POWER CONSUMPTION AND MASS TRANSFER IN AN UNBAFFLED STIRRED TANK FOR AUTOTHERMAL THERMOPHILIC DIGESTION OF SLUDGE

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1 POWER CONSUMPTION AND MASS TRANSFER IN AN UNBAFFLED STIRRED TANK FOR AUTOTHERMAL THERMOPHILIC DIGESTION OF SLUDGE HASSAN Raouf, LOUBIERE Karine 1, LEGRAND Jack GEPEA, UMR 6144, CNRS, Université de Nantes/ENITIAA/EMN CRTT, Boulevard de l'université, BP 406, Saint-Nazaire Cedex, France Abstract New strategies for the disposal and reuse of sewage sludge are developing to face their growing amount while satisfying the new European directives. The most widely adopted process for stabilizing sludge remains mesophilic anaerobic digestions. As enabling to respect stringent specifications regarding pathogens, a migration to higher temperatures is requested. In this context, an interesting option is the Autothermal Thermophilic Aerobic Digestion in which the biological reaction produces itself most of the energy required to achieve thermophilic conditions. This is possible because of the high load of organic matter to be degraded, but the counterpart is a great need of dissolved oxygen. The latter constraint makes necessary tailored investigations for optimizing mixing and aeration in these complex reactors where transfer phenomena and biological conversion kinetics are strongly coupled. This paper reports a preliminary study carried out to characterize a benchscale ATAD reactor in terms of power consumption and gas-liquid mass transfer. This digester was made of an unbaffled vessel equipped with four nozzle spargers and a non-standard impeller (anchor-type paddle). Under ungassed conditions, the power number curve was established and the constants that characterize the laminar and turbulent regimes determined. A model describing the variation of N P over the whole range of Re was proposed. Under gassed conditions, the gas-liquid regimes were visualized and the flow maps defined. The Relative Power Demands were also quantified as a function of gas flow numbers. Overall volumetric mass transfer coefficients were measured in water, with paying a special attention on the effect of mechanical agitation. Keywords: ATAD reactor; gas-liquid flow; power consumption; mass transfer coefficient. 1- Introduction The treatment and disposal of urban and industrial wastewater sludge is an expensive and environmentally sensitive world-wide problem. Increasing environmental and legislative constraints for sludge disposal have intensified the interest to develop sustainable and economical methods for sludge reduction. In Europe, the Sewage Sludge Directive 86/278/EEC regulates the uses and properties of stabilised sludge for being either recycled or disposed. The Auto-thermal Thermophilic Aerobic Digestion (ATAD) is classified as a high temperature aerobic digestion process technology. It is a very suitable alternative solution for implementing thermophilic conditions in sludge treatment. Contrary to the usual ones, this process offers the advantage to provide simultaneously stabilization and disinfection thanks to the temperatures involved (typically between 40 C and 70 C). If the amount of organic matter and the oxygen supply are sufficient, no additional heat is required, leading thus to a self-heating operation. The mass of sludge is commonly reduced to 35-50% and of Class A (Layden, 2007). In a chemical engineering point of view, an ATAD reactor is a very complex process to model, as all the following phenomena interact with others: multiphase system (air bubbles dispersed into a biological liquid phase containing solid particles), presence of biological reactions and coupling between heat and mass transfers. Moreover, sludge are non-newtonian fluids: they present a shear-thinning behaviour with, possibly, a yield point, thixotropicity and/or viscoelasticity (Seyssiecq et al., 2003). These rheological 1 Corresponding author. New affiliations: Université de Toulouse, Laboratoire de Génie Chimique (LGC), CNRS, INP, UPS. New address : LGC, Site Basso-Cambo BP 1301, 5 rue Paulin Talabot, Toulouse France. Karine.Loubiere@ensiacet.fr.

