in agreement with the correlation proposed by Hughmark (1980), whose correlation presented was for six bladed disc turbines and were applicable over a wide range of system configurations and operating parameters.

This data is represented in Table 4.

Table 4: illustrates the correlation for six bladed disc turbines.

The above equation is considered reliable and is based on 391 data points with a standard deviation of 0.117 Yawalkar, Pangarkar and Beenackers (2001). This equation may also be used to derive Power input for a wide range of system configuration and operating parameters for its corresponding applications. The following correlation was obtained with respect to Disc Turbines Yawalkar, Pangarkar and Beenackers (2001);

εg = 0.515(εm)0.25(Vg)0.40

Figure 9: illustrates the Parity plot of its stated equation

This equation is used on 122 data points. Figure 9 shows a Parity Plot of this Equation: This figure illustrates the significant scatter that exists around the correlating line with a maximum

deviation of 50%.

Considering gas hold up data of large STRs (T=0.62 m and 2.67 m) which is equipped with Disc Turbine of D= T/3, the following equation;

εg = 1.28(εm)0.26(Vg)0.66

is used to satisfactorily correlate the gas hold up data Yawalkar, Pangarkar and Beenackers (2001).

The correlations of both Equations 2 and Equation 3, for Disc Turbines predict εg to correspond to each other for given Power Input and Superficial Gas velocity. Yawalkar, Pangarkar and Beenackers (2001) show that gas hold up on T=2.67 m are based on a narrow range of impeller speeds of N=1.42-1.67 rev/s and a gas flow rate in the range of 66.4 to 185.8×10-3 m3/s. The minimum impeller speed for complete gas dispersion is denoted by Ncd, has values between 1.66 and 2.77 rev/s. Therefore high scatter of the data points as seen in Figure 9 may be as a result of the wide range of system configurations (T=0.57 m to 2.7 m and D/T=0.25 to 0.5) and the range of operating conditions which were proposed by

researchers (0.5 N/ Ncd 1.60) Yawalkar, Pangarkar and Beenackers (2001).

Impeller speed also has a large significance on gas hold up in STRs. Ncs denotes the minimum impeller speed necessary for complete suspension of solid particles in solid-liquid systems in STRs. Ncs may be found experimentally via simple methods such as visual

observation, power measurement, mixing time study and through radioactive tracer technique Yawalkar, Pangarkar and Beenackers (2001) or it may also be determined from correlations which are available in literature Zwietering, 195; Chapman et al., 1983a; Rewatkar et al., 1991. In gas-liquid systems, Ncd denotes the minimum impeller speed for complete dispersion of gas in the liquid phase of a STR. Therefore Ncd is representative of the minimum impeller speed required to have all of the liquid of the tank in contact with the sparged gas. Ncd is a very important parameter to consider so that complete dispersion of the gas in the liquid phase may be obtained at Ncd with minimum power input, when reviewing the economical gas-liquid mixing operation of STRs. At impeller speeds below Ncd the mixing process is not sufficient and will result in incomplete dispersion of gas Yawalkar, Pangarkar and Beenackers (2001). Therefore the overall performance of the STR will not be up to standard operating conditions as a result of very little or no gas dispersion in the region below the impeller.

Figure 10: this figure illustrates the minimum impeller speed for complete dispersion of the gas Ncd in STR

A) Illustrates the dispersion of gas in STR with increasing Impeller Speed

B) Illustrates the gassed to un-gassed power number ratio (NPg/NPo) against gassed flow number (Flg).

From Figure 10, Yawalkar, Pangarkar and Beenackers (2001), it is shown that the degree of dispersion of gas in the STR increases as the impeller speed is increased. The bulk mixing stages are highlighted at differentiating gas flow rates. (a) illustrates little or perhaps no gas dispersion; (b) illustrates sufficient dispersion for the upper part of the vessel to act as a bubble column; (c) illustrates gas circulation in the upper region of the vessel and an indication of the start of movement in the lower region; (d) illustrates circulation throughout the vessel after a small increase in the speed of the impeller; (e) illustrates recirculation of the gas with and secondary loop formation.

