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Critical analysis of submerged membrane sequencing batch reactor operating conditions
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Water Research 39 (2005) 4011–4019 www.elsevier.com/locate/watres
Critical analysis of submerged membrane sequencing batch reactor operating conditions Ewan McAdama,, Simon J. Judda, Rene´ Gildemeisterb, Anja Drewsb, Matthias Kraumeb a School of Water Sciences, Cranfield University, Bedfordshire, MK43 0AL, UK Department of Chemical Engineering, Technical University of Berlin, Ackerstr. 41-76, 13355 Berlin, Germany
b
Received 19 November 2004; received in revised form 13 July 2005; accepted 19 July 2005 Available online 29 August 2005
Abstract To evaluate the Submerged Membrane Sequencing Batch Reactor process, several short-term studies were conducted to define critical flux, membrane aeration and intermittent filtration operation. Critical flux trials indicated that as mixed liquor suspended solids increased in concentration so would the propensity for membrane fouling. Consequently in order to characterise the impact of biomass concentration increase (that develops during permeate withdrawal) upon submerged microfiltration operation, two longer term studies were conducted, one with a falling hydraulic head and another with a continuous hydraulic head (as in membrane bio-reactors). Trans membrane pressure data was used to predict the maximum possible operating periods at 10 and 62 days for the falling hydraulic head and continuous hydraulic head respectively. Further analysis revealed that falling hydraulic head operation would require 21% more aeration to maintain a consistent crossflow velocity than continuous operation and would rely on pumping for full permeate withdrawal 80% earlier. This study concluded that further optimisation would be required to make this technology technically and economically viable. r 2005 Elsevier Ltd. All rights reserved. Keywords: Sequencing batch reactor; Microfiltration; Submerged; Fouling
1. Introduction Interest has developed recently in applying membrane technology to the sequencing batch reactor (SBR), the advantages of which include the ability to modify the process conditions dependent upon influent characteristics (Kang et al., 2003), operation of all cycles in one tank, improved process control (Andreottola et al., 2001) and no entrainment of oxygen from pre-denitriCorresponding author. Tel.: +44 1234 750111x3335;
fax: +44 1234 751671. E-mail address: e.j.mcadam.2003@cranfield.ac.uk (E. McAdam).
fication (as in Membrane bio-reactors (MBR)) (Krampe and Krauth, 2001). The ability to continuously modify the process implies that a reduction in operating cost (especially aeration) is possible, as well as an improvement in effluent quality especially with respect to nutrient removal (Chang et al., 2000), thus potentially improving upon the established MBR process. However the recent development of membrane coupled SBR, means that little is known about the operational effectiveness of the membrane in these conditions. With all reaction phases occurring in one reactor, investigators have been divided on where best to site permeate withdrawal. Shin and Kang (2002) observed the importance of withdrawing permeate after aerated
0043-1354/$ - see front matter r 2005 Elsevier Ltd. All rights reserved. doi:10.1016/j.watres.2005.07.028
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react to allow full biological treatment and so sited permeate withdrawal within the anoxic period. With no subsequent aeration, the membrane fouled instantly causing an instantaneous increase in trans membrane pressure (TMP) thus reducing the available operational period of the membrane and establishing a cake more cohesive than that formed in the presence of shearing (Defrance and Jaffrin, 1999b). To remedy this, Kiso et al. (2000) induced agitation at the membrane surface using a mechanical stirrer but concluded that mixing could not set appropriate flow conditions for membrane cleaning during the anoxic period. Several investigators have also sited permeate withdrawal in place of the traditional SBR phases of settle and draw (Bae et al., 2003; Chang et al., 2002). This prolongs the cycle time leading to an increase in hydraulic retention time (HRT), a decrease in chemical oxygen demand (COD) loading rate, potential endogenous respiration (Shin and Kang, 2002) and approximately doubles the aeration requirement. Accordingly more recent investigation has led to siting withdrawal within aerated react albeit with variable success (Krampe and Krauth, 2001; Shin and Kang, 2002). This paper therefore investigates the operational performance of submerged microfiltration with permeate withdrawal sited in the aerated react period as a cost saving measure and as part of a wider study to determine the feasibility of submerged membrane SBR (SMSBR) technology.
Fig. 1. Experimental reactor design.
Table 1 Ingredients for composition of synthetic greywater Ingredients of concentrate
Greywater daily amount per day (per 150 l/ per person)
Toothpaste Showergel Handsoap Domestic cleaner Oil/ lotion Tertiary effluent Shampoo Bubble bath Urea Na2CO3 K2PO4 NH3Cl
1.5 g 10 ml 0.7 g 60 ml 0.5 ml 50 ml
Bath amount per bath (per 100 l)
2. Materials and methods 2.1. Experimental setup The experimental setup consisted of a 29 l bioreactor with a submerged microfiltration plate and frame module (Maxflow, A3 GmbH) comprising 12 elements with nominal porosity 0.4 mm, manufactured from polyphenol resin and a total membrane area of 0.38 m2 (Fig. 1). Permeate was removed using a peristaltic pump (Ecoline VC, ISM 912C) sited on the permeate line. Reactor volume and TMP were measured using pressure transducers (BD Sensors GmbH, DMP 331P) located at the base of the reactor and on the permeate line respectively. All information from the pressure transducers was transferred and recorded directly on a computer. A fine bubble membrane diffuser (ENVICON GmbH) was used to supply air for both biological aeration and membrane air scour. Airflow was controlled via a needle gauge (Platon Instruments, 0–790 l h1). A motor driven stirrer was used to ensure complete mixing during the anoxic phase. Automated control of SBR phases was via Programmable Logic Controller (PLC) (Siemens, Logo 12/ 24 RC). The PLC also controlled feed mode, stirrer and aeration via electrically actuated solenoids (Buschjost GmbH). With
5 ml 25 ml 7.0 g 5.5 g 500 mg 7g
the exception of sampling no biomass was removed from the reactor, resulting in a solids retention time (SRT) of approximately 39 days. 2.2. Feed solution and analysis A synthetic feed solution was adapted from previous literature (Komvuschara, 2002; Nolde, 1999) in order to produce an analogue greywater (Table 1). The concentrate was diluted to maintain a near constant COD concentration of 200 mg l1.
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Pumping was suspended between tests, the module removed from the reactor, chemically cleaned and tested with clean water to detect for signs of permanent fouling prior to the next test being undertaken. Cleaning involved an initial 1 h circulation with 0.5 wt%. sodium dichlor isocyanurat solution through the module followed by a 12 h soaking. The module was subsequently mechanically washed with clean water followed by a throughput of clean water for 1 h prior to use. 2.4. Data analysis and testing protocol During short term testing, permeate was returned to the top of the reactor to maintain a practically constant pressure on the retentate side of the membrane thus changes in TMP were assumed to be solely due to fouling (Le-Clech et al., 2003). TMP was taken to equate to the difference between the average hydraulic head at the mid-point of the membrane module and the permeate pressure (Le-Clech et al., 2003). The permeate flux was calculated by collecting a volume of filtrate over a set time period which was then repeated three times and an average calculated. Flux was adjusted to a reference temperature of 20 1C using standard methods (Ueda et al., 1997). The reactor was allowed a minimum 24 h to recover between tests.
3. Results and discussion 3.1. Definition of operational conditions 3.1.1. Sub-critical flux The fixed flux method was adopted in this investigation as it avoids overfouling of the membrane in the initial stages of operation (Defrance and Jaffrin, 1999a). The weak form of critical flux was used and is defined as the final flux step at which TMP stabilises (Kwon and Vigneswaran, 1998). Below critical flux, TMP was found to rise slowly and then stabilise in less than 15 min, which agrees with a similar investigation conducted by Defrance and Jaffrin (1999a). The threshold point equating to critical flux was approximately
-1
2.8 g l
25
20
Flux Increments
20
15
15
10
10
Flux (l m-2 h-1)
25
30
5
5 0
2.3. Cleaning protocol
30
40
TMP (kPa)
During experimentation, analysis of the feed and permeate was conducted for COD, total nitrogen (TN) and total phosphorous (TP) using Dr Lange test solutions followed by determination via spectrophotometry (Dr Lange, ISIS 9000 MDA Photometer). Mixed liquor suspended solids (MLSS) and mixed liquor volatile suspended solids (MLVSS) concentration was determined by standard methods (APHA, 1992).
