Clarifier -1

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DESIGN AND OPTIMISATION OF FINALCLARIFIER PERFORMANCE WITH CFD MODELLING

D. J. Burt*, BEng, MSc, CEng, MIMechEJ. Ganeshalingam+, BSc, MSc, PhD, AMIChemE

Presented at the CIWEM / Aqua Enviro joint conferenceDesign and Operation of Activated Sludge Plants

19th April 2005.

ABSTRACT

A Computational Fluid Dynamics (CFD) prediction procedure for computing the internalhydrodynamic behaviour of final clarifiers is presented. Calculations are carried out andpresented for comparison with experimental measurements of point velocities in a shallowcircular clarifier in operation at a UK waste water treatment works. The calculations comparefavourably with the data and the model has subsequently been used for many design studies. Aseparate study is described where CFD predictions are used to investigate the influence ofintroducing energy dissipating influents (EDI), varying the stilling well diameter, addingStamford baffles at the side wall or placing an influent floor baffle (McKinney baffle) below thestilling well. From the study it is possible to determine the sizes and combinations of internalbaffling likely to give the best clarification performance across a range of operating conditions.The success of the optimum design is judged by the depth of the settling sludge bed and themagnitude of the effluent suspended solids (ESS) for the state points considered.

Key words: Computational Fluid Dynamics (CFD), Energy Dissipation Influent (EDI), StirredSludge Volume Index (SSVI), McKinney-baffle, Stamford-baffle.

*Senior Engineer, MMI Engineering, Bristol, UK. and Dept of Mechanical Engineering, QueensBuilding, University of Bristol.+

Project Engineer, MMI Engineering, Bristol, UK.

INTRODUCTION

According to the CIWEM handbook(1), Activated sludge plants (ASPs) are responsible for thetreatment of about 50% of all sewage treated by biological oxidation in the UK. These plantsare able to produce effluents in compliance with the current legislative requirements forsuspended solids (SS), biochemical oxygen demand (BOD) and nitrate content (ammoniacalN). However, European legislation is demanding improvements in effluent quality and therequirement to treat larger volumes in our expanding towns and cities, is putting considerablepressure on the existing sewage treatment infrastructure. Consequently, there is much interestin the enhancement and improvement of the performance of the final stages of the activated

sludge process. For, it is in the final clarifier that the sludge settles and it is the separationefficiency of the clarifier that largely determines the effluent quality of the ASP.

It is well known that the settling efficiency of the clarifier is greatly affected by the hydrodynamicflow paths within the tank. In 1940, Anderson (2) published a paper showing how the internalflows within the tank led to a billowing of suspended solids immediately below the effluent weir.

Figure 1 is taken from the work of Anderson and shows a section through a typical circularclarifier of 40mdiameter and 3.5mside wall depth. The influent has an axial riser deliveringfeed into the tank through a central radial diffuser. Flow rates for this tank are in excess of 1000m3/hrgiving inlet velocities of order 0.1 m/s. A conventional stilling well is used to attenuate theturbulent inlet flow; the diameter and depth of this well are significant design parameters. Themeasurements show the typical density driven current observed in all circular clarifiers. The


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dense feed with a concentration of mixed liquor suspended solids (MLSS) typically ranging from2000 to 6000 mg/l, falls from the influent and generates a radial underflow. In order to balancethis momentum, the underflow is matched by a return flow at higher depths in the tank. Theresulting re-circulation is one of the reasons why the sediment blanket appears to lift at the sidewall below the effluent weir. This early work clearly shows that the flow in a circular clarifier isfar from one dimensional but can be considered as close to two dimensional with axissymmetry.

The standard technique for designing a secondary clarifier is mass flux theory, sometimesreferred to as state point analysis. This method uses a one-dimensional settling model thatcannot take account of the complex internal flow patterns flow present in the clarifier. Short-circuiting of the influent flow, scouring of solids and re-entrainment of solids into the re-circulating flow pattern all contribute to the effluent suspended solids (ESS). Excess influentmomentum, perhaps from an undersized stilling well, can invoke a complete failure withoverflow of the sludge blanket. Even when the clarification surface area is over sized, relative tomass flux theory, the tank may still fail, or perform badly in practice because of these internalflow features. Clearly there is a need to use a new design method for clarifiers that is capable ofovercoming the limitations of mass flux theory.

