The activated sludge process part 2. Application of the general kinetic model to the contact stabilization process

WmtTgaem.t'k Vol. 14, pp. 1737 to 1747 INwlptmonPrms Ltd 1990. Printed in Great Britain

THE ACTIVATED SLUDGE PROCESS PART 2. APPLICATION OF THE GENERAL KINETIC MODEL TO THE CONTACT STABILIZATION PROCESS W. V. ALEXANDER1, G. A. EKAMA2 and G. v. R. MARXtS~ 1Scott & de Wanl Inc., Johannesburg and "Water Resources and Public Health Engineering, University of Cape Town, Rondebosch, South Africa

(Received August 1979) Akatzagt--The general aerobic activated sludge model including nitrification for pr _oo~__~streating principally municipal wastewnter is applied to the contact stabilization process treating municipal wastewaters. The application involves two chanlp= to the model: (i) a change in one of the values of the kinetic constants in the e x ~ of the substrata utilization rates; (fi) a change in the enmeshment ma:hmigm by a ~ that a fraction of the particulate COD which is mot admrbed onto the active m'ganilm~ does not.become mmeslmi in the sludge floes and escapes with the effluent. Accepting only these dumges it was found possible to satisfactorily simulate the behaviour of the contact and stabilization reazto~ of the pro©ess under both constant and cyclic conditions of loading. For design, the general activated sludge model, as applied to the contact stabilization process, requires the Woet~8 configuration to be completely specified. To aid in the initial design of the process, a pr01il~ design procedure is presented by means of which the volumes, sludge concentrations and retehtion times of the contact and stabilization reactors may be determined from five independent iamtmeters whist are m u m e d t o govern the proee~ These are the sludge age, recycle ratio, fractional distribution of the sludge mass between contact and stabilization reactors, dally COD mass load and the average sludge concentration in the process,

NOMENCLATURE bh* - a n g u s n~q.piration rate (d-l) CMASP - Completely Mixed Activated Sludge Process CSASP ,- Contact Stabilization Activated Sludge Proosss f m anbiodegrudable fraction of the active mass (mg VSSmg VSS -x) f , I, tfitrolpm fraction of the sludge mass (mg N mg VSS -s) fup " fraction of total influent COD which is in the particulate unbiodegradable form f., ,= fmct/on of the total influent COD which is in the soluble unbiodegradable form Kin* =, maximum specific substrata utilization rate for readily assimilable COD (mg COD mg VSS- x d - ~)

M = prefix denoting mass, i.e. M(Su) = Su" Q == daily COD mass (ms COD d-x)

load

M(X~) = X,.,Vp

mass of sludge in process (rag VSS) N =. general symbol for nitrogen concentration (mg N 1= xg Subscripts n or t refer to nitrate or TKN concentrations respectively. Additional subgripts c~ e, i or s refer respectively to the values in the contact reactor, effluent, influent or stabilization reactor * Additiomtl subec~pt Tot 20 refers to the value at T°C or 200C. 1737

0 -- general symbol for oxygen consumption rate (mgO1-1 h-X). Subscripts c, n and t refer to carbonaceous, nitrification and total values respectively. Additional subscripts c, p and s refer ~ e l y to values in the contact reactot, overall procoss and stabilization reactor P -= COD to VSS ratio (mgCODmgVSS -t) Q = influent flow rate (ld -s) q = waste flow rate 0 d - x) Rt = hydraufic retention time (d). Subscripts a or n refer to actual or nominal mention times respectively. Additional subscripts c or s refer to values in the contact or stabilization reactor respectively R, = sludge age (d) r = recycle ratio with respe~ to average influent flow S = gmeral symbol for COD concentration (rag COD I- s). Subscript t refers to total concentrationL Additional subscripts c, e, i, s refer respectively to the values in the contact reactor, elHuent, influent and stabilization reactor T = temperature in °C V = general symbol for volume 0.~ Subscripts c, p and s refer tO contact reactor, overall process and stabilization reactor respectively w = waste I]ow ratio with respect to average influent flow X = smefal symbol for volatile sludge ~ c e n tration (mgVSSl-1). Subseripts a, n and v refer to active volatile, nitrifier and total volatile sludge concentrations respectively. Additional subscripts c, p and s refer respectively to valties in the contact reactor, overall process and stabilization reactor

1738

W.V. ALEXANDER,G. A. EKAMAand G. v. R. MARAIS

Yh= yield coefficient (rag VSS mg COD - ~) = fraction of mass sludge in process in the contact reactor 0 = temperature sensitivity coefficient ~ , , * = maximum specific growth rate of the nit riflers (d- ')