The minimum requirement of impeller speed for bulk gas mixing is in the transition of stages (c) and (d). This may be classified as the onset of flooding with an impeller speed of Ncd at this point. The gassed to un-gassed power number ratio (NPg/NPo) against gassed flow number (Flg) plot is shown as Figure 2B. At Ncd, the growing gas cavity covers the rear face of the impeller blade. This results in a decrease of the power input to the impeller.

A comparison was carried out on a large range of system configurations and operating parameters by Yawalkar, Pangarkar and Beenackers (2001). Correlations for Ncd were studied for STRs equipped with a standard six blade disc turbine, four blade mixed blade mixed flow impellers pumping up (MFU) and pumping down (MFD), six blade MFU and MFD, Axial flow impellers pumping up (AFU) and pumping down (AFD), Disc Turbines (DT) and Angular-Bladed Disc Turbines (ADT).

Turbulent fluctuations also have a significant effect on relative gas dispersion in STRs. Turbulent fluctuations in the gas-liquid dispersion control the drag on the bubble within the vessel; hence have an effect on the bubble size. When the surface tension force is outweighed by the hydrodynamic stresses, the bubbles break up. From Yawalkar, Pangarkar and Beenackers (2001), it was found that bubble break up is as a result of eddies of the inertial sub range. As mentioned, the turbulence increases the drag on the bubbles, which in turn reduces the bubble rise velocity. With this reduces rise velocity, the bubbles are more susceptible to be entrained in the downward liquid flow which is generated by the impeller. Therefore, an increase of turbulence in the vessel will result with an increase in the gas hold up (εg) in the STR.

Complete dispersion of the sparged gas is achieved at an impeller speed of Ncd. Therefore it may be concluded that, at an impeller speed of Ncd, the bubble size generated close to the impeller has buoyancy less than the downward drag, which is caused by the downward liquid flow and the bubbles are then pulled down to the lower part of the vessel Yawalkar, Pangarkar and Beenackers (2001). Therefore at Ncd, the turbulence intensity is sufficed to allow complete dispersion of the gas throughout the liquid in the vessel to be achieved. Now it is apparent that the drag on the gas bubble, therefore the bubble size also and the gas hold-up depend on turbulence intensity. From this it was found that εg/ εgcd= f(N/Ncd). N/Ncd denotes relative dispersion of the gas and also gas hold up (from just mentioned; εg/ εgcd= f(N/Ncd) ) in STR for a given system configuration and gas flow rate. Therefore gas hold up may be defined in terms of the relative dispersion parameter, N/Ncd which in turn, represents the amount of gas retained in the liquid at the impeller speed N with respect to that at Ncd.

From this, it may be concluded that the gas hold up may be predicted over a range of system configurations and operating parameters, by a correlation based on the dispersion parameter N/Ncd.

6. Heat and Mass Transfer

The removal of heat from large scale STRs may be a critical design requirement of STRs. Large D/T ratios at an equivalent (εT)g are sometimes necessary to produce higher heat transfer rates A W Nienow (1998).

Enhanced gas-liquid mass transfer is an important topic which is required when considering a suitable impeller design for STRs. It is difficult to categorise the agitators, and determining whether one specific agitator is better than an alternative agitator. From work carried out by A W Nienow (1998), it was found that six impeller types produced equal kLa values at different scales in the range of 0.26 to 2.0 m diameter, with (εT)g and superficial gas velocity being kept constant. Even though the absolute kLa value is dependent on the trace chemicals which appear in the vessel, an example which is from the work of A W Nienow (1998), found that an electrolyte in water and in water plus antifoam the kLa varied considerably from case to case. The use of Rushton turbines and axial flow hydrofoils give similar results. Lower shear of bulk blending may cause a change to the value of kLa. The change of agitators, as may be required in cases where they must be frequently replaced (e.g fermentation processes) may also produce a varying kLa result to allow the system to adjust to the new conditions which are produced to the vessel to stabilize. The kLa value may be increased via increasing the power into the system (e.g increasing the power input to the broth or by using higher aeration rates), while ensuring the impeller is not then susceptible to flooding conditions A W Nienow (1998).