In this work, a CFD modelling technique was developed, verified and validated to determine the

internal hydrodynamic performance of secondary clarifiers. The model is based on anadaptation of the IAWQ drift flux model part of which is discussed in the Scientific andTechnical report No 6 by Ekama et al(3). Results from the model were used to demonstrateseveral characteristic internal flow features that are thought to be causes of poor clarificationperformance.

A validation study is presented which compares local velocity data with experiment for the RyeMeads clarifier, measured extensively by Richardson et al (4)(5). Also a single design study ispresented for a typical UK clarifier where a number of internal modifications have beeninvestigated in order to understand the influence of introducing energy dissipating influents(EDI), varying the stilling well diameter, adding Stamford baffles or placing an influent floorbaffle (McKinney) below the stilling well. In the design study, the flow and settling behaviours

were calculated for a base case configuration and then for alternative designs. The aim of thework was to determine the internal design likely to give best clarification performance across arange of operating state points. The success of a design was judged by the height of thesettled sludge bed and the magnitude of the ESS for all state points.

MASS FLUX THEORY

Assuming sludge settling velocity is a unique function of the solids concentration; flux, theproduct of settling velocity and solids concentration, can be plotted against solids concentration.This generalized flux curve can be calculated from any of the standard correlations for settledvolume index (SVI) or stirred sludge volume index (SSVI), e.g. Pitman (6) and White (7) orWahlberg and Keinath(8). On this same curve two operating lines, over flow and under flow, arealso plotted. If the intersection of these lines is above the flux curve then clarifier failure ispredicted. At any state point, the crossing of the underflow and overflow lines must be belowthe flux curve for safe operation. If the underflow line is below the flux curve, but becomestangential to it at higher concentrations, then the tank is said to be critically loaded and on thepoint of failure. There are other definitions of tank failure obtainable from mass flux theory thatare not discussed here, see Ekama et al(3).

The main difficulty in estimating the limiting situation with mass flux theory is a poor correlationbetween SSVI and flux, largely due to the flow patterns and physical processes previouslydiscussed. Therefore, a safe design is not usually based on the generalized flux curve but ononly 80 % to 90 % of the value at any point. Care must also be taken in use of correlations for


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the mass flux curve as there are significant differences in these. Figure 2 shows howcorrelations can differ greatly when used in a practical selection application.

CFD MODELLING

Most of the CFD modelling work for secondary clarifiers reported in the literature uses a form ofthe algebraic slip (ASM) or drift flux model of Wallis (9) to represent the two-phase mixture of

water and activated sludge. Zhou and McCorquodale(10)

describe how the density variationsand settling velocity relationships may be modelled. Lakehal and Krebs(11) have extended thismodel to include variations in fluid mixture rheology. Two recent PhD theses discuses themodel in greater detail and include extended rheology functions, see Armbruster (12) anddeClercq(13).

In this implementation, a modified version of the CFX code was used as the modelling tool withan adaptation of the IAWQ drift flux model(3). The simulations were performed in two-dimensional, axis-symmetric co-ordinates with models for sludge mixture density, viscosity,following Bokil and Bewtra(14), or Dahl(15), Dick and Ewing(16), and the double exponential

settling function, following Takcs(17). A low Reynolds number k - turbulence model(18) was

used in deference to the large variation in mixing time scales present in a clarifier. The model

uses a multiphase method where the sludge is able to move independently with respect to thewater but only in the direction of a slip vector; where, in this implementation the slip only acts inthe direction of gravity.

Boundary Conditions

Even with its limitations, there is still value in using mass flux theory to calibrate the boundaryconditions for subsequent CFD analysis. In the process of developing the generic CFD studiesreported here it was discovered that, without applying any factors of safety, tanks with sufficientsurface area to satisfy mass flux theory always fail before the mass flux limit when modelledwith CFD, in other word mass flux theory is the absolute upper bound of performance. Massflux theory also provides a useful indicator of the likely RAS solids concentration.

Physical Properties for Sludge

The Water Research Council (WRc) standard for clarifier design refers to the application of amethod based on 30 minute SSVI settling tests following the Pitman(6) and White (7) correlation.However, in validation studies for this model it was discovered that not all UK sites exhibit agood match to this correlation(19) and a more rigorous assessment of sludge settleability asdocumented by Ekama et al(3) should be used when obtaining the settling or Vesilind(20)

coefficients for use in the CFD model. Coefficients for the rheological constitutive relationshipsare difficult to obtain as there is no firm agreement on the type of viscosity model that should beapplied to activated sludge. Consequently there are no standard test procedures for

characterising sludge rheology. In these studies several rheological models were comparedand it is thought that models incorporating a yield stress term, see deClercq(13), are the mostappropriate.