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INTRODUCTION STASILtZATIO

In the literature, the kinetic theories describing the behaviour of the contact stabilization activated sludge process (CSASP) have been essentially of an ad hoc nature. No attempt to describe the behaviour of this process by a general activated sludge kinetic theory has been successful, principally because the theories proposed to date are deficient in themselves. However. the general hi-substrate death-regeneration aetivated sludge kinetic theory for the process treating principally municipal wastewater, proposed by Dold et al. (1980), appears to be sufficiently comprehensive to encompass also the kinetic behaviour of the CSASP treating mainly municipal wastewaters. This paper describes the integration of the CSASP with the general activated sludge theory, In presenting this paper, it is assumed that the reader has studied the paper setting out the general activated sludge theory (Dold et al., 1980)--only the aspects necessary for the particular requirements of the CSASP will be presented, INITIAL PROCESS DESIGN A diagrammatic representation of the CSASP is given in Fig. 1. The contact reactor receives the influent flow and the recycle from the stabilization reactor. In the contact reactor, which has a short nominal hydraulic retentionf, a small fraction of the sludge mass removes the bulk of the carbonaceous material from the influent--principally by adsorption and enmeshment. A minor fraction of the adsorbed organic material is metabolized in the contact reactor and consequently the sludge mass contains a high proportion of unmetabolized COD. The effluent from the contact reactor is discharged to the settling tank. The settling tank overflow leaves the process as effluent; the undertow is discharged to the stabilization reactor, which has a long nominal retention time in which the densifled mixed liquor is stabilized. Stabilized sludge is returned to the contact reactor via a sludge recycle flow. Close control of the sludge age of the process is possible by hydraulic control, i.e. by wasting an appropriate volume of mixed liquor directly from either the contact or stabilization reactor, The advantage of the CSASP over the Completely Mixed Activated Sludge Pro~ess (CMASP) stems from its configuration. The major fraction of the ~"N ~ i n a l hydraulic retention time is given by the volume -~ the reactor divided by the daily average influent flow.

nEACTO~ Fig. 1. Diagrammatic representation of the contact stabilization process. sludge mass in the CSASP system is contained in the stabilization reactor in a densified form, giving a correspondingly smaller volume to the whole system. Furthermore, if the excess sludge is wasted from the contact reactor, as shown in F.if, 1, the overall process oxygen demand is slightly reduced because the sludge mass in the contact reactor contains a relatively high fraction of unmetabolized COD. The CSASP is a spatially del)endcat system even under steady flow and load c o , ditions. The concerttration of sludge and the food/micro-organism ratio differ significantly in the contact a n d : s ~ l i z a t i o n reactors; these parameters are depfndeat on the sludge recycle rate and the fractional distribution of the mass of sludge between the two reactors. Therefore, unlike the CMASP which essentially is defined by the sludge age and the orgamc toad in the process, in the CSASP the process definition must include, in addition to the above parameters, the recycle ratio, r, and some parameter fixing the fractional distribution of the mass of sludge betwe~l the contact and stabilization reactors. Ohron & Jenkins (i972) selected.the fraction of the mass of sludge in contact reactor relative to the total mass of sludge in the process, ~, to define the fractional parameter, i.e. ~t = VcXvc/(VcX,c ~- V,X~,)

{1)

where Vc, Vs = volumes of the contact and stabilization reactors respectively (1.) X,0, X,, = volatile sludge concentrations in the contact and stabilization reactors respectively (mg VSS 1-1). It will be apparent from the above, that the design of a specific CSASP is not as straightforward as the CMASP and must proceed on a trial and error basis. In designing a CSASP configuration it is particularly useful to have a procedure for initially estimating the sizes of the reactors and their respec(ive sludge conoentrations in order to decide whether the liquid solid separation efficiency in the settling tank and the actual hydraulic retention time in the contact reactor are within acceptable limits. A basis for making an initial estimate of the reactor sizes is to accept that the C S A S P prodlg~:tke'~ame mass of sludge as the C M A S P for the same sludge age,

The activated sludge process--Part 2 R,, and daily organic mass load, M(S~). This assumption is based on the findings of Gujer & Jenkins (1975a) and has been verified during this investigation. Therefore, an estimate of the total mass of slUdge in the process M(X~) can be obtained from the steady state equaU'ons i'or the CMASP (Marais & Ekama, 1976), i.e.