4013
3.5 g l-1 0
80
160
240 320 400 Time (mins)
480
560
640
0
Fig. 2. Comparison of flux trials conducted at two different MLSS concentrations.
13.5 l m2 h1, when MLSS concentration was stable at 3.5 g l1. A foaming incident prior to the final critical flux trial resulted in a loss of MLSS from a stable 3.5 to 2.8 g l1 inducing a shift in critical flux to 14.8 l m2 h1 (Fig. 2). It has been reported that fouling is more significant at higher MLSS concentration, and conversely that higher MLSS concentration results in less fouling under certain conditions (Kang et al., 2003; Lee et al., 2003). However, Defrance et al. (2000) conducted fixed TMP trials within biomass concentration range 2–6 g l1 noting only minor flux decline. Whilst the previous research implies an element of ambiguity, this experimental data indicates that an increase in biomass concentration within the reactor will increase the operational TMP (corresponding to a reduction in operating flux under hydraulic pressure operation). Unlike in a continuous process, it is common practice in SBR to separate fill and draw from one another (Ketchum, 1997), creating a variation in the hydraulic head by up to 50% of reactor volume dependent upon the volumetric exchange ratio (VER) adopted. Consequently as withdrawal continues and hydraulic head decreases, biomass concentration will increase. Therefore even if initially operating at sub-critical flux, concentration of the biomass may lead operation into a ‘supra-critical state’ (Defrance and Jaffrin, 1999a) due to the formation of an irreversible cohesive cake (Defrance and Jaffrin, 1999b), the ramifications of which include increased chemical cleaning frequency, greater potential of irreversible fouling and an increase in pumping costs due to the variability in available hydraulic pressure and increased TMP (or declining flux dependent upon control adopted). 3.1.2. Membrane aeration rate The reactor was designed as an internal loop airlift reactor, thus it was possible to develop a simulation model based on Chisti et al.’s (1988) work
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to mathematically determine superficial liquid velocity (ULr), and so investigate shear effects at the membrane surface, by applying the reactors dimensional parameters to Eq. (1): 2 30:5 6 2ghD ðr d Þ 7 U Lr ¼ 4 2 5 K B AAdr ð11 Þ2
(1)
d
16
0.04
14
0.035
12
0.03
10
0.025
8
0.02
6
0.015
4
0.01
2
0.005
0
0
0.1
0.2
0.3
0.4
0.5
0 0.6
Aeration Intensity (m3 m-2 s-1)
TMP (kPa)
where g is acceleration due to gravity, hD is the height of the gas–liquid dispersion, er and ed are the gas hold-up in the riser and downcomer respectively, KB is the frictional loss coefficient for the bottom connecting section, and Ar and Ad are the cross sectional areas of the riser and downcomer respectively. More recently several authors have applied the model to submerged membrane bioreactor design (Liu et al., 2000; Kishino et al., 1996; Sofia et al., 2004). Liu et al. (2000) determined a crossflow velocity from the model of 0.3 m s1 and found an equally good correlation between theoretically and experimentally derived data operating at an MLSS 4 g l1 and a flux of 5.2 l m2 h1. Sofia et al. (2004) also achieved good correlation when investigating the difference between coarse and fine bubble aeration, concluding fine bubble aeration to have the smaller error between data obtained experimentally and theoretically. Flux was set just below 13.5 l m2 h1 throughout the trial and each airflow rate subjected to an operating period of 20 h. Aeration intensities were calculated using airflow rate and effective cross sectional area of the riser (Liu et al., 2000) and were between 0.0014 and 0.036 m3 m2 s1 corresponding to superficial liquid velocities in the range of 0.15 to 0.5 m s1. As depicted in Fig. 3, an almost linear relationship between critical pressure (defined as the intercept of the plateau that occurs upon TMP stabilisation under set conditions) (Ueda et al., 1997) and crossflow velocity was observed up to 0.4 m s1, similar to the relationship noted by Defrance and Jaffrin (1999a). This change of rate of increase of TMP is thought to denote a change in fouling
Crossflow Velocity (m s-1)
Fig. 3. Effect of crossflow velocity on TMP (m) with corresponding aeration intensity (’). Operating at an MLSS of 3–3.5 g l1 and a flux of 12.2 l m2 h1.
rate, i.e. rate of increase of TMP decreases with increasing crossflow velocity (Liu et al., 2000). Beyond a ULr of 0.4 m s1 (aeration intensity 0.017 m3 m2 s1) the observed linear relationship is assumed to cease to exist and any additional aeration intensity applied had very little effect on the pressure. Several authors have postulated that this point is due to effects of the system hydrodynamics where gas hold up diminishes and/ or the onset of circulation patterns occur and is termed the point of critical aeration intensity (Liu et al., 2000; Sofia et al., 2004). Sofia et al. (2004) investigated a submerged plate and frame MBR treating municipal wastewater and attained a plateau at 0.69 m s1 corresponding to an aeration intensity of 0.017 m3 m2 s1 identical to the assumed critical aeration intensity observed in this investigation. The resultant non linear relationship observed beyond the point of critical aeration intensity implies that the cake removing efficiency of aeration does not increase proportionally with increase of air flow and that this critical value may be adopted as the operational air flow rate to avoid over supply of air (Ueda et al., 1997). It is also evident that transferring critical aeration intensity to ‘normal’ SBR operation could be difficult because, as can be seen from Eq. (1), liquid dispersion height is of significance to ULr. The simulation model was reconfigured to determine the air flow increase required to maintain a constant ULr of 0.4 m s1 whilst the hydraulic head reduces during permeate withdrawal. Fig. 4 reports a range of hydraulic height from 2.00 m to 0.38 m (the latter being the hydraulic height at the end of withdrawal) indicating that as the hydraulic head decreases, aeration intensity must be increased to maintain the same ULr. In fact based on a VER of 0.5 and the same aeration intensity, ULr will fall from 0.4 m s1 to 0.3 m s1 by the completion of withdrawal within this system. It is further evident that the initial hydraulic height adopted above the membrane will influence how much additional aeration will be required, i.e. the greater the initial hydraulic head the lesser the additional aeration
16 Volume of Air (l min-1)
4014
14 12 10 8 6 4 2 0 0.38 0.58 0.78 0.98 1.18 1.38 1.58 1.78 1.98 Height (m)
Fig. 4. Increasing air requirement in order to maintain crossflow velocity at 0.4 m s1 as the hydraulic head falls.
