Supercritical Water Oxidation of Oily Wastes at Pilot Plant: Simulation

Ind. Eng. Chem. Res. 2011, 50, 775–784

775

Supercritical Water Oxidation of Oily Wastes at Pilot Plant: Simulation for Energy Recovery Francisco Jimenez-Espadafor,† Juan R. Portela,*,‡ Violeta Vadillo,‡ Jezabel Sańchez-Oneto,‡ Jose´ A. Becerra Villanueva,† Miguel Torres Garcıá,† and Enrique J. Martıńez de la Ossa‡ Departamento de Ingenierıá Energe´tica y Mecańica de Fluidos, Escuela Tećnica Superior de Ingenieros, UniVersidad de SeVilla, Camino de los descubrimientos s/n 41092, SeVilla, Spain, and Departamento de Ingenierıá Quı´mica y Tecnologıá de Alimentos, Facultad de Ciencias, UniVersidad de Ca´diz, 11510 Puerto Real (Ca´diz), Spain

The destruction of industrial wastewaters by supercritical water oxidation (SCWO) has been studied intensively in the last two decades due to the powerful and promising advantages of this technology. However, the SCWO process is not yet commercially established due to several drawbacks that limit its application as a general treatment, process costs being one of those limitations. In an effort to enhance the viability of SCWO as a commercial process, a study was performed in a pilot plant (25 kg/h) used to treat industrial oily wastes by SCWO, and a simulation was carried out to evaluate the viability of energy production on an industrial scale. The SCWO pilot plant effluent is good for producing hot water or steam by recovering heat of waste organics. Both alternatives are evaluated for a SCWO industrial plant design with 1000 kg/h, with it being possible to recover a maximum of 118 kW, that is, 71% of the energy content of the wastewater. 1. Introduction The supercritical water oxidation (SCWO) process is based on the oxidation of wastewaters in supercritical water (temperature >374 °C and pressure >22 MPa), an excellent reaction medium due to its special properties, which include low densities and viscosities, high diffusivities, low dielectric constant, and high miscibility with organics and oxygen.1 Based on these properties, some advantages of the SCWO process are the single one reaction phase (no mass transfer limitations), very high reaction rates, complete oxidation (removal efficiencies >99.99%), and nonharmful products (NO, NOx, SOx, etc.). The process advantages and the application of this approach to numerous wastewaters have been extensively described in the literature.2-9 However, the SCWO process is not yet commercially established due to three main drawbacks that limit its application as a general treatment: corrosion due to the severe operating conditions in the presence of an oxidant, for example, chlorides, etc.,10,11 salt precipitation due to the very low solubility of inorganics in supercritical water,12,13 and process costs14,15 due to the expensive materials and reactor configurations required to solve the first two drawbacks. The commercialization of SCWO as an alternative to conventional methods for wastewater treatment will only be achieved when the main technical and economic problems are solved. Numerous authors have proposed different technical solutions to minimize corrosion,16-18 to handle salts,19-21 and to optimize the energy balance of the process.22-24 Because oxidation reactions in SCWO are highly exothermic, once an energetically self-sufficient process has been achieved, the main goal today is the recovery of energy surplus from the destruction of a waste. An economical analysis of the SCWO process is hard to perform due to the lack of real data about investment and operational costs related to industrial scale plants. Several cost evaluations * To whom correspondence should be addressed. E-mail: [email protected]. † Universidad de Sevilla. ‡ Universidad de Ca´diz.