M(X.p) = M(S,)R,

advantage, Higher X,p concentrations for the CSASP are permissible because the settling tank of the process has to deal with the relatively low concentration of sludge from the contact reactor. Where the sludge age, R,, of the CSASP is controlled by wasting sludge directly from the contact reactor, i.e. hydraulic control of the sludge age, the sludge age of the process is given by

{~(l+fbhrR,) + bhrR,)

R, = 3

x (1 - fu. - fupP) + fup/

1739

(2)

mass of sludge in process mass of sludge wasted per day

--- XvpVp/(Xvcq)

(6)

where M(Xv)--mass of volatile (rag VSS)

sludge in

process

where q = waste flow rate (1 d - t)

= XvpVp Xvp = average process volatile sludge concerttration (rag VSS 1- J) Vp = volume of process (1.) M(Stl) = daily organic mass load (mg COD d - l )

= Q" Sit Sti - total influent COD concentration (rag COD 1-1) R, = sludge age (d) Yh - yield coefficient (rag VSS mg COD-1) = 0.45 f = unbiodegradable fraction of the active

In the analysis of the CSASP significant simplification in the formulations is attained by expressing q as a fraction of the average influent flow rate, i.e. q = wQ

(7)

where Q = average influent flow (1 d - 1) w = sludge waste flow rate from the contact reactor as a fraction of the average influent flow rate. Hence, for a given sludge age, w is given by

volatile mass (mg VSS mg VSS- 1)

w -- XvpVp/(RsQXvc)

= 0.20 bhr = endogenous respiration rate at T°C (d)

In order to size the two reactors of the CSASP, a decision on the two additional independent par-

_._

bh2o = 0= T= fu, =

(8)

b h 2 o ( T T - 2o)

endogenous respiration rate at 20°C 0.24 temperature sensitivity coefficient 1.029 temperature (°C) soluble unbiodegradable fraction of the influent COD (rag COD mg C O D - 1)

f~p = particulate unbiodegradable fraction of the inttuent COD (rag VSS mg C O D - i) P = COD to VSS ratio (mgCODmgVSS -1) Once an estimate of M(Xvp) is available, a decision must be made on the overall mean sludoe concentration, Xvt# for the proces~ The overall volume of the CSASP, Vp, is determined by Vp --- M(X~p)/Xvp

,

(3)

where

ameters ~ and r is required. Knowing ~, r, R8 and X,p as well as Vp, which can be determined from equations (2) and (3), then by doing a sludge mass balance around each reactor and manipulating equations (4)-(7), the concentration of the sludge in, and the volumes of the two reactors can be expressed in terms of the five known parameters, i.e. Vc = Vpa(1 + r - w)/(~t + r - otw)

(9)

V, = Vpr(1 - ~)/(~ + r - ~tw) X,c = Xvp(~ + r - uw)/(1 + r - w)

(10) '(11)

X~, = X~(~ + r - ~w)/r X,, = X~(1 + r - w)/r.

(12) (13)

An expression for w in equations (9)-(13) can be derived as follows: Substituting equation (5)for VpXvp in equation (8):

Vp - V, + V~

(4)

w = (X~cVc + X~,V,)/(R,QXvc).

(14)

and Substituting equation (13) for Xv, and equation (I0) VpXvp = VcXvc + VtXvs. (5) If X,p for the CSASP is chosen to be the same as that for a CMASP the volumes of the two processes will be identical, resu!tin_g m " n o space s a ~ Consequently, ~,p for the' CSASP usually is selected to be appreciably higher ~than that normally accepted for the CMASP in order to gain a plant volumetric

for V, into equation (14) and rearranging, yields: w = Vp(1 + r - w)/{R,Q(,, + r - ~tw)}.

(15)

Under normal operating conditions, ~ov is small relafive to (c( + r~ Consequently, ignoring ~,w in equation (15)and solving for w, an expression for w in terms of

W.V. ALEXANDER,G. A. EKAMAand G. v. R. MARAIS

1740

the known parameters is obtained, i.e. w = (1 + r)/{ 1 + (~ + r)QRs/Vp}.

(16)

F r o m equations (9) and (10), the actual and nominal retention times in the two reactors can be found, i.e. R~c

w)/[(1 + r)(0c + r - 0~w)~] (17) • Rhas = Vp(I - ~t)/[(~ + r - 0tw)Q] (18) Rh.¢ = (1 + s)" Rhoc (19) = Vp(1

+

r -

Rhns = Rh~

(20)