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requirement (this is also dependent on the relationship between draught tube height and hydraulic height). Liu et al. (2000) also identified that preliminary dimensional analysis with such a model could reduce the airflow required to achieve the same ULr. 3.1.3. Intermittent filtration The most suitable intermittent operating protocol for submerged activated sludge systems is to maintain constant superficial liquid velocity and vary applied flux or TMP thereby employing a cyclical stop/ start filtration operation (White and Lesecq, 1993). This technique relies on producing a sufficiently high shear stress to erode the top layer of cake formed at the membrane surface. However, it has been shown recently that build up of a cake cannot be reversed simply by lowering the TMP unless the fluid velocity is high enough to reach the yield stress or critical erosion stress for the particles in the cake (Kwon and Vigneswaran, 1998; Si-Hassen et al., 1996). The critical aeration intensity has been defined as 0.017 m3 m2 s1 in this investigation, and whilst it may correspond to the critical erosion stress, it does not necessarily. However, as this intensity was deemed most beneficial in previous trials, the same value was adopted for this trial. Three cycles were selected for comparison as follows: continuous mode; 10 min operation/2 min break; and 20 min operation/2 min break and were operated for approximately 20 hours per cycle (Fig. 5). The 2 min break time was selected arbitrarily but has been adopted in several previous investigations (Krampe and Krauth, 2001; Ueda et al., 1997). The recovery of TMP after each break during intermittent mode was similar to that found by Defrance and Jaffrin (1999b) where the TMP rise is initially rapid followed by a gradually rising plateau. The authors interpreted the results to indicate that when filtration stops, partial loosening of the cake occurs and that when filtration is resumed the cake is reorganised by packing, causing pressure to increase 35 30
TMP (kPa)
25 20 15 10 5
4015
(Defrance and Jaffrin, 1999b). An obvious distinction between Defrance and Jaffrin’s (1999b) work and this current investigation however can be made when comparing the slight difference in TMP reported between continuous and intermittent operation compared to the significant difference encountered between intermittent and continuous operation in this investigation. During this trial complete recovery of initial TMP was not achieved as observed during long intermittency trials by Ueda et al. (1997) and may be as a result of operating below the critical erosion stress, however the distinction shown between operating continuously and intermittently emphasises the benefit of intermittent operation. The two intermittent cycles adopted were found to be comparable yielding a minimal difference in pressure change (1.5 104 kPa min1) between the initial startup period and where operation reached a plateau (corresponding to 38 and 873 min of operation respectively). The total time used for breaks (dead operation time) during each cycle was calculated for the 10 and 20 min cycles and comprised 16.7% and 9.1% of the total cycle time respectively, thus indicating the twenty minute cycle to be more applicable simply on the basis of cost. 3.2. Long term studies 3.2.1. Membrane performance Having identified that the biomass concentration will increase during the withdrawal period, its influence on membrane performance was then investigated by comparing two testing protocols each with identical cycle times (4.5 h), HRT (9 h), VER (approximately 0.5), flux, aeration rate and intermittency (all defined previously) over a seven day operating period, the only difference being the location of fill (Table 2). Protocol 1 applied a ‘continuous head’ by replacing the withdrawn permeate with fresh feed in the reactor during aerated react. Protocol 2 applied an instant fill strategy at the start of the anoxic period followed by continuous withdrawal during aerated react (hence resulting in a biomass concentration increase by a factor of approximately 2). Aerated react was set at 3 h to ensure that time was available for permeate withdrawal (approximately 2.75 h). A 1.5 h anoxic period followed aerated react. Table 2 Cycle periods
0 0
100 200 300 400 500 600 700 800 900 1000 1100 1200
Time (mins)
Fig. 5. Comparison of two intermittent filtration strategies (10 min on/2 min off (’); 20 min on/ 2 min off (~)) versus a continuous (—) filtration strategy, operating at a flux of 13.5 l m2 h1 and a crossflow velocity 0.4 m s1.
Time (minutes) Continuous head operation Falling head operation
Aerobic
Anoxic
180 Fill/withdrawal Withdrawal
90 — Instant fill
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40
20
35 30 TMP (kPa)
TMP (kPa)
15 10 5
25 20 15 10 5
0 0 -5
20
40
60
80
100
120
140
160
0 0
Time (mins)
2000 4000 6000 8000 10000 12000 14000 16000 Time (mins)
Fig. 6. Comparison of continuous (’) and falling (~) hydraulic head operation during the 1st withdrawal cycle whilst operating at an MLSS 2.5 g l1, flux 12.2 l m2 h1 and crossflow velocity 0.4 m s1.
Fig. 7. Comparison of TMP cyclic peaks for both continuous (~) and falling (’) hydraulic head operation, whilst operating at an MLSS 2.5 g l1, flux 12.2 l m2 h1 and crossflow velocity 0.4 m s1.
In both cases where filtration was restarted, TMP recovered to that recorded prior to the intermittent break within the first few minutes of operation (Fig. 6). During continuous operation beyond this point TMP appeared to stabilise after approximately 4 min until the following intermittent break, with only a negligible increase throughout the withdrawal cycle. This is qualitatively similar to the work of White and Lesecq (1993) who interpreted the effect as a dissipation of a reversible concentration polarisation fouling layer where near total flux could be recovered. During falling hydraulic head operation, once TMP has recovered, it continues to climb without a period of stabilisation. In MBR systems producing low filtration fluxes a rapid increase in membrane hydraulic resistance from the beginning of filtration is anticipated due to the accumulation of bioparticles on the membrane surface however, this increase is usually followed by stabilisation (Tardieu et al., 1998). The instability of the falling hydraulic head system indicates that flux operation may be supra-critical from the beginning of the cycle (Defrance and Jaffrin, 1999a). An approximate linear relationship can be observed in TMP increase with time over the withdrawal cycle which this author directly attributes to the linear increase in biomass concentration as permeate withdrawal continues and hydraulic head reduces. It is generally recognised that fouling of the membrane is most critical at the start of the filtration run, or when the change in pressure or flux is most rapid (Gander et al., 2000), thus by operating within falling hydraulic head conditions, fouling will occur throughout each complete cycle resulting in the significant dP/dt relationship observed. This rapid development of fouling at the membrane surface is also thought to lead to a greater risk of irreversible fouling (Liu et al., 2000). Furthermore the rapid traverse of TMP across each complete withdrawal cycle implies that pumping
intervention will be required from the earliest point of operation. Analysis of the relationship dK/dt (change in permeability with time) further supports this assumption with linear decline in permeability corresponding to 5.1 103 and 2.8 104 (l m2 h1) kPa1 min1 for falling hydraulic head and continuous hydraulic head operation respectively. To simplify the long term data, the final TMP peak from each withdrawal cycle over the full operation period has been reported in Fig. 7. Both data sets have certain discrepancies which are explained in Table 3. The subsequent effect of the final fault has been reported by Ueda et al. (1997) where an extended period of aeration without withdrawal has returned TMP to near its initial value upon recommencement of filtration. In order to complete the data set, operation was allowed to continue for a further 5000 min. Examination of the full data indicates a linear increase of TMP with time as a consequence of progressive fouling and agrees with a trend observed during other investigations (White and Lesecq, 1993; Tardieu et al., 1998). The steady TMP rise is due to the gradual loss of ‘active’ membrane area due to fouling in each cycle. This loss leads to increased local flux even though the averaged value is constant (Ognier et al., 2004). It is this phenomenon that drives the TMP rise over the series of cycles. In order to evaluate fouling in the absence of sludge viscosity data, the rate of change relationship dP/dt (gradient from Fig. 7) was used. The filtration period was periodical due to operating an SBR cycle however, the linear relationship was defined using the full cycle time which was therefore substituted for dt. The TMP during the continuous hydraulic head test increased very slowly from 12.5 to 16.5 kPa over the full period of operation (dP/ dt ¼ 4 104 kPa min1). Conversely during the falling hydraulic head test a significant increase in TMP from 15.0 to 34.4 kPa was recorded over the full operating
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Table 3 Indication of data faults for Fig. 7 Study
Visual observation
Problem
Result
Continuous Continuous Falling
‘dip’ 2972–3512 min ‘peak’ at 7121 min ‘break’ 1829–4647 min & 4647–6058 min
Solenoid fault Solenoid fault PLC fault
Reactor overfilled Extended withdrawal Complete aeration, no withdrawal
period (dP/dt ¼ 1.7 103 kPa min1). The operational time (top.) between startup and chemical regeneration was estimated using Eq. (2): top: ¼
TMPf TMPi , dP=dt