VALIDATION CASE STUDY

The Rye Meads sewage treatment works (STW) has a population equivalent of 360,000 and a95 percentile effluent discharge consent of 15:08:03 (ESS:BOD:NH3) mg/l. The works has 3stages of activated sludge plant with a total of 16 final settlement tanks. The circular final tanksat Rye Meads are 28mdiameter with a side wall depth of 2.2m. They are shallow with a floor


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DESIGN STUDY

The design study is for a typical UK clarifier 55 (16.76 m) diameter and 8 (2.438 m) side walldepth. The tank has a 7.5 sloping floor with bridge scrapers and a pumped central hopper RASextraction. As built, the centre well is only 2m(12%D) diameter and 1.5m deep. Modificationsbased on a larger stilling well were considered either with a central EDI, similar to that shown in

Figure 4a or by mounting a baffle plate, Figure 5c, directly below the enlarged stilling well. Thissecond design, known as a McKinney floor baffle is described in Ekama(3), it is located at a

depth close to the side water depth with a gap sized for densimetric Froude number, Frd 0.7,at the highest flow rate. The key design parameter for the McKinney baffle is the slot gapbetween the bottom of the stilling well and the baffle. This is sized to keep 0.5 < Frd< 0.7 suchthat re-entrainment into the stilling well is prevented(12),

=

w

w

d

gh

UFr

where Uis the average velocity through the slot, his the height of the slot, is the mixturedensity at influent and w is the density of water, gis the gravitational constant.

Alternative designs were analysed at various state points representative of the bounds ofoperation. Comparisons are presented here for a forward flow of 156.3 m3 /hr, with mixedliquors at influent of 3700 mg/l, RAS ratio 0.51 or 1.0 and SSVI of 80 ml/g.

CFD Models Flow ratio ESS Depth below TWL Status

R (mg/l) (m)

As Built 0.51 312 0.245 Fail

a) 1 9 1.450 Pass

EDI 0.51 328 0.255 Fail

b) 1 12 1.925 PassMcKinney 0.51 24 0.825 Pass

c) 1 10 2.170 Pass

Table 1: Summary of bed depth and ESS for the model variations.

Figure 5 and Table 1 summarise the results for the study. Two state points are considered forthree geometries allowing the performance of each design to be compared. The first thing tonote from table 1 is that a critically loaded clarifier can be brought back into compliance bysimply increasing the RAS rate. Figure 5a shows how the smaller stilling well gives rise to ahigher sludge bed but, where the influent flow is now directed through the bed, the ESS is seento be the lowest of the three configurations. Figure 5c shows how the McKinney baffle acts to

separate the stilling zone from the settling zone and this design has the ability to carry thegreatest volume flow rate, the results in Table 1 show that, in this case, only the McKinneydesign is compliant at both of the RAS rates investigated. Comparisons for the EDI design withand without a Stamford baffle showed no difference in the effluent quality, see Figure 6.

DISCUSSION OF INTERNAL DESIGN PARAMETERS

Masss flux theory only gives guidance on the likely surface area required to effect clarification,and, as has been stated, this can be a significant under estimate. There are many designs ofcircular clarifier using different shapes and depths in use in the UK. Side walls are generally aminimum of 2m in depth but floor angles can vary from flat to very steep in excess of 45.


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Although there may be merit in deep designs (this is still a topic of investigation), the modelingtechniques applied here have largely been used to obtain maximum performance from flatbottom or shallow angle clarifiers with various forms of RAS removal system. From thesestudies a number of observations have emerged that have provided guidance on internalbaffling which can improve effluent quality and allow the tank performance to approach themass flux limit.

The Stilling Well

One of the first things to consider in internal design is the diameter and depth of the stilling well.The concept of a flocculating stilling well gained some credibility in the 1990s; however, ascan be seen in Figure 2a, a large stilling well is likely to generate significant re-entrainment fromthe settling region of the tank back into the stilling pond and this can persist even when an EDIis included. If the stilling pond is too small then the resultant down flux of momentum from theinfluent can be sufficient to disrupt the settled bed leading to higher beds and early failure. Thisimpact of a small stilling well on bed height is shown in Figure 5a. In various studies a stillingwell diameter of 20%D has been found to be most effective with depth set to half of water depthat the radius of the stilling well. Any shallow tank which has an influent based on a stilling wellonly design will suffer from strong density current effects at certain state points.