where Rh = hydraulic retention time (d). Second subscripts a or n refer to actual or nominal values respectively. Additional subscripts c or s refer to contact or stabilization reactors respectively. F r o m the equations set out above estimates of Vc, V,, X,o, and X,, in terms of the values a, r, M(St, g X~p and R, are now available a n d constitute an initial design. Checks can be made to ensure that the actual hydraulic retention in the contact reactor is adequate and that the densification ratio in the settling tank is not too great, Experience with experimental plants and the general activated sludge model as applied to the CSASP, has indicated that this preliminary design procedure predicts X~c and X,, accurately. Also, the average process carbonaceous oxygen demand for the system approximately equals that for the equivalent C M A S P [i.e. for the same R, and M(S~)]. However, the procedure provides no estimation of the relative Oxygen demands in the two reactors, nor the C O D , T K N and nitrate concentration in the reactors or e~uent under constant or cyclic loading conditions. Under constant flow conditions these process variables may be estimated from models with constants determined on an ad hoc basis (Gujer & Jenkins, 1975a, bg However, under cyclic loading conditions the only reliable procedure for determining these process variables is by utilizing the general activated sludge model as applied to the CSASP. EXPERIMENTAL INVESTIGATION It has been shown that there are five major parameters which influence the response of the CSASP, i.e. dm'ly COD load. M(S~), sludge age, R,, temperature T, sludge recycle rate, r, and the fra~ioaal dlairibetion of sludge mass between the contact and stabilization reactors in terms of, a. For this investigation it was ~ that t h e temperature, r and a would be kept cezf,taat, while the :O~lge age and daily COD load pattern would be varied. The reason for maintaining r and a constant was that Gujer & Jenkins (1975a) reported that these two parameter-have less effect on the efficiency of carbonaceous material removal by the prooeM than the' Other tiwee. ~ ~ v t [ i l i ~ ~ tion is also consistent withthe bi-m~l~mt¢ ~ ~ et al., 1980), Although r and 0c have a marked ~ 9~a the contact reactor hydraulic retention time, the i~'tteulate

carbonaceous substrate, which forms the bulk of the influent COD, is either adsorbed by the organisms or enmeshed in the sludge floes and removed from the effluent in the settling tank. Adsorption and enmeshment are rapid processes, so that they are expected to be near completion even in relatively short actual hydraulic retention times. The removal of the soluble COD fraction in the influent by the contact reactor is affected by the length of the contact time: any soluble COD concentration not metabolized during the contact period in the contact reactor escapes with the effluent flow. However, as only about 25% of the biodegradable COD is in soluble form (Dold et al., 1980) and the removal is rapid, the escape of the soluble COD concentration is unlikely to affect the overall COD removal efficiency. With regard to ~, a value of 0.1 was selected as with this value only one tenth of the mass of sludge in the process is responsible for the initial removal of the carbonaceous material from the influent. It represents, therefore, an extreme situation and constitutes a severe test of the predictive capacity of the model With regard to the choice of r, a high value (r > 4) in the CSASP results in a smaller relative difference between the concentrations of sludge in the contact and stabilization reactors and the process approximates the CMASP which would tend to detract from the uniqueness of the CSASP configuration. In contrast, a low recycle ratio (r < 1) results in a low process nitrification efficiency, as only a small fraction of the influent ammonia is recycled to the stabilization reactor in which the bulk of the sludge mass is retained (Gujer & Jenkins, 1975b). Consequently, a recycle ratio of 2:1 with respect to average flow was selected. The experimental investigation was to be undertaken at 20°C. To ensure complete nitrification of the recycled ammonia in the stabilization reactor, sludge ages longer than about 4 days are required. Consequently, sludge ages of 6 and 10 days were selected. If the CSASP operates under cyclic flow conditions, the actual hydraulic retention time in the contact reactor is reduced during the peak flow period. Should the actual hydraulic retention time under peak flow conditions become too short, particulate COD removal by adsorption and enmeshment and soluble COD removal by metabolism will be insufficient, resultin$ in a ~ o r effluent quality. Consequently a lower limit to/lie ~ i h / U / f i a~c/~-i hydraulic retention time in the contact reactor needs to be set. To satisfy this lower limit, it will be found that lower values of X~p or higher values of a need to be chosen as the peak/average flow ratio increases, than would be required under the average flow conditions. In this experimental investigation, the minimum actual hydraulic retention time in the contact reactor was set at about 20 min under an expected peak/average flow ratio of 1 : 1.5. With this restriction, the chosen design parameters a, r, R, and M(S,t), the design of the experimental CSASP units was obtained by a trial and error procedure. All the conditions could be met only when X~p was less than about 2000 mg VSS 1-1 a relatively low value. By doubling a to 0.2, and keeping the minimum actual hydraulic retention time in the contact reactor at 20 min under the peak/average flow ratio of 1:1.5. X~ can be increased to 3800 mg VSS 1- :, giving a more practical design. However, from a model verification point of view, the latter design would not have provided as severe a test on the predictive capaoty of the model as the former. The chosen design parameters of the experimental units are given in Table 1 and a diagrammatic layout of the units is shown in Fig. 2 The experimental investigation was divided into two phases: (a) a phase during which the two experimental units were operated under constant flow and load conditions, and (b) a phase during which one of the units was operated under cyclic flow and load conditions.