can be found in the literature, but all of these are based on estimations.25,26 Moreover, operational costs for SCWO are strongly influenced by the energy recovery in the process, the throughput, and the heating value of the waste. Several authors have also studied the SCWO process for generating power. Bermejo et al.27 presented a preliminary theoretical study on the SCWO of coal, which showed higher efficiencies (4-8%) than the processes in conventional coal power plants under the same steam conditions. That work relied on the possibility of separating solid particles from the process stream at 650 °C and 30 MPa, a situation that is not possible nowadays for technical reasons. Marias et al.28 attempted to solve this problem by studying the energy recovery from SCWO of biomass by using an auxiliary fluid, but they concluded that the production of electricity using a steam turbine cannot be carried out effectively. Svanstro¨m et al.29 studied the energy recovery from sewage sludges by SCWO, showing that sludge (at 10 wt % dry solids) can be oxidized with virtually complete recovery of the sludge heating value as hot water or high-pressure steam. In the work described here, the viability of energy recovery from a SCWO process was evaluated. The first part of the study concerns the results obtained in a pilot plant (25 kg/h) used to treat industrial oily wastes by SCWO. Cutting oil wastes were chosen as a suitable wastewater for SCWO process, because they are a serious environmental problem, they are almost saltfree, and they do not contain chlorides, meaning that reactor plugging and corrosion are not technical problems. Besides, our research group has studied the SCWO treatment30 of this type of waste, and all relevant kinetic parameters required for the simulation are available. In the second part of the study, a complete simulation was carried out to fit the experimental results obtained in the pilot plant and to evaluate the energy production in a SCWO industrial plant design with a capacity of 1000 kg/h. The results obtained in a pilot plant (25 kg/h) used to treat industrial oily wastes by SCWO are described along with a simulation carried out to evaluate the viability of energy production from the process on an industrial scale. It was demonstrated that power generation based on Rankine cycle is

10.1021/ie101166j  2011 American Chemical Society Published on Web 12/17/2010

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Ind. Eng. Chem. Res., Vol. 50, No. 2, 2011

Figure 1. Schematic diagram of the supercritical water oxidation pilot plant. Numbers 1-8 refer to key points used for experimental analysis and simulation.

not appropriate for this pilot plant configuration. More appropriate methods for energy recovery with a low temperature and a high fluid pressure, such as water heating and steam generation, are proposed for a typical SCWO industrial plant design with a capacity of 1000 kg/h. 2. Experimental Section A schematic diagram of the SCWO pilot plant is shown in Figure 1. The wastewater is pressurized by a high pressure liquid pump, preheated, and introduced into the tubular reactor, which consists of a stainless steel (AISI 316 L) tube of 12.32 and 19.05 mm internal and external diameters, respectively. The reactor tube is divided into three horizontal sections of 2920, 2960, and 3000 mm, with two short vertical tubes that are 69.3 and 69.9 mm long. The air stream is pressurized by a high pressure compressor and separately preheated. The oxidant and the organic compounds are mixed at high temperature (around 400 °C), and the oxidation reactions take place at a high rate, releasing a significant amount of heat. To minimize the loss of heat produced by the oxidation reactions, the continuous flow reactor is thermally insulated, and it was possible to obtain temperatures up to 550 °C. At the reactor exit, the pilot plant is fitted with a coaxial counter-current heat exchanger that is used to preheat the liquid feed by using the excess energy of the reactor effluent. After this first heat exchanger, the effluent crosses a second coaxial counter-current heat exchanger, which is used to preheat the air feed. Both heat exchangers are insulated so that this system can operate autothermally and heat from the wastewater treatment can be utilized. The first heat exchanger is a tubular countercurrent heat exchanger and is formed by two concentric tubes made of stainless steel (AISI 316 L). Hot fluid from the reactor flows through the inner pipe, and the feed stream from the high pressure pump flows through the outer pipe. The exterior pipe

has the same dimensions as the reactor tube, and the interior pipe has external and internal diameters of 9.53 and 5.5 mm, respectively. The external and internal tube lengths are 15.6 and 11.5 m, respectively (see configuration in Figure 1). The whole heat exchanger is covered with thermal insulation material (39.0 mm thickness) of the same characteristics as the reactor insulation (see section 4.3). The air heat exchanger has the same configuration as the water heat exchanger, although the tube length is different: 5.26 and 6.95 m for the external and internal pipes, respectively. Hot fluid flows through the interior pipe, and cold air from the compressor flows through the outer pipe. The excess energy released in the plant between points 7 and 8 is used in the simulation to estimate the viability of energy generation from a SCWO process on an industrial scale. In the pilot plant, a cooler (refrigerated by water) is used to ensure that the effluent temperature of the system is below 50 °C before depressurization in a back pressure regulator. Finally, the product stream is separated into liquid and vapor phases in a gas-liquid separator. Biocut (Houghton Ibe´rica S.A.) is the commercial name of the semisynthetic cutting fluid studied in this work, having a COD of 2.264 ( 0.041 (g O2/g concentrated cutting fluid). The C, H, N, and S contents were determined by elemental analysis (wt % dry basis), and it is assumed that the remainder is oxygen. The values obtained were C ) 70.1 ( 0.4, H ) 16.5 ( 2.6, N ) 0.26 ( 0.05, S ) 0.36 ( 0.04, and O ) 12.74. On the basis of these results and neglecting the small contribution of N and S, the molecular formula of the cutting oil can be represented as C6H17O. The chemical oxygen demand (COD) content of liquid samples was monitored. COD was analyzed by a closed reflux colorimetric method (5220D) according to the standard method for water and wastewater analysis.31