Improving the Influent

A central EDI, spreads the load out uniformly near the top of the stilling well reducing thedensity variation in the stilling pond. If it is designed well it can limit the influence of flow re-entrainment into the stilling pond and can give good performance across a range of statepoints. Studies on the influence of vanes suggest that these are largely inconsequential acceptwhen the combination of slot size and vane angle gives rise to excessive swirl and invoke re-suspension of the sludge below the stilling pond. The important requirement is to keep themomentum exiting the EDI ports at a level high enough to promote homogenisation but not sohigh as to produce re-suspension. Some workers have made much of the idea that the stillingpond is contributing to re-flocculation, however, it has been shown in the measurements of

deClercq(13) that there is little variation in the Particle Size Distribution (PSD) within the stillingpond. Other CFD studies of Camp number and this authors own concept of a G Scalar historyfunction(21) suggest that levelsof G high enough to promote orthokinetic flocculation of activatedsludge, as defined by Biggs(22), are only present in the firstfew tens of seconds following entryinto the stilling pond.

A McKinney baffle cuts the density current and, if designed correctly, completely separates thestilling and settling zones. However, the McKinney design shows sensitivity to bed depth as itoperates most effectively when the baffle sits at the same level as the sludge blanket; in thisway the flow exiting the influent forms a level radial jet across the top of the settling bed. Forlow flow or low SSVI situations a second density waterfall can form at the end of the McKinneybaffle which has a degrading effect on tank performance. Ideally, a McKinney baffle would trackthe bed height in operation or the bed height would be maintained (through RAS control) at thelevel of the McKinney. When including a McKinney baffle it is usually necessary to increase thedepth of the stilling pond to suit the required slot height.

Side Wall Baffles

In other design studies, modifications to the influent were augmented by a variety of side wallbaffling options, see Figure 4b. In shallow tanks with low floor angles, no clear benefit fromusing a Stamford baffle was observed and this is consistent with Figure 6. The CFD predictionsindicate that shear layer separation always occurs before the flow reaches the side wall in such


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a way that the effluent flow effectively by passes the location of the Stamford. In flat bottomeddeep tanks, with low beds, the shear layer tends to persist all the way to the side wall and theStamford can have a much more significant influence on the effluent quality.

CONCLUSIONS

An integrated CFD and mass flux modelling approach has been developed that may be used to

optimise the internal designs of final clarifiers. During the course of the development work anumber of key conclusions have arisen both with reference to current clarifier designmethodology and to current design practice.

Flows in secondary clarifiers are characterised by a strong density current arising fromthe influent that drives a radial shear layer above the settling sludge bed. Several flowfeature exist that must be designed out to optimise the tank performance.

The flow is far from one dimensional but it can be considered as close to twodimensional and in a circular clarifier there is axis-symmetry.

Tanks with sufficient surface area to satisfy mass flux theory will always fail before the

mass flux limit when modelled with CFD, in other words mass flux theory is the absoluteupper bound of performance.

The stilling well dimensions are important, too large a diameter or too shallow, can serveto enhance the density current momentum through re-entrainment. Two options arecurrently favoured as a means of breaking the density current. An EDI or a McKinneybaffle.

A central EDI helps to diffuse the density current by spreading the load uniformly nearthe top of the stilling well and reducing the density gradients in the stilling pond.

A McKinney baffle cuts the density current and, if designed with 0.5 < Frd < 0.7 itdistinctly separates the stilling zone from the settling zone introducing the flow as a directradial jet into the settling zone.

The McKinney design tends to favour a limited operating range as it provides maximumbenefit only when the baffle sits slightly above the settling sludge bed. In its bestoperating range it will tend to out perform an EDI.

Sidewall baffling has only a limited influence in shallow clarifiers but can be beneficial inflat bottomed clarifiers with deep side walls.

It is now possible to check clarifier retrofit and final design options with CFD modelling

prior to executing a civil engineering project.

ACKNOWLEDGEMENTS

The author would like to thank Dr Pete Pearce of Thames Water for his help, advice andcontinued support. Thanks also to other colleagues at Thames Water, United Utilities,Montgomery Watson Harza, Severn Trent and Yorkshire Water who have all supported projectscontributing to the overall understanding of internal clarifier flows.


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REFERENCES

[1] Chartered Institution of Water and Environmental Management (CIWEM), Activated SludgeTreatment, Handbooks of UK Wastewater practice, London, 1997.