2 properties depend on many parameters (total solid content, biological flocs, temperature, and particle size) and are not reproducible (season, origin, pre-treatments, operating conditions). They strongly influence hydrodynamics (mixing), heat transfer (operating temperature), gas-liquid mass transfer (oxygen supply for microorganisms), and biological degradation. With regard to this complexity, tailored investigations are required for optimising mixing and aeration and for better understanding the phenomena involved and their coupling. A special attention should be paid on the relationship between the gas-liquid mass transfer performances and the biological degradation (efficiency, thermophilic temperature, autothermal conditions). For that, a lab-scale ATAD reactor has been initially designed at the Rovira I Virgili University (Tarragona, Spain) aiming to studying the biological degradation kinetics of sludge. It was made of an unbaffled vessel equipped with four nozzles and a non-standard impeller (anchor-type paddle). In keeping with a larger scientific context (European program REMOVALS), the paper is focused on some preliminary investigations carried out to characterise this original agitation system in terms of power consumption and aeration performance. 2- Material and Methods 2.1 Experimental set-up The lab-scale reactor consists of a 10 L unbaffled vessel, with a cylindrical and curved-bottomed shape, made of PMMA and having a diameter (T) of 21.2 cm and a total height (H T ) of 31.6 cm. It is equipped with a square double jacket ( cm 2 ). Air is introduced inside the tank through four nozzles (equally dispatched with an inner spacing of 12 cm). These spargers are made of stainless steel fibre mesh and have a circular shape (1.4 cm in diameter). Air flow rate is regulated by four manometers (Samson ) and measured by using four volumetric flowmeters (Brooks R2-25-D), enabling thus the same air flow rate in each nozzle to be imposed. The total air flow rate is ranged from 0.33 to 3.0 NL/min. Agitation is provided with an impeller fixed on a shaft which rotation speed (N) is regulated by an electrical motor (Ikavisc MR-D1 Messrührer, Janke & Kunkel, Ika ). The impeller has a non-standard shape, looking like an anchor-type paddle, with a diameter (D) of 90 mm, a height of 50 mm, a width of 20 mm and a thickness of 3.0 mm. The vertical distance between the vessel bottom and the lower horizontal plan of the impeller, called the impeller clearance (C), varies between T/3, T/4 and T/7. Water, aqueous solutions of glycerine (30, 67, 85, 93 and 100% percent in weight) and viscous oil (Emkarox ) are used as liquid phases. Viscosities are measured by a rotational rheometer (PAAR Physica MCR500) and the density by a hydrometer. The physical properties of the liquids are reported in Table 1. Table 1. Physical properties of liquid phases at 20 C. Solution Viscosity (µ) (mpa s) Density (ρ) (kg/m 3 ) Water Glycerine (30%)(67%)(85%)(93%) (2.5)(17.7)(82.3)(253)(367) (1072)(1171)(1223)(1245) pure Emkarox oil Power consumption measurements The power consumption is deduced from measurements of the torque induced by the agitator in rotation. The torquemeter used (IKAVISC MR-D1) has a sensitivity of 0.01 N.cm. In practice, the value of the torque was variable due to the mechanical friction in the bearings; that is why, for more accuracy, the maximum and the minimum values of the torques were systematically recorded for each experiment. The power consumed by the impeller (P expressed in W) is calculated using the torques measured when the impeller is submerged in the liquid (M) and when the impeller rotates in air (M O ): ( M ) P torque = 2 πn M (1) Ο Under gassed conditions, the total power consumption is deduced from: P = P + ρ gu V (2) torque L G L