(2)
where TMPi is the initial pressure, TMPf is the membrane manufacturers recommended maximum operating pressure (40 kPa) and dP/ dt is the gradient of the line of best fit (Fig. 7). The operating time between chemical cleans were calculated to be 10 and 62 days for the falling hydraulic head and continuous hydraulic head systems respectively. The continuous hydraulic head operating time is similar in duration to those reported for MBR (Kiso et al., 2000), however longer MBR operating periods have been reported upto 8 months between cleans (Sofia et al., 2004). The linear increase in TMP corresponds to a linear decrease in permeability, less dramatic than that observed in a single cycle, recording 1.3 104 and 4.7 105 (l m2 h1) kPa1 min1 for falling and continuous hydraulic head operation respectively, but nevertheless indicating a palpable distinction between the two operating conditions. The falling hydraulic head permeability data also compares poorly to data from an investigation by Bae et al. (2003) who operated an SMSBR in similar hydraulic conditions but at an operating flux that was only 37% of the critical flux (compared to 90% adopted in this investigation). This compounds the importance of selecting an appropriate operating flux however, it should be noted that Bae et al. (2003) were able to select such a low operating flux due to an abundance of membrane surface area. Pumping requirements were estimated by comparing the TMP data from each falling hydraulic head cycle with the available hydraulic pressure based on an initial hydraulic head of 2 m (assumed as a standard operating hydraulic height) (Gander et al., 2000). From withdrawal cycle 1, pumping assistance is required after approximately 90 min of operation corresponding to a hydraulic head of approximately 1.25 m. Full reliance on pumping for permeate removal occurs at withdrawal cycle 16, corresponding to 5.3 days of operation. Conversely in the continuous system, no pump intervention is required until 24.4 days of operation has lapsed, beyond this point full reliance is placed on pumping for permeate withdrawal. Thus the falling head
operation will require full pumping intervention 80% earlier than the continuous method. Finally, additional aeration requirement to maintain the ULr during falling head operation was calculated by estimating the area under the graph in Fig. 4 (again assuming a standard height of 2 m) identifying that 21% more air is required than in continuous operation. 3.2.2. Biological performance Establishment of biological performance was peripheral to analysis of membrane performance and in such a short-term study is difficult to characterise fully. Whilst good COD removal was achieved (91% and 92% for falling and continuous hydraulic head respectively), the low total nitrogen removal efficiency (50% and 54% for falling and continuous head tests respectively) can be attributed to the fill mode selected (Ketchum, 1997) and siting permeate withdrawal within aerated react rather than after a further anoxic period as with conventional SBR. Shin and Kang (2002) attributed a decrease or low nitrogen removal efficiency to an extended withdrawal period in SMSBR as it extended aerated react (due to declining flux), leading to long HRT and lower F: M resulting in the decay of biomass activity by endogenous respiration of sludge. A resultant reduction in the withdrawal period (by chemical restoration of the membrane) restored nitrogen removal efficiencies (Shin and Kang, 2002). As with Shin and Kang’s (2002) investigation, MLSS concentration initially declined but was then followed by a slow recovery (to approximately 2.5 g l1) corresponding to a period of endogenous respiration whilst the biomass adapted to the feed conditions. It is anticipated that this was because aerated react was designed for permeate withdrawal rather than biological treatment. An investigation by Krampe and Krauth (2001) agrees reporting nitrogen removal efficiencies up to 80% in SMSBR by operating at a lower VER (0.1–0.3) thus being able to abbreviate or extend the react period and site permeate withdrawal in the latter part of aerated react, withdrawing permeate once biological treatment has neared completion.
4. Conclusions From this study it is apparent that fouling of the membrane occurs when operating in ‘normal’ SBR
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conditions even when sub-critical flux has been adopted and membrane aeration conditions optimised. The TMP changes dramatically as the hydraulic head lowers throughout the withdrawal period forcing operation into a theoretically supra-critical state even during the initial stages of filtration. The ensuing increase in TMP (corresponding to a linear decline in permeability) during the long-term study indicated only a short period of operation is possible before reaching the maximum recommended TMP, thus comparing poorly to continuous hydraulic head membrane technology such as MBR. Additionally pumping and aeration costs attributed to falling hydraulic head operation coupled with a high chemical cleaning frequency reinforce that this current SMSBR operational strategy is both technically and economically unfeasible. Biological performance was fairly consistent throughout the operating period achieving good COD removal but poor total nitrogen removal. The poor nitrogen removal can be accredited to removing permeate in aerated react and the extension of aerated react to allow for permeate withdrawal thus inducing endogenous respiration. Imposing a lower volumetric exchange ratio can reintroduce the flexibility into SMSBR. A reduction in the volume withdrawn will lower the biomass concentration with which the membrane must finally operate in and can then be combined with selection of a lower operating flux, extending the operative period between chemical interventions. Additional pumping and aeration costs can be correlated to the VER, thus a lowering in VER will directly translate to a reduction in pumping and aeration costs. The reduction in VER and subsequent reduction in length of permeate withdrawal implies that aerated react can be varied in length in accordance with the treatment requirements and permeate withdrawal can be preferentially sited toward the end of aerated react which should improve total nitrogen removal.
Acknowledgements The authors would like to thank those at the Department of Chemical Engineering, Technical University of Berlin for their help in developing the experimental rig outlined in this paper and for the support of A3 Gmbh in donating the membrane module used in the design.
References Andreottola, G., Foladori, P., Ragazzi, M., 2001. On-line control of a SBR system for nitrogen removal from industrial wastewater. Water Sci. Technol. 43 (3), 93–100.
APHA, 1992. Standard methods for water and wastewater Examination, Seventeenth ed. American Public Health Association, Washington, DC. Bae, T.-H., Han, S.-S., Tak, T.-M., 2003. Membrane sequencing batch reactor system for the treatment of dairy industry wastewater. Process Biochem. 39 (2), 221–231. Chang, C.-N., Ho, C.-F., Chen, Y.-R., Chao, A., Ma, Y.-S., 2002. Reclamation of wastewater by sequencing batch membrane bioreactor (SBMBR). J. Chin. Inst. Environ. Eng. 12 (4), 295–305. Chang, H.N., Ra, K.M., Byung, G.P., Seong-Jin, L., Dong, W.C., Woo, G.L., Seok, L.S., Yong, H.A., 2000. Simulation of sequential batch reactor (SBR) operation for simultaneous removal of nitrogen and phosphorous. Bioprocess Eng 23 (5), 513–521. Chisti, M.Y., Halard, B., Moo-Young, M., 1988. Liquid circulation in airlift reactors. Chem. Eng. Sci. 43 (3), 451–457. Defrance, L., Jaffrin, M.Y., 1999a. Comparison between filtrations at fixed transmembrane pressure and fixed permeate flux: application to a membrane bioreactor used for wastewater treatment. J. Membr. Sci. 152 (2), 203–210. Defrance, L., Jaffrin, M.Y., 1999b. Reversibility of fouling formed in activated sludge filtration. J. Membr. Sci. 157 (1), 73–84. Defrance, L., Jaffrin, M.Y., Gupta, B., Payllier, P., Geaugey, V., 2000. Contribution of various constituents of activated sludge to membrane bioreactor fouling. Bioresource Technol. 73 (2), 105–112. Gander, M., Jefferson, B., Judd, S.J., 2000. Aerobic MBRs for domestic wastewater treatment: a review with cost considerations. Sep. Purif. Technol. 18 (2), 119–130. Kang, I.S., Lee, C.-H., Kim, K.-J., 2003. Characteristics of microfiltration membrane in a membrane coupled SBR. Water Res. 37 (5), 1192–1197. Ketchum, L.H., 1997. Design and physical features of SBRs. Water Sci. Technol. 35 (1), 11–18. Kishino, H., Ishida, H., Iwabu, H., Nakano, I., 1996. Domestic wastewater reuse using a submerged membrane bioreactor. Desalination 106 (1–3), 115–119. Kiso, Y., Jung, Y.-J., Ichinari, T., Park, M., Kitao, T., Nishimura, K., Min, K.-S., 2000. Wastewater treatment performance of a filtration bio-reactor equipped with a mesh as a filter material. Water Res. 34 (17), 4143–4150. Komvuschara, K., 2002. UV disinfection in comparison with other small scale technologies for the treatment of domestic greywater. Ph.D. Thesis, University of Hertfordshire, United Kingdom. Krampe, J., Krauth, K., 2001. Sequencing batch reactor with submerged hollow fibre membranes for the biomass separation. Water Sci. Technol. 43 (3), 195–199. Kwon, D.Y., Vigneswaran, S., 1998. Influence of particle size and surface charge on critical flux of crossflow microfiltration. Water Sci. Technol. 38 (4–5), 481–488. Le-Clech, P., Jefferson, B., Chang, I.S., Judd, S.J., 2003. Critical Flux Determination by the Flux Step Method in a Submerged MBR. J. Membr. Sci. 227 (1–2), 81–93. Lee, W., Kang, S., Shin, H., 2003. Sludge characteristics and their contribution to microfiltration in submerged membrane bioreactors. J. Membr. Sci. 216 (1–2), 217–227.