The activated sludge process--Part 2

1741

Table 1. Process design parameters for the Laboratory CSASP units Name

Parameter Symbol

1. Sewage type 2. I~fluent COD 3. influent TKN 4. Average influent flow 5. Temperature 6. Sludge age 7. Waste flow ratio 8. Fractional distribition of sludge in process 9 . Recycle ratio 10. Volumes Process Contact reactor Stabilization reactor 11. Average hydraulic retention times Con tact: Actual Nominal Stabilization: Actual Nominal 12. Length of test period

Sti Nil Q T R, w a

mg COD 1-1 mg N 1-1 I d- ~ °C d ---

Sludge age 6 days 10 days Unsettled 500 ~ 50 36 20 6 0.092 0.10

Unsettled 500 ~ 50 24 20 10 0.083 0.10

2.0

2.0

Vp Vc V, Rh

1. 1. 1. 1. h

14.3 2.0 12.3

14.3 2.0 12.3

Rt,© R~c

h h

0.44 1.32

0.67 1.98

R~,, R~,, --

h h d

4.07 4.07 23

6.27 6.27 25

r

(a) Constant flow and load conditions were established by feeding the required volume of sewage feed per day at a constant rate over the day. The average values of the daily measurements of the process variables in the contact and stabilization reactors and in the effluent of the 6 and 10 day sludge units are given in Table 2, the length of the test period having been about 25 days. (b) In the cy¢fic loading tests the 6 day sludge age unit was operated under two different influent cyclic flow patterns: (i) a square wave pattern, and (ii) a sine wave pattern. In both cases the daily influent COD mass was the same and equal to that fed during the constant flow test, i.e. 18,000mgCODd -t. (i) The square wave loadino pattern was obtained by

i

Units

--

dividing the daily influent feed volume of 361. between two pumps--one operating at a rate of 181 d - : over the full 24 h period and the other at 361 d- 1 for a period of 12 h only. Measurements of the process variables in the contact and stabilization reactors and in the effluent were made at regular intervals over a 24 h period--a typical set of data is shown in Fig. 3 . (fi) The sine wave ioadin 0 pattern was achieved by means of a specially designed pump that produced an approximate sine wave flow with an amplitude of about 0.5 times the average flow. A typical set of data observed over a 24h cycle under sine wave influent flow conditions is shown in Fig. 4.

~ CONTACT R|ACTO~

i "

I SETTLIr~ EFFLUENT RI~C~C LE

Fig. 2. Diagrammatic representation of the laboratory scale contact stabilization process.

W.V. ALEXANDER,G. A. EKAMAand G. v. R. MARAIS

1742

Table 2. Averages of process variable measurements observed in laboratory scale contact stabilization units. The corresponding predictions of the general model are also given Constant flow and load conditions Sludge age Process variable 6 Days l0 Days Position Symbol Measured Predicted Measured Predicted COD mg COD l Influent Contact Effluent Stabilization TKN mg N 1- l

Influent

Sti* Stc St¢* Sts

512 62 90 54

Nti*

Contact Ntc Effluent Nte* Stabilization Nts Nitrate mg N 1- l Influent Nn~ Contact N,c Effluent N,, Stabilization N,s MLVSS mg VSS 1-1 Contact X,¢ Stabilization X, Process X,p Oxygen Demand (mg O 1~ z h - l ) Contact O~ Carbon ~ Stabilization O¢~ I Process Oep Contact On, Nitrif. ~ Stabilization O,~ [. Process O.p ( Contact Ot~ Total ~ Stabilization Ot~ {, Process O,, Mass balances COD+ N:~

512 60 81 45

496 52 61 42

496 57 75 43

56 11 15 1.7

56 15 15 3

47 10.5 12 4.8

47 12 12 3

0.0 27 27 36

0.0 27 27 37

0.0 22 23 29

0.0 23 23 31

1301 1772 1706 31 22 23 30 10 13 61 32 36 96% 94%

1489 2099 2016 29 23 24 24 11 13 54 34 37 100% 100%

1478 2073 1990 24 16 17 18 5 7 42 21 24 97% 96%

1389 1989 1903 23 16 17 19 5 8 42 22 25 100% 100%

* Refers to unfiltered samples. t" Based on a COD/VSS ratio of 1.48. :~ Based on TKN/VSS ratio (f,) of 0.10.

MODEL

VERIFICATION

The general kinetics of the activated sludge process including nitrification described by Dold et al. (1980) were incorporated into a general model for the CSASP (in the form of a computer program). When this model, as well as its associated kinetic constants, was applied to simulate the behaviour observed in the experimental CSASP units described above, the following observations were made: Charu3es in kinetic constants