[2] Anderson, N.E., Design of Settling Tanks for Activated Sludge, Sewage Works J., 17(1),50-63, 1945.

[3] Ekama, G.A., Barnard, J.L., Gunthert,F.W., Krebs,P., McCorquadale, J.A., Parker, D.S. andWahlberg, E.J., Secondary Settling Tanks, Theory, Modelling, Design and Operation,International Association of Water Quality, Scientific and Technical Report No 6, 1997.

[4] Richardson, D.S., Hydraulic Considerations of Final Settlement Tank Design, CranfieldUniversity School of Water Sciences, M.Sc. Thesis, 1998.

[5] Scriven, R. and Richardson, D.S. Rye Meads STW Stage 1 Final Settlement TanksProcess Investigation, Thames Water Technical Report, R19808, August 1998.

[6] Pitman, A.R. Settling Properties of Extended Aeration Sludge, J. Wat. Pollut. Control Fed.52(3), 524-536, 1980.

[7] White, M.J.D., Settling of Activated Sludge, Technical Report TR11, Water ResearchCentre, Stevenage, UK, 1975.

[8] Wahlberg, E.J. and Keinath, T.M. Development of settling flux curves using SVI J. Wat.Pollut. Control Fed. 60 (12), pp2095-2100, 1988.

[9] Wallis, G.B., One-Dimensional Two Phase Flow, McGraw-Hill, 1st Ed, 1969.

[10] Zhou, S. and McCorquodale, J.A., Modelling of Rectangular Settling Tanks, J. Hydr. Eng.,ASCE, 118(10), October 1992.

[11] Lakehal, D., Krebs, P., Krijgsman, J. and Rodi, W. Computing Shear Flow and SludgeBlanket in Secondary Clarifiers, J. Hydr. Eng., ASCE, 125(3), 1999.

[12] Armbruster, M., Untersuchung der mglichen Leistugssteigerung von Nachklrbeken mitHilfe numerischer Rechungen, PhD thesis, University of Karlesruhe, August 2003.

[13] De Clerq, B. Computational Fluid Dynamics of Settling Tanks: Development ofExperiments and Rheological, Settling and Scraper Sub Models, PhD Thesis, Dept of AppliedMath, Biometrics and Process Control (BIOMATH), University of Ghent, Belgium, 2003.

[14] Bokil, S.D. and Bewtra, J.K., Influence of Mechanical Blending on Aerobic Digestion of

Waste Activated Sludge, Proc., 6th Int. IAWPRC Conf. on Water Pollution Res., Int. Assoc. onWater Pollution and Control, London, 421-438, 1972.

[15] Dahl, C.P., Larsen, T. and Peterson, O., Numerical Modelling and Measurement in a TestSecondary Settling Tank, Water Sci. and Technology., 30(2), 219-228, 1994.

[16] Dick, R.I. and Ewing, B., The Rheology of Activated Sludge, J. Water. Pollution ControlFed., 39(4), 543-560, 1967.

[17] Takcs, I., Patry, G.G., and Nolasco, D. A Dynamic Model of the Clarification ThickeningProcess., Water Res, 25(10), 1991


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[18] CFX International, CFX-4.4 Solver Manual, Vol 3, AEA Technology, Harwell, 2001.

[19] Burt D.J. and Ganeshaligam, J. Validation study for the Witney Clarifier, MMI EngineeringReport, MMU035, February 2005.

[20] Vesilind, P.A. Theoretical considerations: Design of prototype thickeners from batchsettling tests, Water and Sewage Works, 115 (July), 302-307, 1968.

[21] Burt, D.J. and Gilbertson, M.A. Flocculation Frameworks and CFD Modelling for ActivatedSludge Clarifiers, 5th Particle Technology Forum, University of Sheffield, July 2003.

[22] Biggs, C. A. Activated Sludge Flocculation: Investigating the Effect of Shear Rate andCation Concentration on Flocculation Dynamics, PhD Thesis, Dept of Chem Eng, University ofQueensland, Australia, 2000.


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Figure 1: Taken from Anderson(1) shows concentration gradients through a circular clarifier andvelocity vectors mapped at discrete points.