ARTICLE IN PRESS E. McAdam et al. / Water Research 39 (2005) 4011–4019 Liu, R., Huang, X., Wang, C., Chen, L., Qian, Y., 2000. Study on hydraulic characteristics in a submerged membrane bioreactor process. Process Biochem. 36 (3), 249–254. Nolde, E., 1999. Greywater reuse systems for toilet flushing in multi storey buildings — over ten years experience in Berlin. Urban Water 1 (4), 275–284. Ognier, S., Wisniewski, C., Grasmick, A., 2004. Membrane bioreactor fouling in sub-critical filtration conditions: a local critical flux concept. J. Membr. Sci. 229 (1–2), 171–177. Shin, H., Kang, S., 2002. Performance and membrane fouling in a pilot scale SBR process coupled with membrane. Asia Environmental Technology. Water Sci. Technol. 47 (1), 139–144. Si-Hassen, D., Ould-Dris, A., Jaffrin, M.Y., Benkhala, Y.K., 1996. Optimisation of an intermittent cross-flow filtration
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process of mineral suspensions. J. Membr. Sci. 118 (2), 185–198. Sofia, A., Ng, W.J., Ong, S.L., 2004. Engineering design approaches for minimum fouling in submerged MBR. Desalination 160 (1), 67–74. Tardieu, E., Grasmick, A., Geaugey, V., Manem, J., 1998. Hydrodynamic control of bioparticle deposition in a MBR applied to wastewater treatment. J. Membr. Sci. 147 (1), 1–12. Ueda, T., Hata, K., Kikuoka, Y., Seino, O., 1997. Effects of aeration on suction pressure in a submerged membrane bioreactor. Water Res. 31 (3), 489–494. White, D.A., Lesecq, P., 1993. Optimisation if intermittently operated microfiltration processes. Chem. Eng. J. 52 (2), 73–77.
Water Research 39 (2005) 4011–4019 www.elsevier.com/locate/watres
Critical analysis of submerged membrane sequencing batch reactor operating conditions Ewan McAdama,, Simon J. Judda, Rene´ Gildemeisterb, Anja Drewsb, Matthias Kraumeb a School of Water Sciences, Cranfield University, Bedfordshire, MK43 0AL, UK Department of Chemical Engineering, Technical University of Berlin, Ackerstr. 41-76, 13355 Berlin, Germany
b
Received 19 November 2004; received in revised form 13 July 2005; accepted 19 July 2005 Available online 29 August 2005
Abstract To evaluate the Submerged Membrane Sequencing Batch Reactor process, several short-term studies were conducted to define critical flux, membrane aeration and intermittent filtration operation. Critical flux trials indicated that as mixed liquor suspended solids increased in concentration so would the propensity for membrane fouling. Consequently in order to characterise the impact of biomass concentration increase (that develops during permeate withdrawal) upon submerged microfiltration operation, two longer term studies were conducted, one with a falling hydraulic head and another with a continuous hydraulic head (as in membrane bio-reactors). Trans membrane pressure data was used to predict the maximum possible operating periods at 10 and 62 days for the falling hydraulic head and continuous hydraulic head respectively. Further analysis revealed that falling hydraulic head operation would require 21% more aeration to maintain a consistent crossflow velocity than continuous operation and would rely on pumping for full permeate withdrawal 80% earlier. This study concluded that further optimisation would be required to make this technology technically and economically viable. r 2005 Elsevier Ltd. All rights reserved. Keywords: Sequencing batch reactor; Microfiltration; Submerged; Fouling
1. Introduction Interest has developed recently in applying membrane technology to the sequencing batch reactor (SBR), the advantages of which include the ability to modify the process conditions dependent upon influent characteristics (Kang et al., 2003), operation of all cycles in one tank, improved process control (Andreottola et al., 2001) and no entrainment of oxygen from pre-denitriCorresponding author. Tel.: +44 1234 750111x3335;
fax: +44 1234 751671. E-mail address: e.j.mcadam.2003@cranfield.ac.uk (E. McAdam).
fication (as in Membrane bio-reactors (MBR)) (Krampe and Krauth, 2001). The ability to continuously modify the process implies that a reduction in operating cost (especially aeration) is possible, as well as an improvement in effluent quality especially with respect to nutrient removal (Chang et al., 2000), thus potentially improving upon the established MBR process. However the recent development of membrane coupled SBR, means that little is known about the operational effectiveness of the membrane in these conditions. With all reaction phases occurring in one reactor, investigators have been divided on where best to site permeate withdrawal. Shin and Kang (2002) observed the importance of withdrawing permeate after aerated
0043-1354/$ - see front matter r 2005 Elsevier Ltd. All rights reserved. doi:10.1016/j.watres.2005.07.028
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react to allow full biological treatment and so sited permeate withdrawal within the anoxic period. With no subsequent aeration, the membrane fouled instantly causing an instantaneous increase in trans membrane pressure (TMP) thus reducing the available operational period of the membrane and establishing a cake more cohesive than that formed in the presence of shearing (Defrance and Jaffrin, 1999b). To remedy this, Kiso et al. (2000) induced agitation at the membrane surface using a mechanical stirrer but concluded that mixing could not set appropriate flow conditions for membrane cleaning during the anoxic period. Several investigators have also sited permeate withdrawal in place of the traditional SBR phases of settle and draw (Bae et al., 2003; Chang et al., 2002). This prolongs the cycle time leading to an increase in hydraulic retention time (HRT), a decrease in chemical oxygen demand (COD) loading rate, potential endogenous respiration (Shin and Kang, 2002) and approximately doubles the aeration requirement. Accordingly more recent investigation has led to siting withdrawal within aerated react albeit with variable success (Krampe and Krauth, 2001; Shin and Kang, 2002). This paper therefore investigates the operational performance of submerged microfiltration with permeate withdrawal sited in the aerated react period as a cost saving measure and as part of a wider study to determine the feasibility of submerged membrane SBR (SMSBR) technology.
Fig. 1. Experimental reactor design.