3 where U G = Q G /(πt 2 /4) is the gas superficial velocity. These results are commonly presented by correlating the dimensionless power number (N p = P/ρN 3 D 5 ) with the Reynolds number (R e = ρnd 2 /µ). Whatever the agitation system, the laminar regime occurs for Re < 10 and is characterised by: N P. Re = K p For turbulent flows (i.e. when Re > Re t, with Re t 10 4 for radial impellers, Re t 10 5 for axial impellers; Nagata, 1975), the power number becomes constant and is independent of liquid viscosity: N p = N pt (4) 2.4 Gas-liquid mass transfer measurements (3) Figure 1. Schematic diagram of the experimental set-up: 1: Valve; 2: Manometer; 3: Volumetric flowmeter; 4: Spargers; 5: Impeller; 6: O 2 and temperature probes; 7: Cover; 8: Electrical motor; 9: Torquemeter; 10: Transmitter; 11: Computer; 12: Double jacket The standard dynamic gassing out method is used to measure the volumetric gas-liquid mass transfer coefficient k L a. The dissolved oxygen concentrations are recorded as a function of time by using two probes (InPro-6050, Mettler Toledo) and an acquisition card. The LabVIEW software is used for data acquisition. The probes positions are represented in Figure 1 (vertically at 4 cm and 18 cm above the bottom of the vessel, horizontally at 1.5 cm of the walls). Assuming that the liquid phase is well mixed, the mass balance in the dissolved oxygen concentration is written: dco 2 = k * La C O C 2 O dt 2 k L a is deduced from the slope of ln(c * O2 C O2 ) versus time and corrected by taking into account the time constants of probes ( 16 s). As T = 20 ± 3 C, a temperature correction is applied (Bewtra et al., 1970): kl a = k ( T ) La T (6) The mean k L a in the tank is an average of the values measured for each probe (three tests per condition). The experiments are performed in water (H L = 21.2 cm, C = 5.3 cm), without and with mechanical agitation (100, 200, and 300 rpm) and for air flow rates Q G varying between 0.33 and 3 NL/min. 3- Results and discussion 3.1 Power consumption In Figure 2-a, the progressive decrease of N P for increasing Re is reported. By fitting the experimental data with the values predicted by either Eq. (3) (if Re < 10) or Eq. (4) (if Re > 10 4 ), the constants K p and N pt are firstly determined. As commonly used, the following correlations are also tested to model the variation of the power numbers over the whole range of Reynolds numbers: K A n + N Re P pt N = + (Nagata et al., 1957) (7) p Re Cn + Re p KP Re0.66 N p = + B. Re Re0.66 (Nagata, 1975) (8) From Table 2, it appears that the liquid height and the impeller level have no major effect on the model s constants, even if a slight increase and decrease of both laminar and turbulent constants can be distinguished when increasing H L and C respectively. More important is that the present laminar constants (5)

4 are significantly smaller than the ones obtained for usual impellers: K P = 35 for Rushton turbine, for anchor and for propellers (Xuereb et al., 2006). The turbulent constants ( ) have the same order of magnitude than anchors or marine propellers, but they remain very small when compared to the radial impellers commonly used for dispersing gas in agitated tanks (Roustan, 2005). For N = 370 rpm (used by our Spanish partner for biological study), the associated tip speed (πdn) and power consumed are close to 1.7 m/s and 60 W/m 3 respectively. According to literature (Paul et al., 2004), these values are very low, suggesting thus that the dispersion of bubble plume by the present impeller would not be sufficient for aerating efficiently the liquid medium. To confirm this, the aeration performances inside the lab-scale reactor are investigated below N P H = 21.2 cm, C = 7 cm H = 21.2 cm, C = 3.2 cm H = 15 cm, C = 7 cm H = 15 cm, C = 3.2 cm Re (a) RPD = P G /P 100 rpm 150 rpm 200 rpm 250 rpm 300 rpm Fl 0.0 G Figure 2. (a) Power curve N P (Re) without gas, (b) Relative Power Demand, RPD, versus Gas Flow number. The well-known equation for impeller power is often modified for gas-liquid systems to give: P = N ( RPD). N 3. D5 p ρ (9) where RPD is the relative power demand or gassing factor (P G /P) which depends on the blade shape. Figure 2-b illustrates, for the present system, the decrease of RPD with increased dimensionless gas rate (or gas flow number Fl G = Q G /ND 3 ). It can be noticed that RPD falls a little at high rotation speeds (RDP > 0.9), whereas a significant decrease (from 0.8 to 0.4) is observed for N=100 rpm. This should be linked to the fact that, for the smallest speeds, the bubbles are regrouped around the shaft, and tend to form some gas cavities behind the blades (loading regime). Working at N > 300 rpm is thus recommended for minimising the power loss under gassed conditions. Table 2. Constants associated with the N P (Re) modelling. (b) H L (cm) d (cm) Nagata et al. (1957) [Eq. (7)] Nagata (1975) [Eq. (8)] K P N Pt A n C n ε (%) B p ε*(%) * The deviations between measurements and predicted values are calculated as ε=1/n.