Ind. Eng. Chem. Res., Vol. 50, No. 2, 2011 Table 1. Summary of Oxidation Experiments Conducted in the UCA Pilot Plant Facility run 1

run 2

run 3

wastewater flow (kg/h) 10.44 10.62 10.43 air flow (kg/h) 77.6 67.8 56.6 oxygen excess (%)a 853 469 264 inlet reactor 430 430 430 temperature (°C) feed COD (g O2/L) 12.20 19.07 28.77 COD conversion (%) 80.3 80.8 84.9 effluent temp before 160 154 152 final cooling (°C) a

run 4

run 5

run 6

10.62 16.53 16.89 110.08 112.08 180.54 287 188 289 430 430 430 50.55 50.55 51.71 95.5 97.7 98.0 195 209 219

Oxygen excess ) (oxygen supplied/stoichiometric oxygen)*100.

777

the fluid composition to change from reactor input to output. The thermodynamical and transport properties are modified, and these changes have to be taken into account in the thermodynamic model of the plant. As shown in the Experimental Section, the organic compound can be defined by the general formula C6H17O. In this plant, oxygen comes from ambient air, and therefore reactants and oxidation products include N2. Other gases from air were not considered and neither was air humidity. On the basis of these considerations, the chemical reaction under thermodynamic equilibrium is given by eq 1: 4C6H17O + 39(O2 + 3.76N2) f 24CO2 + 34H2O + 146.64N2

(1)

Figure 2. Temperature profiles for different feed concentrations of Biocut at 25 MPa with an initial temperature of 430 °C.

3. Results All the experiments were carried out under supercritical conditions, at a constant pressure of 25 MPa, using compressed air as oxidant and maintaining an oxygen excess of more than 100% (relative to the stoichiometic oxygen required for complete oxidation of organics present in the feed), and different temperatures ranging between 400 and 500 °C. Because the cutting fluids are a mixture of several compounds, the efficiency of the oxidation process was followed in terms of the reduction in chemical oxygen demand (COD). Experimental results for COD disappearance at different temperatures are shown in Table 1. As can be seen, it is possible to apply hydrothermal oxidation successfully under supercritical conditions for the treatment of cutting fluid wastes, with more than 98% COD removal achieved. The temperature profiles in the reactor for the six different runs are shown in Figure 2. Despite the reactor being thermally insulated, an adiabatic behavior could not be considered, and a significant heat loss occurred. In this way, for those runs with low feed concentrations (runs 1 and 2), the heat released during oxidation is not sufficient to increase the temperature of the reaction medium, and the temperature profile decreases due to the heat loss. However, for those runs with higher feed concentrations (runs 3-6), despite the heat loss, the heat released during oxidation produces a marked increase in the temperature profile, leading to higher reaction rates and better COD conversion. After an increase in the temperature in the initial tracts, the reactor temperature decreases in all cases. This set of experiments was used to calculate the potential use of the released energy. 3.1. Fluid Composition. The oxidation process undergone by the cutting fluid (organic compound) to be eliminated causes