Table 1 Ingredients for composition of synthetic greywater Ingredients of concentrate
Greywater daily amount per day (per 150 l/ per person)
Toothpaste Showergel Handsoap Domestic cleaner Oil/ lotion Tertiary effluent Shampoo Bubble bath Urea Na2CO3 K2PO4 NH3Cl
1.5 g 10 ml 0.7 g 60 ml 0.5 ml 50 ml
Bath amount per bath (per 100 l)
2. Materials and methods 2.1. Experimental setup The experimental setup consisted of a 29 l bioreactor with a submerged microfiltration plate and frame module (Maxflow, A3 GmbH) comprising 12 elements with nominal porosity 0.4 mm, manufactured from polyphenol resin and a total membrane area of 0.38 m2 (Fig. 1). Permeate was removed using a peristaltic pump (Ecoline VC, ISM 912C) sited on the permeate line. Reactor volume and TMP were measured using pressure transducers (BD Sensors GmbH, DMP 331P) located at the base of the reactor and on the permeate line respectively. All information from the pressure transducers was transferred and recorded directly on a computer. A fine bubble membrane diffuser (ENVICON GmbH) was used to supply air for both biological aeration and membrane air scour. Airflow was controlled via a needle gauge (Platon Instruments, 0–790 l h1). A motor driven stirrer was used to ensure complete mixing during the anoxic phase. Automated control of SBR phases was via Programmable Logic Controller (PLC) (Siemens, Logo 12/ 24 RC). The PLC also controlled feed mode, stirrer and aeration via electrically actuated solenoids (Buschjost GmbH). With
5 ml 25 ml 7.0 g 5.5 g 500 mg 7g
the exception of sampling no biomass was removed from the reactor, resulting in a solids retention time (SRT) of approximately 39 days. 2.2. Feed solution and analysis A synthetic feed solution was adapted from previous literature (Komvuschara, 2002; Nolde, 1999) in order to produce an analogue greywater (Table 1). The concentrate was diluted to maintain a near constant COD concentration of 200 mg l1.
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35
Pumping was suspended between tests, the module removed from the reactor, chemically cleaned and tested with clean water to detect for signs of permanent fouling prior to the next test being undertaken. Cleaning involved an initial 1 h circulation with 0.5 wt%. sodium dichlor isocyanurat solution through the module followed by a 12 h soaking. The module was subsequently mechanically washed with clean water followed by a throughput of clean water for 1 h prior to use. 2.4. Data analysis and testing protocol During short term testing, permeate was returned to the top of the reactor to maintain a practically constant pressure on the retentate side of the membrane thus changes in TMP were assumed to be solely due to fouling (Le-Clech et al., 2003). TMP was taken to equate to the difference between the average hydraulic head at the mid-point of the membrane module and the permeate pressure (Le-Clech et al., 2003). The permeate flux was calculated by collecting a volume of filtrate over a set time period which was then repeated three times and an average calculated. Flux was adjusted to a reference temperature of 20 1C using standard methods (Ueda et al., 1997). The reactor was allowed a minimum 24 h to recover between tests.
3. Results and discussion 3.1. Definition of operational conditions 3.1.1. Sub-critical flux The fixed flux method was adopted in this investigation as it avoids overfouling of the membrane in the initial stages of operation (Defrance and Jaffrin, 1999a). The weak form of critical flux was used and is defined as the final flux step at which TMP stabilises (Kwon and Vigneswaran, 1998). Below critical flux, TMP was found to rise slowly and then stabilise in less than 15 min, which agrees with a similar investigation conducted by Defrance and Jaffrin (1999a). The threshold point equating to critical flux was approximately
-1
2.8 g l
25
20
Flux Increments
20
15
15
10
10
Flux (l m-2 h-1)
25
30
5
5 0
2.3. Cleaning protocol
30
40
TMP (kPa)
During experimentation, analysis of the feed and permeate was conducted for COD, total nitrogen (TN) and total phosphorous (TP) using Dr Lange test solutions followed by determination via spectrophotometry (Dr Lange, ISIS 9000 MDA Photometer). Mixed liquor suspended solids (MLSS) and mixed liquor volatile suspended solids (MLVSS) concentration was determined by standard methods (APHA, 1992).
4013
3.5 g l-1 0
80
160
240 320 400 Time (mins)
480
560
640
0
Fig. 2. Comparison of flux trials conducted at two different MLSS concentrations.
13.5 l m2 h1, when MLSS concentration was stable at 3.5 g l1. A foaming incident prior to the final critical flux trial resulted in a loss of MLSS from a stable 3.5 to 2.8 g l1 inducing a shift in critical flux to 14.8 l m2 h1 (Fig. 2). It has been reported that fouling is more significant at higher MLSS concentration, and conversely that higher MLSS concentration results in less fouling under certain conditions (Kang et al., 2003; Lee et al., 2003). However, Defrance et al. (2000) conducted fixed TMP trials within biomass concentration range 2–6 g l1 noting only minor flux decline. Whilst the previous research implies an element of ambiguity, this experimental data indicates that an increase in biomass concentration within the reactor will increase the operational TMP (corresponding to a reduction in operating flux under hydraulic pressure operation). Unlike in a continuous process, it is common practice in SBR to separate fill and draw from one another (Ketchum, 1997), creating a variation in the hydraulic head by up to 50% of reactor volume dependent upon the volumetric exchange ratio (VER) adopted. Consequently as withdrawal continues and hydraulic head decreases, biomass concentration will increase. Therefore even if initially operating at sub-critical flux, concentration of the biomass may lead operation into a ‘supra-critical state’ (Defrance and Jaffrin, 1999a) due to the formation of an irreversible cohesive cake (Defrance and Jaffrin, 1999b), the ramifications of which include increased chemical cleaning frequency, greater potential of irreversible fouling and an increase in pumping costs due to the variability in available hydraulic pressure and increased TMP (or declining flux dependent upon control adopted). 3.1.2. Membrane aeration rate The reactor was designed as an internal loop airlift reactor, thus it was possible to develop a simulation model based on Chisti et al.’s (1988) work
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to mathematically determine superficial liquid velocity (ULr), and so investigate shear effects at the membrane surface, by applying the reactors dimensional parameters to Eq. (1): 2 30:5 6 2ghD ðr d Þ 7 U Lr ¼ 4 2 5 K B AAdr ð11 Þ2
(1)
d
16
0.04
14
0.035
12
0.03
10
0.025
8
0.02
6
0.015
4
0.01
2
0.005
0
0
0.1
0.2
0.3
0.4
0.5
0 0.6
Aeration Intensity (m3 m-2 s-1)
TMP (kPa)
where g is acceleration due to gravity, hD is the height of the gas–liquid dispersion, er and ed are the gas hold-up in the riser and downcomer respectively, KB is the frictional loss coefficient for the bottom connecting section, and Ar and Ad are the cross sectional areas of the riser and downcomer respectively. More recently several authors have applied the model to submerged membrane bioreactor design (Liu et al., 2000; Kishino et al., 1996; Sofia et al., 2004). Liu et al. (2000) determined a crossflow velocity from the model of 0.3 m s1 and found an equally good correlation between theoretically and experimentally derived data operating at an MLSS 4 g l1 and a flux of 5.2 l m2 h1. Sofia et al. (2004) also achieved good correlation when investigating the difference between coarse and fine bubble aeration, concluding fine bubble aeration to have the smaller error between data obtained experimentally and theoretically. Flux was set just below 13.5 l m2 h1 throughout the trial and each airflow rate subjected to an operating period of 20 h. Aeration intensities were calculated using airflow rate and effective cross sectional area of the riser (Liu et al., 2000) and were between 0.0014 and 0.036 m3 m2 s1 corresponding to superficial liquid velocities in the range of 0.15 to 0.5 m s1. As depicted in Fig. 3, an almost linear relationship between critical pressure (defined as the intercept of the plateau that occurs upon TMP stabilisation under set conditions) (Ueda et al., 1997) and crossflow velocity was observed up to 0.4 m s1, similar to the relationship noted by Defrance and Jaffrin (1999a). This change of rate of increase of TMP is thought to denote a change in fouling
Crossflow Velocity (m s-1)
Fig. 3. Effect of crossflow velocity on TMP (m) with corresponding aeration intensity (’). Operating at an MLSS of 3–3.5 g l1 and a flux of 12.2 l m2 h1.