Performing unsteady –state oxygen adsorption and desorption experiments is another technique which may be performed to assess the ability of various mixing configurations to promote interphase mass transfer in gas-liquid dispersions H.J Palmer (1999). When such experiments have been performed, it is possible to compute the overall mass transfer coefficient, kLa. In an oxygen uptake (oxygen absorption) experiment, nitrogen is bubbled into the tank via the sparge ring at speeds in the range of 550-650 rpm, in order to remove the dissolved oxygen, which allows the gas flow rate and the impeller speed to be adjusted in order to find their desired operating values H.J Palmer (1999). Since the oxygen reading now presents a value approaching zero, it was assumed that the system had reached a state of equilibrium. The impeller speed may be recordered with the use of a data-acquisition program on the micro computer. After a designated time of 60 s the source of gas is changed from nitrogen to oxygen H.J Palmer (1999). The gas flow rate is adjusted carefully to its desired operational value. The dissolved oxygen concentration data at specific intervals and then used to determine the overall volumetric mass-transfer coefficient for the interphase transfer KLa. The KLa is defined by the following equation H.J Palmer (1999);

Where C denotes the dissolved oxygen concentration in the liquid phase at any instant and C*

denotes the dissolved oxygen concentration which is in equilibrium with the feed of gas to the vessel. (Integration of this equation provides an equation which describes the DOC as a function of time.)

The rate of mass transfer is also temperature sensitive. For optimum experimental results, it is desirable to keep the temperature of the vessel constant at all times. However, this is not always possible due to imprecise temperature control from the day to day operation of the STR. This is somewhat impractical, therefore slight variations of temperature must be allowed for. H.J Palmer (1999) provides a correlation which is used to overcome this impracticality;

Where T denotes the temperature of the liquid in the vessel, measured in degrees Celsius.

A sample of mass transfer results from experimental work carried out by H.J Palmer (1999) is illustrated in Figure 11.

Figure 11: illustrates the effect of impeller type on the KLa which is obtained at a gas feed rate of 30 L/min in a conventional tank configuration.

From the plot, it is shown that the KLa increases as the power input is increased. It is also noted that the results of the A310 and A410 impellers are very similar. The KLa values for the R100, which is a radial flow impeller, was at its peak at the highest gas feed rate and was almost non-existent at the lowest feed rate (0.057 cm/s) H.J Palmer (1999). It is also noted that the value of KLa does not increase as sharply with the increase with P/V ratio for the R100 impeller type as it appears to do for the axial flow impellers.

It may be concluded that greater mixing efficiency is obtained with the ease at which the mixing system draws air down towards the impeller region, providing high rates of gas-liquid mass transfer, without causing additional gas sparging H.J Palmer (1999).This is a feature which may prove to be very useful when applied to industrial applications which require contacting of liquids with gases which may be toxic, reactive, or corrosive.

7. Scale-up

Scale up is a very important feature of designing suitable impellers for STRs. The impeller speed N will normally decrease as the size or scale of the vessel increases.

For single turbines, a higher specific energy dissipation rate or power per unit volume is necessary to prevent flooding of STRs when attempting scale up. On scale up, a single impeller will be more likely to experience flooding than the smaller unit with the application of the same power number for both models. Therefore, for scale up, complete air dispersion and gross recirculation will require a much higher specific power input A W Nienow (1998). The gross recirculation is a condition which may be difficult to be produced.