Reaction 1 is exothermic and has a reaction heat of 39.2 MJ/ kg cutting fluid. That value has been determined experimentally according to UNE Standards32 for the determination of the higher calorific value (or “gross calorific value”, GCV). The maximum fluid temperature in the reactor is below 600 °C, and therefore the formation of nitrous species, like NO, NO2, or N2O, was not considered.2 Formation of these nitrous compounds requires a huge activation energy that would only be compatible with fluid temperatures higher than 1200 °C.33 Finally, the cutting fluid is not all transformed, and a certain level was found outside the reactor. The composition of the fluid that circulates in the plant is therefore as follows: at the reactor input, water, cutting fluid, O2, and N2; and at the reactor output, water, cutting fluid, CO2, O2, and N2. 3.2. Thermodynamical and Transport Fluid Properties. The thermodynamical and transport properties of the cutting fluid are only known at pressures and temperatures far from critical conditions. However, the mass percentage of this organic compound with respect to the total mass flow is between 0.3% and 1.4% for all the conditions studied, so the fluid properties were considered to be the same as for water. This assumption is consistent with most SCWO simulations reported in the literature.34-36 For each pressure and temperature considered, the properties of all pure chemical species were calculated with the code EES.37 For those analyses where a unique fluid property is required, the corresponding magnitude was evaluated through a mass average using expression 2: Bi(p, T) )

∑ m˙ B (p, T) ∑ m˙ j ji

(2)

j

where Bi is the property i of the pure chemical species j evaluated at pressure p and temperature T, and m ˙ j is the mass flow of j. 3.3. Reactor Heat Losses. The whole reactor is covered with a 46 mm thick thermal insulation layer made of 43-47% Al2O3 and 53-57% SiO2. The thermal conductivity is 0.14 W/m · K at 600 °C and 0.20 W/m · K at 800 °C. Despite the insulation material, relevant heat losses were measured. To evaluate and model heat losses, several experiments denoted as BLANK were performed by heating pressurized water and atmospheric air without the addition of organic compound. The temperature profile along the reactor tube was measured. The test conditions for three different blank tests and reactor heat losses (absolute and relative) are shown in Table 2. These losses were evaluated through an energy balance between inlet and outlet fluid conditions given by expression 3. Energy losses related to changes in kinetic energy were not considered.

778


Table 2. Test Conditions and Reactor Heat Losses for Three Different Blank Tests

water flow (kg/h) air flow (kg/h) inlet reactor temperature (°C) outlet reactor temperature (°C) inlet reactor pressure (bar) HLreactor(W) HLreactor (%) m ˙ inlet · hinlet

blank 1

blank 2

blank 3

10.62 5.34 495 336.5 250 5014

10.08 2.64 504 356.0 250 4356

9.84 3.96 522 369.7 250 4170

51.7

48.7

45.6

HLreactor ) m ˙ inlet · hinlet(pinlet, Tinlet) - m ˙ outlet · houtlet(poutlet, Toutlet) (3) where HLreactor is the reactor heat losses (W), m ˙ is the mass flow (kg/h), and h is the enthalpy (J/kg) at the temperature and pressure of the stream. It can be observed that total heat losses are close to 50% of the total reactor energy input, and the equilibrium temperature profiles are shown in Figure 3 for the three different tests. Reactor heat losses are similar in all three experiments and are in the range 51.7-45.6% of the total reactor energy input. As can be seen in Figure 3, the first reactor tract, between 0 and 0.519 m of the reactor length, produces the highest temperature drop with an overall specific heat loss of 3200 W/m (around 34% of the total heat lost). For the rest of the reactor length, between 0.519 and 9.48 m, the specific heat losses are around 356 W/m. This huge variation in the specific heat losses along the reactor tube, along with the irregular temperature drop distribution observed, suggests a nonconstant local heat transfer coefficient. This could be due to degradation and contamination of the thermal insulation material. 3.4. Fluid Flow with Organic Compound. The reactor heat losses (absolute and relative) for the Biocut oxidation experiments (see operating conditions in Table 1) are shown in Table 3. These losses were evaluated through an energy balance between inlet and outlet fluid conditions and heat addition from cutting fluid oxidation given by expression 4, where it was only considered the organic fraction that has been converted by oxidation. Energy losses related to changes in kinetic energy were not considered. HLreactor ) m ˙ inlet · hinlet(pinlet, Tinlet) + Hp · m ˙ converted organics m ˙ outlet · houtlet(poutlet, Toutlet) (4) ˙ is the stream mass where HLreactor is the reactor heat losses (W), m flow (kg/h), h is the enthalpy at the temperature and pressure of ˙ converted organics the stream (J/kg), Hp is heat power (J/kg), and m accounts for the mass flow of organics that have been oxidized (kg/h). An important conclusion from the results in Table 3 is that relative ˙ inlet · hinlet + Hp · mconverted organics), diminish reactor heat losses, (HLreactor/m from 47.0% to 14.6% as relative organic mass flow increases (referred to inlet water and air flow). This change could be due to the fact that as the organic mass flow increases, a bigger fraction of the whole chemical energy of the organic compound is released far from the reactor inlet, that is, at reactor tracts with low heat losses. The measured reactor heat losses, for runs 5 and 6, show thermal energy losses of 76% and 63%, respectively, referred to the total energy liberated by the cutting fluid. This finding rules out the possibility of any subsequent efficient energy use. From the point of view of the potential use of the released energy for other purposes such as heating, steam generation, cogeneration, or electrical power generation, the current