Rye Meads Final Clarifier

Q Q

360.00 m3/h 360.00 m3/h

m3/h

RAS = RQ

360.00 m3/h

RecycleRAS_surplus

360.00 m3/h 0.00 m

3/h

Tank Diameter 28

SST Area (m2) 615.75

Forward Flow (Q, m3/h) 360.00

Design Overflow Rate (m/h) 0.58SSVI (ml/g) 100

Influent MLSS ( kg/m3) 2.825

RAS ratio 1.00

RAS_surplus (m3/h) 0

Solids Loading (l/m2h) 220

Estimated RAS Conc. ( kg/m3) 5.6500

Aeration Lane 720.00

Qin=Q(1+R)

Flux theory

0

20

40

60

80

100

120

140

160

0 1 2 3 4 5 6 7 8 9 10

Sludge Concentration (kg/m3)

Solids

Flux(kg/m2/day)

SSVI = 100 [Pitman and White (1980, 1984)]

SSVI = 100 [W ahlberg and Keinath (1998)]

Overflow Line

Underflow Line

Influent Conc.

Figure 2: Mass flux graph comparing Pitman(6) and White (7) with Wahlberg and Keinath(8) for theRye Meads clarifer at average flow and 100 SSVI.


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a)

0.35 0.5 0.64 0.78 0.96Density Waterfall

Radial Shear LayerShear Layer Separation

Bed Conveyor

Entrainment

0.35 0.5 0.64 0.78 0.96Density Waterfall

Radial Shear LayerShear Layer Separation

Bed Conveyor

Entrainment

b)

R / Rmax =0.35

0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1.0

-0.2 0 0.2

V / Uin [-]

H/Hmax[-]

Expt T 3

T 4 T 5

R / Rmax =0.50

0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1.0

-0.2 0 0.2

V / Uin [-]

H/Hmax[-]

Expt T 3

T 4 T 5

R / Rmax=0.64

0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1.0

-0.15 -0. 05 0.05 0. 15

V / Uin [-]

H/Hmax[-]

Expt T 3

T 4 T 5

R / Rmax =0.78

0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1.0

-0.15 -0. 05 0.05 0. 15

V / Uin [-]

H/Hmax[-]

Expt T 3

T 4 T 5

R / Rmax =0.96

0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1.0

-0.05 0 0.05

V / Uin [-]

H/Hmax[-]

Expt T 3

T 4 T 5

c)

R / Rmax =0.350.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1.0

0 1 2 3

C / Cin [-]

H/Hmax[-]

T 3 T 4

T 5 Exp

R / Rmax =0.500.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1.0

0

C / Cin [-]

H/Hmax[-]

T 3 T 4

T 5 Exp

R / Rmax =0.640.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1.0

0 1 2 3

C / Cin [-]

H/Hmax[-]

T 3 T 4

T 5 Exp

R / Rmax =0.780.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1.0

0 1 2

C / Cin [-]

H/Hmax[-]

T 3 T 4

T 5 Exp

R / Rmax =0.960.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1.0

0 0.5 1

C /Cin [-]

H/Hmax[-]

T 3 T 4

T 5 Exp

Figure 3: CFD analysis for Rye Meads, T3, T4 (14) and T5(11) represent alternative rheologicalmodels compared with the experimental data of Richardson et al (4)(5). The results show goodagreement for radial velocity profiles and bed height.


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a) Typical flocculating centre well arrangementwith central EDI, swirling ports and large, 30%D, stilling well.

b) Various alternatives for baffling the sidewall. The Crosby is also sometimes called aStamford.

Figure 4: a) A central EDI, spreads the load out uniformly near the top of the stilling wellreducing the density variation. b) The return current at the side wall can be disrupted by variousbaffle options.

a) Clarifier as built with a 2m (12%D)diameter stilling well.

b) Revised design with flocculatingstilling well at 20%D and enclosingan EDI. The EDI is a variation onFigure 4a.

c) A McKinney baffle design withstilling well at 20%D and gap sizedfor densimetric Froude number, Frd

1, at the highest flow rate.

Figure 5: A typical UK Clarifier diameter 16.76m. Comparisons between designs are presentedhere for a forward flow of 156.3 m3/hr, mixed liquors at influent of 3700 mg/l, a RAS ratio of 1.0and SSVI around 80 ml/g. A logarithmic scale is used to show the solids distribution 1 mg/l to10,000 mg/l. The lightest shading at 1000 mg/l is approximately at the transition into the settledbed.

CROSB Y VARIATI ON

McKINNEY

PLAIN

CANTILEVERED

TROUGH


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With Stamford baffle Without Stamford baffle

Figure 6: A Stamford baffle was included for the EDI design shown in Figure 5b. The flowpatterns near the effluent are redirecated but there is no significant difference in the effluentquality.

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