The predictions of the response of the contact reactot is very sensitive to the values of the maximum specitic growth rate of the nitrifiers (/~m2o) and the maximum utilization rate of soluble carbonaceous substrate (Km~zo). With regard to P~m20, Dold et al. found an averacje value of 0.65 d - ~ for data. The value o f / z w z o associated with each batch of sewage can be determined by obtaining the best fit between experimental and pre-

dicted T K N and nitrate concentrations. F r o m a comparison of many e x p e r i m ~ t a t and predicted responses, it became evident that the value of /q,,z0 often changes for each batch of sewage probably due to changes in the content of inhibitory substances. In this investigation; it was found that the best overall correspondence between experimental and predicted T K N and nitrate concentrations was obtained when P~=2o = 0.55d-~. All the CSASP simulations are Imsed on this value. With regard to K,,2o, Dold et al. found that a value of 8.0 nag C O D mg VSS d - 1 gave the best correspondence between experimental and predicted oxygen consumption rates for single and series C M A S processes. When this value for Km~2o was utilized to simulate the behaviour of the CSASP under cyclic loading conditions it was found that the prcgiicted oxygen consumption rate in the contact rca~tor (Otc) showed a marked cyclic variation over the d a y - - a result in conflict with the experimental obs~vations. Also, the variation in the predicted soluble C O D con-

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1743

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Fig 3 Comparison of experimental and predicted responsesof the contact stabilization process under daily cyclic square wave loading conditions at 20°C centration in the contact reactor (S,=) showed a far 8reater degree of attenuation than the experimental

ever the need for changing K=,~o for the CSASP is indicative of a deficiency in the formulations of the

dam. The best correspondence between predicted and experimental Or= and St= responses was obtained when Km2o = 2.0 mg COD mg VSS d - L When this value was incorporated in the CMASP model, predictions of the precipitous decrease in the oxygen consumption rate at feed termination in square wave cyclically loaded single reactor CMAS units (see Dold et al., 1980) was not accurately predicted. No explanation can be advanced for this apparent discrepancy between the Km,~o values for the CSAS and CMAS processes, The kinetic constants / ~ o and K=,,0 were the only two which required adjustment of their values, The need for adjusting ~m~O is not necessarily indicative of a deficiency in the formulations of the general model, because there appears to be strong evidence that P~=~o does in fact change not only between different sources but also between batches of sewage obtained from the same source---the sensitivity of the nitrifiers to inhibitory or toxic substances in the wastewater is well documented in the literature. How-

general model. The lowering of Km,20 indicates that the soluble carbonaceous substrate is utilized at a slower rate in the contact reactor of a CSASP than in the first reactor of a multiple CMASP. However, once the different value of Km,2o is accepted, the general model is capable of describing the CSASP under both constant and cyclic conditions of loading. Modification to enmeshment mechanism A conceptual modification had to be made to the general activated sludge model in order to satisfactorily simulate general CSASP behaviour. In the experimental data observed in the CMASP, it was found that there was very little difference between the filtered and unfiltered effluent COD concentration. Consequently, in the general model, it was assumed that all the unadsorbed particulate COD remains enmeshed in the sludge flocs, is densified with the sludge in the settling tank and recycled to the reactor. However, the CSASP experimental data (see Table 2 and Figs 3 and 4) show a marked difference between

1744

W.V. ALEXANDER,G. A. EKAMAand G. v. R. MARAIS

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Fig. 4. Comparison of experimental and predicted responses of the contact stabilization process under daily cycfic sine wave loading conditions at 20°C. filtered and unfiltered effluent COD concentrations, the unfiltered values being always significantly higher than the filmed values*. This appeared to indicate that the floceulation and ~ t of the particulate substrate was not complete in the contact reactor, This effect was empirically incorporated into the model by hypothesizing that a fraction'of the particulate COD not adsorbed onto the active mass does not become enmeslmt in the sludge floes and e~capes with the effluent. A good correspondence between predicted and measured filtered and unfiltered effluent COD concentrations was obtained when this fraction was a~=umod to be 0.50. The ~ appears to have relevan(~e to only the C S A ~ because in series or single rea~or CMASP, the rete~ltion times are sufilciently long and sludge conceatrations sufficiently high for adsorption and enmeshment to be virtually complete before solid-lk~id separation takes place,

• The comparison is ~ om the ~ that the filtered contact and Wtefed ~ t COD concentrations are equal.

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General CSASP model The changes discussed above were accepted and incorporated in the CSASP model. The theoretical predictions of this model with the correspondi~ experimental data observed under constant and cyclic loading conditions at 20°C are given in Table 2 and Figs 3 and 4 respectively. The comparison of the predicted and experimental data is discussed below. COMPAmSON OF EXPERIMENTALAND THEORETICAL DATA Constant flow conditions (Table 2) In general there is a good correspondence between predicted and experimental results at both 6 and 10 days sludge age. The predicted sludge concentration in the 6 day sludge age unit is higher than the corresponding experimental data. However, subsequently, when the unit was Opersted u n d ~ cyclic l o a ~ conditions with the sludge age remaining at 6~days, the experimental sludge conoentgations s t a ~ at vabJcs very close to those predicf~i by the mod~ ~tee Figs 3 and 4). Hence the low experimental vatnes