Figure 3. Temperature profiles along reactor length for three blanks and inlet temperatures.

plant performance is not suitable. To analyze the potential for energy use from SCWO in a well isolated reactor, a complete plant model has been developed in the following section. Once the simulator has been validated, it will be used to make predictions for energy production from a similar SCWO plant on an industrial scale. 4. Reactor Model 4.1. Cutting Fluid Chemical Kinetics. The reaction rate is dependent upon temperature according to the Arrhenius law. Moreover, this rate depends on the local concentration of waste and oxygen.30 Finally, the reaction rate is expressed as follows: rCOD ) -

(

)

Ea d[COD] ) A · exp · [COD] · [O2]0.579 dτ R·T

(5)

In eq 5, A stands for the pre-exponential factor of Arrhenius law (A ) 35 (mg O2/L)-0.579 · s-1), and Ea is the activation energy of reaction 1 (Ea ) 63 000 J/mol). 4.2. Governing Equations. The governing equations, shown in their derivative form in Table 4, postulate that in the stationary mode of operation, mass, species, momentum, and energy are conserved over a control volume in the reactor. The density, viscosity, and specific enthalpy of each species depend on pressure and temperature and were obtained using EES.37 Steam properties were obtained from the IAPWS formulation for pure water. The thermodynamic properties of carbon dioxide were obtained from the fundamental equation of state developed by Span and Wagner,39 and viscosity was determined in the manner described by Vesovic et al.40 The thermodynamic and transport properties of oxygen and nitrogen were obtained from the literature.41-43 4.3. Reactor Heat Losses. Heat transfer from the fluid to ambient air is considered to be axisymmetric and is produced mainly in the radial direction. As a consequence, neither the heat transferred along the pipe nor the length of isolation material were considered. The heat balance for the calculation of heat losses was formulated using an electrical analogue (see Figure 4). Fluid reactive flow transfers heat by convection to the reactor pipe via resistance R1 and by conduction through the reactor thickness (resistance R2). Heat is evolved through the insulating material (resistance R3), and finally heat is lost to the surroundings by convection (resistance R4). Conventional correlations were used to estimate of the heat transfer coefficients for fluid flow and air. Heat losses are not


779

Table 3. Reactor Heat Losses for Biocut Oxidation Experiments

m ˙ converted organics(g/h) m ˙ air inlet(kg/h) m ˙ water inlet(kg/h) m ˙ organic compound (%) m ˙ water inlet + m ˙ air inlet heat power from converted organic compounds (W) inlet reactor temperature (°C) outlet reactor temperature (°C) HLreactor(W) HLreactor (%) m ˙ inlet · hinlet + Hp · mconverted organics

constant due to the evolution of the fluid temperature along the reactor length (dimension x). The following equations hold for heat losses: R1 )

1 hfluid2πr1∆x

R2 )

R3 )

ln(r2/r1)

(7)

2πkAISI∆x ln(r3/r2)

(8)

2πkINSULATE∆x

R4 ) Qloss(x) )

(6)

1 hair2πr3∆x

(9)