rate, i.e. rate of increase of TMP decreases with increasing crossflow velocity (Liu et al., 2000). Beyond a ULr of 0.4 m s1 (aeration intensity 0.017 m3 m2 s1) the observed linear relationship is assumed to cease to exist and any additional aeration intensity applied had very little effect on the pressure. Several authors have postulated that this point is due to effects of the system hydrodynamics where gas hold up diminishes and/ or the onset of circulation patterns occur and is termed the point of critical aeration intensity (Liu et al., 2000; Sofia et al., 2004). Sofia et al. (2004) investigated a submerged plate and frame MBR treating municipal wastewater and attained a plateau at 0.69 m s1 corresponding to an aeration intensity of 0.017 m3 m2 s1 identical to the assumed critical aeration intensity observed in this investigation. The resultant non linear relationship observed beyond the point of critical aeration intensity implies that the cake removing efficiency of aeration does not increase proportionally with increase of air flow and that this critical value may be adopted as the operational air flow rate to avoid over supply of air (Ueda et al., 1997). It is also evident that transferring critical aeration intensity to ‘normal’ SBR operation could be difficult because, as can be seen from Eq. (1), liquid dispersion height is of significance to ULr. The simulation model was reconfigured to determine the air flow increase required to maintain a constant ULr of 0.4 m s1 whilst the hydraulic head reduces during permeate withdrawal. Fig. 4 reports a range of hydraulic height from 2.00 m to 0.38 m (the latter being the hydraulic height at the end of withdrawal) indicating that as the hydraulic head decreases, aeration intensity must be increased to maintain the same ULr. In fact based on a VER of 0.5 and the same aeration intensity, ULr will fall from 0.4 m s1 to 0.3 m s1 by the completion of withdrawal within this system. It is further evident that the initial hydraulic height adopted above the membrane will influence how much additional aeration will be required, i.e. the greater the initial hydraulic head the lesser the additional aeration
16 Volume of Air (l min-1)
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14 12 10 8 6 4 2 0 0.38 0.58 0.78 0.98 1.18 1.38 1.58 1.78 1.98 Height (m)
Fig. 4. Increasing air requirement in order to maintain crossflow velocity at 0.4 m s1 as the hydraulic head falls.
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requirement (this is also dependent on the relationship between draught tube height and hydraulic height). Liu et al. (2000) also identified that preliminary dimensional analysis with such a model could reduce the airflow required to achieve the same ULr. 3.1.3. Intermittent filtration The most suitable intermittent operating protocol for submerged activated sludge systems is to maintain constant superficial liquid velocity and vary applied flux or TMP thereby employing a cyclical stop/ start filtration operation (White and Lesecq, 1993). This technique relies on producing a sufficiently high shear stress to erode the top layer of cake formed at the membrane surface. However, it has been shown recently that build up of a cake cannot be reversed simply by lowering the TMP unless the fluid velocity is high enough to reach the yield stress or critical erosion stress for the particles in the cake (Kwon and Vigneswaran, 1998; Si-Hassen et al., 1996). The critical aeration intensity has been defined as 0.017 m3 m2 s1 in this investigation, and whilst it may correspond to the critical erosion stress, it does not necessarily. However, as this intensity was deemed most beneficial in previous trials, the same value was adopted for this trial. Three cycles were selected for comparison as follows: continuous mode; 10 min operation/2 min break; and 20 min operation/2 min break and were operated for approximately 20 hours per cycle (Fig. 5). The 2 min break time was selected arbitrarily but has been adopted in several previous investigations (Krampe and Krauth, 2001; Ueda et al., 1997). The recovery of TMP after each break during intermittent mode was similar to that found by Defrance and Jaffrin (1999b) where the TMP rise is initially rapid followed by a gradually rising plateau. The authors interpreted the results to indicate that when filtration stops, partial loosening of the cake occurs and that when filtration is resumed the cake is reorganised by packing, causing pressure to increase 35 30
TMP (kPa)
25 20 15 10 5
4015
(Defrance and Jaffrin, 1999b). An obvious distinction between Defrance and Jaffrin’s (1999b) work and this current investigation however can be made when comparing the slight difference in TMP reported between continuous and intermittent operation compared to the significant difference encountered between intermittent and continuous operation in this investigation. During this trial complete recovery of initial TMP was not achieved as observed during long intermittency trials by Ueda et al. (1997) and may be as a result of operating below the critical erosion stress, however the distinction shown between operating continuously and intermittently emphasises the benefit of intermittent operation. The two intermittent cycles adopted were found to be comparable yielding a minimal difference in pressure change (1.5 104 kPa min1) between the initial startup period and where operation reached a plateau (corresponding to 38 and 873 min of operation respectively). The total time used for breaks (dead operation time) during each cycle was calculated for the 10 and 20 min cycles and comprised 16.7% and 9.1% of the total cycle time respectively, thus indicating the twenty minute cycle to be more applicable simply on the basis of cost. 3.2. Long term studies 3.2.1. Membrane performance Having identified that the biomass concentration will increase during the withdrawal period, its influence on membrane performance was then investigated by comparing two testing protocols each with identical cycle times (4.5 h), HRT (9 h), VER (approximately 0.5), flux, aeration rate and intermittency (all defined previously) over a seven day operating period, the only difference being the location of fill (Table 2). Protocol 1 applied a ‘continuous head’ by replacing the withdrawn permeate with fresh feed in the reactor during aerated react. Protocol 2 applied an instant fill strategy at the start of the anoxic period followed by continuous withdrawal during aerated react (hence resulting in a biomass concentration increase by a factor of approximately 2). Aerated react was set at 3 h to ensure that time was available for permeate withdrawal (approximately 2.75 h). A 1.5 h anoxic period followed aerated react. Table 2 Cycle periods
0 0
100 200 300 400 500 600 700 800 900 1000 1100 1200
Time (mins)
Fig. 5. Comparison of two intermittent filtration strategies (10 min on/2 min off (’); 20 min on/ 2 min off (~)) versus a continuous (—) filtration strategy, operating at a flux of 13.5 l m2 h1 and a crossflow velocity 0.4 m s1.
Time (minutes) Continuous head operation Falling head operation
Aerobic
Anoxic
180 Fill/withdrawal Withdrawal
90 — Instant fill
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40
20
35 30 TMP (kPa)
TMP (kPa)
15 10 5
25 20 15 10 5
0 0 -5
20
40
60
80
100
120
140
160
0 0
Time (mins)
2000 4000 6000 8000 10000 12000 14000 16000 Time (mins)
Fig. 6. Comparison of continuous (’) and falling (~) hydraulic head operation during the 1st withdrawal cycle whilst operating at an MLSS 2.5 g l1, flux 12.2 l m2 h1 and crossflow velocity 0.4 m s1.
Fig. 7. Comparison of TMP cyclic peaks for both continuous (~) and falling (’) hydraulic head operation, whilst operating at an MLSS 2.5 g l1, flux 12.2 l m2 h1 and crossflow velocity 0.4 m s1.