σ (x mod -x exp )/x exp * Overall volumetric mass transfer coefficients Firstly, gas flow pattern is visualised to qualitatively appreciate the degree of recirculation and back mixing of the gas phase. Depending on the rotation speed and gas flow rate, three dispersion regimes are usually distinguished: (i) flooding, where the impeller is overwhelmed by gas and the mixing very poor (ii) loading, where the impeller disperses the gas through the upper part of the tank; (iii) complete dispersion, where the gas bubbles are distributed throughout the tank. The gas-liquid flow maps are defined by using the Froude number (Fr = N 2 D/g) and the gas Flow number (Fl G = Q G /ND 3 ). It can be seen from Figure 3-c that, for most of experiments, the loading regime is observed; some zones free of bubbles appear then at the bottom of the tank and close to the wall (Figure 3-a). For four experiments (at N = 300 rpm), the complete dispersion of the gas phase is approximately reached; nevertheless, the areas free of bubbles outside of the spargers, even restricted, still exist and the vortex is quite big. When

5 compared to literature, the present flow numbers are rather small (Fl G <0.04). It is mainly due to the small gas flow rates tested, which have been imposed by some constraints own to the operation of the ATAD reactor in presence of sludge (minimisation of evaporation and foam formation). The present flow map is compared with several relationships developed in water in tanks with straight baffles and Rushton turbine (Figure 3-c): Nienow et al. (1977) defined the maximum gas-holding for full recirculation, Warmoeskerken et al. (1981) reported the minimum (Fr) for gas dispersion and Nienow et al. (1985) determined the maximum gas flow rate before flooding the impeller. Even a slight shift exits, this confirms that experiments are mainly in the loading regime 1,00 Fr Nienow et al.(1977) Fl G=13.Fr 2 (D/T) 5 Nienow et al.(1985) FlG=30.Fr.(D/T) 3.5 0,10 Warmoeskerken et al. (1981) : Fr=0.045 Loading Complete dispersion 0,01 0,0001 0,0010 0,0100 Fl G 0,1000 1,0000 (a) (b) (c) Figure 3. Photographs of the dispersion regime: (a) Loading regime (N = 200 rpm and Q G = 2.3 NL/min), (b) Complete dispersion (N = 300 rpm and Q G = 3 NL/min), (c) Gas-liquid flow map. To illustrate the effect of agitation on the gas liquid mass transfer, the variations of the mean volumetric mass transfer coefficients with the air flow rates are shown in Figure 4-a (water). As expected, k L a increases with increased gas flow rates and rotation speeds. This is mainly because, as the impeller speed increases, gas dispersion and bubble break-up are favoured, enhancing thus the surface area available for mass transfer. These results can be modelled according to the relationships usually available in literature, where the variation of k L a is related, by means of a power law, either to N and Q G, or to Fr and Fl G, or to (P G /ρv) and U G. As shown in Figure 4-b, the average deviations between experimental and predicted values are below 10% whatever the conditions. When compared to the well-known correlation of Van t Riet (1979), the exponent on U G is found identical (0.51 against 0.5), the exponent on P G /ρv is here twice smaller (0.17 against 0.4) and the constant significantly higher (0.35 against 0.026) k La (s -1 ) rpm 200 rpm 300 rpm without agitation Predicted k La (s -1 ) +10 % -10 % Correlation (N, Qg) Correlation (Fr,Flg) Correlation (Pg, Ug) k L a = 0.41.N 0.59 Q G 0.51 k L a/n = Fr 0.05 Fl G 0.51 k L a = 0.35.(P G /ρv) 0.17 U G 0.51 Q G (NL/min) (a) Experimental k La (s -1 ) (b) Figure 4 (a) Gas-liquid mass transfer coefficient (mean value) versus air flow rate. (b) Experimental k L a compared with predicted values by usual correlations In order to test the axial homogeneity of oxygen mass transfer in the tank, the following criterion (h) is used (Cabaret et al, 2008): 1 h = k Latop k Labottom / k La (10) mean n n In this equation, n represents the number of experiments, in our study n = 15 (average over all the Q G ). It is clear that h represents the mean deviation between k L a at the top and the bottom of the tank. For N =