1 (T (x) - T∞) (R1 + R2 + R3 + R4) f

(10)

where hair and hfluid are the convective heat transfer coefficients for air and fluid, respectively; kINSULATE and kAISI are the conductivities of the insulating and pipe materials, respectively; and ∆x is the incremental reactor length considered in each calculus step. For calm air, hair can be taken as 6-10 W/m2 · K. The convective heat transfer coefficients for the reactor fluid flow were calculated from the Gnielinski correlation:44

run 1

run 2

run 3

run 4

run 5

run 6

45.3

72.5

112.9

227.0

361.7

379.1

4.66

4.07

3.40

6.60

6.72

10.83

10.39

10.55

10.32

10.39

16.17

16.51

0.30

0.47

0.82

1.32

1.55

1.37

493.5 430.0 346.0 4205

789.4 430.0 370.0 3936

1229.0 430.0 389.0 2618

2472.0 430.0 411.0 2827

3938.0 430.0 480.0 2517

4128.0 430.0 458.0 3141

47.0

42.3

27.5

25.0

14.6

17.2

hfluid )

(f/8)(Re - 1000)Pr 1/

1 + 12.7( /8) (Pr f

2

2/

3

k 2r - 1) 1

where hfluid is the fluid convective heat transfer coefficient (W/ m2 · K), f is the Darcy friction factor, k is the thermal conductivity (W/m · K), r1 is the radious (m), and Pr and Re are the Prandtl and Reynolds numbers. The fluid properties were evaluated at the mean pressure and temperature for each differential reactor length ∆x. The Reynolds number for mass flow studied in section 3.3 and the pressure and temperature conditions at the reactor entrance is close to 20 000. Therefore, turbulence is fully developed, and eq 11 can be applied. This equation gave a very high hfluid value (>2500). As a result, the thermal resistance R1 is much smaller than R3 + R4, and the heat transfer problem is almost independent of heat resistance R1 and is therefore also independent of hfluid. 5. Model Results The relevance of the thermal resistance R3 on reactor performance was highlighted by the simulation model applied to the flow conditions of run 5, see Table 3, and to the maximum reactor flow rate capacity (25 kg/h) denoted as BIOCUT 25, see Table 5. The analysis carried out shows the heat losses and reaction evolution for insulating materials with different thermal conductivities and isolation thicknesses. Variations in the convective heat transfer coefficient for insulating surrounding

Table 4. Governing Equations in Derivative Form F stands for the local density of the mixture and u for the velocity of the mixture

total mass conservation:

∂Fu )0 ∂x

momentum conservation:

-

∆ξ ∂P )0 ∂x ∆x

P stands for the local pressure of the reacting media, direction x, and (∆ξ)/(∆x) represents the local linear pressure drop inside the reactor38

species conservation:

∂u · Fyj - Rj ) 0 ∂x

Rj is total chemical reaction rate (on a mass basis) of species j; yj stands for the mass fraction of species j

energy conservation:

∂Fuh )∂x

∆rHi stands for the standard heat of reaction i, h represents the specific enthalpy of the medium, and the term wlost(x) is equal to Qloss(x)/(∆xπr21) (Qloss(x) is defined in section 4.3)

Nreac

wlost -

∑r∆H i

i)1

r

i

(11)

780


Table 5. Operating Conditions and Simulation Results for the Experiment Denoted BIOCUT 25 BIOCUT 25 water flow (kg/h) air flow (kg/h) inlet reactor temperature (°C) m ˙ organic compound(g/h) m ˙ organic compound (%) m ˙ water inlet + m ˙ air inlet heat power from organic compound (W) COD inlet (g O2/L) reactor pressure (MPa)

24.8 10.1 430 555.1 1.59

6044.6 50.55 25

air are in the expected range 6-10 W/m2 · K, and these only have a minor influence on the reactor performance. As a result, this factor was not analyzed. The reactor fluid temperature and the evolution of chemical species are shown in Figures 5 and 6, respectively, for run 5 and Figures 7 and 8 for BIOCUT 25. Relative heat loss (percentage) was evaluated as the heat power loss relative to the total input power (Hp · m ˙ from the organic compound). The huge influence of thermal conductivity and isolation thickness on heat power loss through the equivalent thermal resistance R3 + R4 (see eqs 6-10) can be appreciated, and this was evaluated for a constant ∆x ) 0.05 m. Although a higher thermal resistance could be considered to reduce heat losses, this situation was not simulated because the fluid temperature would surpass the reactor safety limit. Although increasing the mass flow rate up to the maximum fluid flow compatible with plant design, that is, in BIOCUT 25, reduces the global heat losses

Figure 6. Species evolution for run 5 conditions (see Table 3) for different insulations.