In both cases where filtration was restarted, TMP recovered to that recorded prior to the intermittent break within the first few minutes of operation (Fig. 6). During continuous operation beyond this point TMP appeared to stabilise after approximately 4 min until the following intermittent break, with only a negligible increase throughout the withdrawal cycle. This is qualitatively similar to the work of White and Lesecq (1993) who interpreted the effect as a dissipation of a reversible concentration polarisation fouling layer where near total flux could be recovered. During falling hydraulic head operation, once TMP has recovered, it continues to climb without a period of stabilisation. In MBR systems producing low filtration fluxes a rapid increase in membrane hydraulic resistance from the beginning of filtration is anticipated due to the accumulation of bioparticles on the membrane surface however, this increase is usually followed by stabilisation (Tardieu et al., 1998). The instability of the falling hydraulic head system indicates that flux operation may be supra-critical from the beginning of the cycle (Defrance and Jaffrin, 1999a). An approximate linear relationship can be observed in TMP increase with time over the withdrawal cycle which this author directly attributes to the linear increase in biomass concentration as permeate withdrawal continues and hydraulic head reduces. It is generally recognised that fouling of the membrane is most critical at the start of the filtration run, or when the change in pressure or flux is most rapid (Gander et al., 2000), thus by operating within falling hydraulic head conditions, fouling will occur throughout each complete cycle resulting in the significant dP/dt relationship observed. This rapid development of fouling at the membrane surface is also thought to lead to a greater risk of irreversible fouling (Liu et al., 2000). Furthermore the rapid traverse of TMP across each complete withdrawal cycle implies that pumping
intervention will be required from the earliest point of operation. Analysis of the relationship dK/dt (change in permeability with time) further supports this assumption with linear decline in permeability corresponding to 5.1 103 and 2.8 104 (l m2 h1) kPa1 min1 for falling hydraulic head and continuous hydraulic head operation respectively. To simplify the long term data, the final TMP peak from each withdrawal cycle over the full operation period has been reported in Fig. 7. Both data sets have certain discrepancies which are explained in Table 3. The subsequent effect of the final fault has been reported by Ueda et al. (1997) where an extended period of aeration without withdrawal has returned TMP to near its initial value upon recommencement of filtration. In order to complete the data set, operation was allowed to continue for a further 5000 min. Examination of the full data indicates a linear increase of TMP with time as a consequence of progressive fouling and agrees with a trend observed during other investigations (White and Lesecq, 1993; Tardieu et al., 1998). The steady TMP rise is due to the gradual loss of ‘active’ membrane area due to fouling in each cycle. This loss leads to increased local flux even though the averaged value is constant (Ognier et al., 2004). It is this phenomenon that drives the TMP rise over the series of cycles. In order to evaluate fouling in the absence of sludge viscosity data, the rate of change relationship dP/dt (gradient from Fig. 7) was used. The filtration period was periodical due to operating an SBR cycle however, the linear relationship was defined using the full cycle time which was therefore substituted for dt. The TMP during the continuous hydraulic head test increased very slowly from 12.5 to 16.5 kPa over the full period of operation (dP/ dt ¼ 4 104 kPa min1). Conversely during the falling hydraulic head test a significant increase in TMP from 15.0 to 34.4 kPa was recorded over the full operating
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Table 3 Indication of data faults for Fig. 7 Study
Visual observation
Problem
Result
Continuous Continuous Falling
‘dip’ 2972–3512 min ‘peak’ at 7121 min ‘break’ 1829–4647 min & 4647–6058 min
Solenoid fault Solenoid fault PLC fault
Reactor overfilled Extended withdrawal Complete aeration, no withdrawal
period (dP/dt ¼ 1.7 103 kPa min1). The operational time (top.) between startup and chemical regeneration was estimated using Eq. (2): top: ¼
TMPf TMPi , dP=dt
(2)
where TMPi is the initial pressure, TMPf is the membrane manufacturers recommended maximum operating pressure (40 kPa) and dP/ dt is the gradient of the line of best fit (Fig. 7). The operating time between chemical cleans were calculated to be 10 and 62 days for the falling hydraulic head and continuous hydraulic head systems respectively. The continuous hydraulic head operating time is similar in duration to those reported for MBR (Kiso et al., 2000), however longer MBR operating periods have been reported upto 8 months between cleans (Sofia et al., 2004). The linear increase in TMP corresponds to a linear decrease in permeability, less dramatic than that observed in a single cycle, recording 1.3 104 and 4.7 105 (l m2 h1) kPa1 min1 for falling and continuous hydraulic head operation respectively, but nevertheless indicating a palpable distinction between the two operating conditions. The falling hydraulic head permeability data also compares poorly to data from an investigation by Bae et al. (2003) who operated an SMSBR in similar hydraulic conditions but at an operating flux that was only 37% of the critical flux (compared to 90% adopted in this investigation). This compounds the importance of selecting an appropriate operating flux however, it should be noted that Bae et al. (2003) were able to select such a low operating flux due to an abundance of membrane surface area. Pumping requirements were estimated by comparing the TMP data from each falling hydraulic head cycle with the available hydraulic pressure based on an initial hydraulic head of 2 m (assumed as a standard operating hydraulic height) (Gander et al., 2000). From withdrawal cycle 1, pumping assistance is required after approximately 90 min of operation corresponding to a hydraulic head of approximately 1.25 m. Full reliance on pumping for permeate removal occurs at withdrawal cycle 16, corresponding to 5.3 days of operation. Conversely in the continuous system, no pump intervention is required until 24.4 days of operation has lapsed, beyond this point full reliance is placed on pumping for permeate withdrawal. Thus the falling head
operation will require full pumping intervention 80% earlier than the continuous method. Finally, additional aeration requirement to maintain the ULr during falling head operation was calculated by estimating the area under the graph in Fig. 4 (again assuming a standard height of 2 m) identifying that 21% more air is required than in continuous operation. 3.2.2. Biological performance Establishment of biological performance was peripheral to analysis of membrane performance and in such a short-term study is difficult to characterise fully. Whilst good COD removal was achieved (91% and 92% for falling and continuous hydraulic head respectively), the low total nitrogen removal efficiency (50% and 54% for falling and continuous head tests respectively) can be attributed to the fill mode selected (Ketchum, 1997) and siting permeate withdrawal within aerated react rather than after a further anoxic period as with conventional SBR. Shin and Kang (2002) attributed a decrease or low nitrogen removal efficiency to an extended withdrawal period in SMSBR as it extended aerated react (due to declining flux), leading to long HRT and lower F: M resulting in the decay of biomass activity by endogenous respiration of sludge. A resultant reduction in the withdrawal period (by chemical restoration of the membrane) restored nitrogen removal efficiencies (Shin and Kang, 2002). As with Shin and Kang’s (2002) investigation, MLSS concentration initially declined but was then followed by a slow recovery (to approximately 2.5 g l1) corresponding to a period of endogenous respiration whilst the biomass adapted to the feed conditions. It is anticipated that this was because aerated react was designed for permeate withdrawal rather than biological treatment. An investigation by Krampe and Krauth (2001) agrees reporting nitrogen removal efficiencies up to 80% in SMSBR by operating at a lower VER (0.1–0.3) thus being able to abbreviate or extend the react period and site permeate withdrawal in the latter part of aerated react, withdrawing permeate once biological treatment has neared completion.
4. Conclusions From this study it is apparent that fouling of the membrane occurs when operating in ‘normal’ SBR
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conditions even when sub-critical flux has been adopted and membrane aeration conditions optimised. The TMP changes dramatically as the hydraulic head lowers throughout the withdrawal period forcing operation into a theoretically supra-critical state even during the initial stages of filtration. The ensuing increase in TMP (corresponding to a linear decline in permeability) during the long-term study indicated only a short period of operation is possible before reaching the maximum recommended TMP, thus comparing poorly to continuous hydraulic head membrane technology such as MBR. Additionally pumping and aeration costs attributed to falling hydraulic head operation coupled with a high chemical cleaning frequency reinforce that this current SMSBR operational strategy is both technically and economically unfeasible. Biological performance was fairly consistent throughout the operating period achieving good COD removal but poor total nitrogen removal. The poor nitrogen removal can be accredited to removing permeate in aerated react and the extension of aerated react to allow for permeate withdrawal thus inducing endogenous respiration. Imposing a lower volumetric exchange ratio can reintroduce the flexibility into SMSBR. A reduction in the volume withdrawn will lower the biomass concentration with which the membrane must finally operate in and can then be combined with selection of a lower operating flux, extending the operative period between chemical interventions. Additional pumping and aeration costs can be correlated to the VER, thus a lowering in VER will directly translate to a reduction in pumping and aeration costs. The reduction in VER and subsequent reduction in length of permeate withdrawal implies that aerated react can be varied in length in accordance with the treatment requirements and permeate withdrawal can be preferentially sited toward the end of aerated react which should improve total nitrogen removal.
Acknowledgements The authors would like to thank those at the Department of Chemical Engineering, Technical University of Berlin for their help in developing the experimental rig outlined in this paper and for the support of A3 Gmbh in donating the membrane module used in the design.
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