6 100 rpm, the criterion h is equal to 4.6 %, for N = 200 rpm to 2.7 % and for N = 300 rpm 2.8 %. Rising the rotation speeds enables thus to better homogenise the mass transfer. At last, it is of importance to outline that the present values of k L a are insufficient. Indeed, the associated oxygen transfer rates are ranged between 0.04 and 0.2 kgo 2 /m 3 /h, which would correspond (for the most favourable conditions) to m 3 air/m 3 /h. The latter values are significantly low when compared to the 4 m 3 air/m 3 /h theoretically required for ATAD process (EPA/625/10-90/007, 1990). It is evident that the viscous and non-newtonian properties of sludge will accentuate this tendency, i.e. will lead to worse aeration performances than the present ones in water. 4- Conclusion The objectives were to characterize a lab-scale ATAD reactor in terms of power consumption and aeration performances. The curve relating the power number to the Reynolds number was determined and modelled for the original agitation system involved (anchor-type paddle). The laminar and turbulent constants were found equal to 10 and respectively. They were very low when compared to the impellers usually used for dispersing gas. The Relative Power Demand was also quantified as a function of gas flow numbers, as well as the gas-liquid flow maps. The increase of impeller rotation speed allowed both the volumetric mass transfer coefficient and the oxygen homogenization to be improved. All these findings indicated that the aeration capacities of the present device were not sufficient (even in water) to face the biological needs when the autothermal digestion of sludge operates. That is why the agitation system is at present changed: a more standard gas-liquid system is chosen (four baffles, ring sparger, concave blade turbine). In future works, the effect of rheology (i.e. the shear thinning behaviour characteristic of sludge) on mass transfer performances will be studied in the new configuration. Acknowledgements The development of this work was made possible by the financial support of the Egyptian Ministry of High Education. This work being a part of the European project REMOVALS, the authors would like to acknowledge our partners for the Roviri I Virgili Universitat (Tarragona, Spain), Prof. C. Bengoa, Dr. E. Torrens. The authors are also especially grateful for the technical support provided by J-C Jouin. References Bewtra J.K., W.R. Nicholas, and L.B. Polkowski, 1970, Effect of temperature on oxygen transfer in water. Water Res. 4(2): 115. Cabaret F., L. Fradette, and P.A. Tanguy, 2008, Gas liquid mass transfer in unbaffled dual-impeller mixers. Chem. Eng. Sci. 63: EPA/625/10-90/007, Autothermal Thermophilic Aerobic Digestión of Municipal Wastewater Sludge. Layden, N.M., 2007, An evaluation of autothermal thermophilic aerobic digestion (ATAD) of municipal sludge in Ireland, Journal of Environ. Eng. Sci., 6: Nagata S., 1975, Mixing Principles and applications, A Halsted press book Kodansha LTD, John Wiley Nagata et al (1957). Nagata S., K. Yamamoto, and T. Yokoyama, 1957, Memo, Eng. Tokyo Univ. Japan 19:274. Nienow, A.W., D.J. Wisdom, and J.C. Middleton, 1977, The effect of scale and geometry on flooding, recirculation and power in gassed stirred vessels. Proceedings of the 2nd European Conference on Mixing, Cambridge Nienow, A.W., Warmoeskerken, M.M.C.G., J.M. Smith, and M. Konno, 1985, On the flooding/loading transition and the complete dispersal condition in aerated vessels agitated by a Rushton-turbine. Proceedings of the 5 th European Conference on Mixing, Wurzburg, West Germany, June 10-12, Paul E.L., V.A. Atiemo-Obeng, S.M. Kresta, 2004, Handbook of industrial mixing: science and practise. J. Wiley & sons inc. publication. Roustan M., 2005, Agitation. Mélange. Caractéristiques des mobiles d agitation, Techniques de l Ingénieur Form. J Form.J Seyssiecq I., A. Tolofoudyé, H. Desplanches and Y. Gaston-Bonhomme, 2003, Viscoelastic liquids in stirred vessels- Part I: Power consumption in unaerated vessels, Chemical Engineering Technology, 26 (11): Van t Riet K., 1979, Review of measuring methods and results in non-viscous gas-liquid mass transfer in stirred vessels, Ind. Eng. Chem. Process Des Dev, 18(3), Warmoeskerken, M.M.C.G., J. Feijen, and J.M. Smith, 1981, Institution of Chemical Engineers Symposium Series, 64, J1-J14. Xuereb C., M. Poux, and J. Bertrand, 2006, Agitation et mélange Aspects fondamentaux et applications industrielles, L Usine Nouvelle Dunod.