The activated sludge process--Part 2

1745

measured during this period are likely to be a conse- daily cycle, the recycle ratio varied over the cycle quence of the variation in behaviour usually associ- resulting in fluctuating MLVSS concentrations in the ated with biological systems, contact stabilization reactors. For example, under the The over-prediction of the T K N concentration in square wave loading conditions (Fig. 3), the recycle the contact reactor, Ntc, appears to indicate that more ratio was 4:1 (2 x 36/18) during the low flow period nitrification took place in the experimental unit than and 1.33:1 (2 x 36/54) during the peak flow period. the model predicts. However, this is not reflected in The period of high recycle ratio (4:1) caused the the measured and predicted nitrate concentrations. CSASP to approach the CMASP in behaviour so that The reason for this discrepancy is that the nitrogen the difference between the sludge concentrations in fraction of the waste sludge, f,, in the experimental the contact and stabilization reactors was small. In unit was slightly higher, i.e. 0.12mgNmgVSS -l, contrast, the low recycle ratio period causes a prothan the value incorporated in the general model, i.e. nounce'd deviation from the CMASP so that the dif0.10 mg N mg VSS-~. This results in a greater mass of ' ference between the sludge concentration in the connitrogen being removed from the experimental unit tact and stabilization reactors is large. via the waste sludge than predicted by the model. 3. Oxygen consumption rate. In the contact reactor, This difference is reflected in the effluent T K N con- the experimental total oxygen consumption rate, Otc centrations. Furthermore, by assuming an fn value of remains virtually constant throughout the cycle. 0.12 mg N mg VSS- ~ for the experimental unit, an im- Compared to the CMASP, such a high degree of atproved nitrogen balance will be obtained (Table 2). tenuation in the oxygen demand is unexpected under The experimental nitrification oxygen demand may cyclic influent loading conditions~ However, considerbe calculated from the experimental data by consider- ing square wave loading conditions (Fig. 3) this being a nitrate mass balance over the contact or stabillz- haviour can be explained as follows: Upon comation reactors. The carbonaceous value is given by the mencement of the peak flow period, the availability of difference between the total and nitrification values. A carbonaceous and ammoniacal substrates increases in comparison between the experimental and predicted the contact reactor, resulting in a greater activity of values is given in Table 2. The contact nitrification the heterotrophic and nitrifying organisms. However, oxygen demand (One) calculated from the nitrate concomitant with the higher sludge activity, the measurements (N~c), is extremely sensitive to changes sludge concentration (X~¢,X~¢) is reduced. The net in the values of these measurements--a 1 mg N 1- ~ result is a relatively constant total oxygen consumpdifference in N,c results in a 10.7 mg O 1- ~ h- ~ differ- tion rate. ence in O~c. In contrast, the nitrification oxygen Dividing the experimental total oxygen consumpdemand in the stabilization reactor (O,,) is relatively tion rate measured in the contact reactor (Or©) into insensitive to changes in the input and output nitrate the carbonaceous (O©c) and nitrification (O~) cornconcentrations. Taking into account these factors, it is ' ponents by deducting One from Or© led to widely flucclear that there is a good correspondence between the tuating results----O~c is extremely sensitive to small model predictions and experimental observations, variations in the measured nitrate concentration. Consequently, only the predicted and experimental Cyclic flow and load conditions (Figs 3 and 4) Ot~ responses could be reliably compared. In general the correspondence between the preThe model satisfactorily predicts the observed atdicted and experimental responses under both square tenuated Ot~ response. However, under square wave and sine wave loading conditions is very good. A loading conditions (Fig. 3), the model incorrectly predetailed discussion of the comparison of each process dicts an increase in Ot~ immediately after cessation of variable is given below, the peak flow period. In the theoretical model this 1. COD concentration. As in the constant flow and. increase is caused by the increase in sludge concenload tests, a difference between the filtered COD con- tration (X~, X~) at a time when the carbonaceous centration in the contact reactor and the unfiltered and ammoniacal substrates concentrations are still value in the effluent is apparent. It appears that the high from the peak flow period. Considering the difflempirical partial enmeshment modification in the culties associated with accurately measuring oxygen CSASP model allows accurate predictions of the effiu- consumption rates in the contact reactor (Ekama & ent COD concentration (both filtered and unfiltered) Marais, 1979), the discrepancy between the predicted even under different cyclic loading conditions, and experimental Or© responses is as likely to be a 2. M L V S S concentration. The correspondence consequence of a deficiency in the theoretical model between the predicted and experimental MLVSS con- as experimental error. centrations is excellent in both cyclic tests. An interIn both cyclic tests, the total oxygen consumption esting_L observation is the significant "washing out" rate in the stabilization reactor (Oa) is slightly overeffect of the MLVSS from the contact reactor during predicted. Dividing the experimental Os response the peak flow period. It was stated above that as the into the carbonaceous (Oo.) and nitrification (On,) recycle ratio increased in the CSASP, the process components is possible because On. is relatively inapproached the CMASP. As the recycle flow was kept sensitive to small variations in the nitrate concertat twice the average influent flow throughout the tration (N..). Considering the square wave test, a