For multiple turbines, the lower impeller of a stack of Rushton type turbines is expected to behave as a single impeller, provided adequate impeller spacing is present A W Nienow (1998). Even though the upper impellers behave similarly to each other, the bottom impellers behave somewhat differently. The bottom impeller acts as a gas distributer, so that the upper impellers do not carry the same extent of air as the lowermost impeller. This is also true for the arrangement of the stack of impellers during conditions of flooding.

Figure 35 from A W Nienow (1998)’s work describing possible improvements on Rushton turbines, illustrates the Pg/P ratio for the upper and lower impeller of a dual Rushton turbine system. The data contained in this graph highlights the lower degree of flooding and higher power draw requirement of the upper and lower impellers A W Nienow (1998). It is also evident from the graph that upper and lower impellers have an equal effect on the power draw, when a speed is sufficient to cause gross re-circulation within the vessel. This condition is not obtained as frequently, however, for larger scale vessels. A W Nienow (1998) give an illustration of an established relationship for the power drawn for the multiple impeller system which is shown below;

=

Where may be estimated from equations used to estimate the Power draw for impeller

design.

The dept and back pressure also have an influence on scale up of STRs and impeller design. Increasing the back pressure also increases the power draw in proportion to the reduction in volumetric flow rate A W Nienow (1998). Therefore, impellers which are close to the base (apart from the lowermost impeller) will not draw a considerable amount more power than

impellers which are close to the top of the vessel. This is because the volumetric flow rate is decreased with increasing dept A W Nienow (1998).

The high viscosity broths are another important factor to be considered in the scale up of STRs and impeller design. Good blending and closer impeller are essential for in high viscosity broths in agitators A W Nienow (1998). Single agitator power characteristics are collectible to give a result of combined aerated power characteristics A W Nienow (1998).

Conclusion

This basis of this report is to highlight the development of impeller design for mixing in agitated reactors in recent years. An example of industrial applications of mixing in agitated reactors is with the use of Stirred Tank Reactors. Stirred Tank Reactors and stirred vessels are widely used in the Chemical, Pharmaceutical, food industries and also in waste water treatment plants.

It may be concluded that the gas dispersion performance, in such applications, may be significantly improved with the use of vertically asymmetric blades. The longer upper overleaf of the blade surface, ensures effective gas capture and very good gas dispersion. This highlights the reduced intensity of the vortices which may develop behind the blades of the impeller. The convective rotation of these vortices which are trailing, carry away the majority of the input shaft energy to the bulk of the fluid. At this point, the rotation will decay into random turbulence. The Ruston turbine has significantly higher Power numbers than the hollow blade turbines which have un-aerated power numbers.

Taking the energy dissipation rates into account, it was noted that the energy consumption in STRs was an important parameter in the design of equipment for different impeller types. The impeller geometry displays a significant influence on the mass transfer efficiency (kLa) of STRs. However, these steady state methods of volumetric mass transfer are suitable only for the measurements of kLa for lab scale applications. Whereas dynamic volumetric mass transfer methods are much more suitable for application, for industrial applications.

The scale up of impeller design from lab scale to industrial scale is another very important feature of designing suitable impellers for STRs. The impeller speed normally causes a notable decrease as the scale of the vessel increases.

Mass Transfer is also an important factor to consider with respect to mixing in agitated reactors. It is concluded that a higher mixing efficiency is obtained with ease, as the mixing system draws air downwards in the direction of the impeller region, providing high rates of gas-liquid mass transfer, without causing additional gas sparging. This is a feature which may prove to be very useful when applied to industrial applications which require contacting of liquids with gases, which may be toxic, reactive, or corrosive.

Considering the effect of gas on power draw of selected impellers and on the fluid mechanics of two-phase dispersion, it was concluded that the gas hold up can be predicted over a range

of system configurations and operating parameters, by a correlation based on the dispersion parameter N/Ncd.

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