W.V. ALEXANDER,G. A. EKAMAand G. v. R. MARAIS

1746

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This paper provides a means w h ~ b y , i ~ i t i a l ¢stimates can be mad concentrations in tors of a CSASP from the six i n ~ t : ~ p a r ameters assumed to govern the i a c h a ~ o u r ' d t ~ process. i.e. sludge age, R,, recycle ratio, r, ~ a t distribution of the sludge mass-botwcen the two reactors defined by ~, daily COD mass load, M ~ ) and the average process sludge concentration, X~. The sixth design parameter, i.e, temperature, was not considered in this paper, Two changes to the general activated ~ a d p kixmtic model proposed by Dold et aL (1980) w ~ e ~ i.e. (1) a change in the value of one of the kinetic constants in the expressmns of the carbonaceous sub-

strate utilization rates; and (2) a change in the enmeshment mechanism by accepting that a fraction of the particulate COD which does not become adsorbed onto the active mass, does not become enmeshed in the sludge floes and escapes with the effluent. With these changes it was found that the model is sufficiently general to give a good description of the behaviour of the CSASP under constant and cyclic conditions of loading. The investigation into the cyclic behaviour of the CSASP indicated that the peak flow periods reduce the carbonaceous material removal and nitrification efficiencies of the process in two ways, i.e. by reducing (i) the actual hydraulic retention as well as {ii) the Sludge concentration in the contact reactor. This behaviour of the process results in the effluent COD and T K N concentrations being very sensitive to cyclic loading conditions. More stable effluent qualities under cyclic loading conditions may be produced by increasing the design parameters • and r and reducing X~. However, such changes reduce the specific advantagea that the CSASP has over the CMASP, and prodace configurations coafonning more to the behavioaral characteristics of the ~ S P . A major ~ l t y in a ~ , ~ g carbonaceous material removal a n d nitrification models for the CSASP, for ¢xa~ Jenkins (1975a, b), t( the values of the models under the environnmatal cmalitiotts in the field. These difficulties are also m c o t m ~ ~.4he application of the computer m, altltough the kinetic constants of material removal mcchanimas unchanged for different domestic sewages, the process

The activated sludge process--Part 2 response does depend on the influent sewage characterisfics, such as the unbiodegradable soluble and particulate COD fractions ( f , and fup respectively). Also, the maximum specific growth rate of the nitrifiers at 20°C (/~m,o) has been found to vary considerably between different sewages. It appears that the only way of overcoming these difficulties is to determine these constants in laboratory scale investigations utilizing the wastewater to be treated, A further difficulty in applying the existing models of the CSASP to full scale design is an estimation of the influence of cyclic loading conditions on the process behaviour. Once the sewage characteristics have been established, the general activated sludge theory as applied to the CSASP allows the prediction of the dynamic behaviour of the process under cyclic loading conditions. This model is very cumbersome and can only be utilized in the form of a computer program. Design engineers prefer simple design charts and models. Existing models and the initial design procedure presented in this paper satisfy this preference. However, dynamic solutions of the process behaviour cannot be obtained by simplistic models, computer models have to be employed, if such solutions are required. The use of a simplistic model for the preliminary determination of the independent process

1747

variables and subsequent analysis of the dynamic behaviour of the process by computer models seems a satisfactory compromise and indeed appears to be a logical approach to design. Acknowledgements--This research was carried out under contract with the Water Research Commission of South Africa. The authors wish to thank the Commission for permission to publish this paper.

REFERENCES Dold P. L., Ekama G. A. & Marais G. V. R. (1980) The activated sludge process Part I. A general model for the activated sludge process. Pro¢. 10th Int. Conf. 1.A.W.P.R.,Toronto. In Prog. War. Technol. Vol. 12, 1980. G. A. & Marais G. V. R. (1979) Letter to the EdiEkama tor. Water S.A. 5, 57-60. Gujer W. & Jenkins D. (1975a) The contact stabilization activated sludge process---oxygen utilization, sludge production and efficiency. Water Res. 9, 553-560. Gujer W. and Jenkins D. (1975b) A nitrification model for the contact stabilization activated sludge process. Water Res. 9, 561-566. Marais G. V. R. & Ekama G. A. (1976) The activated sludge process. Part 1--Steady state behaviour. Water S.A. 2, 163-200. Ohron D. M. & Jenkins D. (1972) The mechanism and design of the contact stabilization activated sludge process. Adv. Wat. Pollut. Res. 6, 353-362.