Figure 7. Fluid temperature evolution for BIOCUT 25 (see Table 5) for different insulations.

Figure 4. Electrical analogue used for the calculation of heat losses.

Figure 8. Species evolution for BIOCUT 25 (see Table 5) for different insulations.

Figure 5. Fluid temperature evolution for run 5 conditions (see Table 3) for different insulations.

to 14% for the best insulation design considered (see Figure 7), heat losses continue being very high. It can be concluded that this reactor configuration is not compatible with the maximum feasible insulation, which precludes the recovery of all the energy content of the waste fluid. To increase recovered energy, the reactor design must be modified.


781

Table 6. Flow Conditions at the Inlet and Outlet for Both Heat Exchangers point (Figure 1)

fluid composition

fluid temperature (°C)

2 2_air 3 3_air 4 5 6 7

water with cutting fluid air water with cutting fluid air air and water with cutting fluid O2, CO2, N2, and water O2, CO2, N2, and water O2, CO2, N2, and water

30.0 30.0 447.1 195.1 430.0 581.0 196.1 180.1

5.1. Water and Air Heat Exchanger Models. To evaluate the potential for energy recovery from the plant, we considered the BIOCUT 25 flow conditions (see Table 5), and the length of the heat exchangers required to obtain a temperature of 430 °C at point 4 was calculated (see Figure 1). The heat exchanger model was developed on the assumption that heat losses from the exterior tube are effectively zero. Hot inlet flow conditions, point 5 in Figure 1, are given by Figures 7 and 8 with the Reynolds number evaluated at 581 °C having a value of 63 500. Therefore, turbulent flow is completely developed, and the heat transfer coefficient at the inlet can be estimated from eq 11, taking r1 ) 2.75 mm, and gives a value of 2580 W/m2 · K. The cold water inlet flow, point 2 in Figure 1, is 24.8 kg/h at 30 °C and 25 MPa. Under these flow conditions, the Reynolds number is 500, and therefore water flow is laminar. However, a very high temperature increase is expected for a cold inlet flow, with a final temperature above 400 °C; this changes the Reynolds number to 14 000, and therefore fully developed turbulent flow should be considered at the heat exchanger outlet, point 3 in Figure 1. Furthermore, at a constant water pressure, the specific heat changes abruptly in the proximity of critical point, and therefore the heat exchanger performance should be based on the evaluation of fluid properties at each pressure and temperature value along heat exchanger length, instead of using mean temperatures between the inlet and outlet. For both heat exchangers, the fluid model is the same as that developed in section 4.2, except that the chemical reaction is stopped, and therefore the fluid composition is fixed at the reactor outlet. Heat transfer for both heat exchangers was evaluated through local heat transfer coefficients: the Gnielinski correlation44 for turbulent flow and according to Shah and London45 for laminar flow conditions. Transitional flow is assumed to occur for Reynolds numbers between 2300 and 3000, and interpolation is applied between the laminar and turbulent correlations. The flow conditions at the inlet and outlet for both heat exchangers are given in Table 6 for the reactor design with the minimum heat losses, that is, the system with insulation material that has a conductivity of 0.05 W/m · K and a thickness of 45.7 mm (see Figures 7 and 8). The fluid evolution through the water-fluid heat exchanger is represented in Figure 9. A flat temperature profile can be observed for both fluids when the temperature is close to the critical water temperature. The fluid evolution through the air-fluid heat exchanger is represented in Figure 10. 6. Waste Heat Recovery A previous analysis shows that this plant design produces low-quality heat energy, mainly because the flow temperature at point 7 (Figure 1) is very low, and therefore the number of feasible energy demanders is reduced. For a temperature range

Figure 9. Temperature evolution through the water-fluid heat exchanger length for BIOCUT 25 (see Table 5).

Figure 10. Temperature evolution through air-fluid heat exchanger length for BIOCUT 25 (see Table 5).

of 150-250 °C (and despite the fact that the fluid pressure is above the supercritical point for water, CO2, and O2), the fluid specific enthalpy is an almost linear function of temperature, which means that the specific enthalpy is far from the extremely high value at the critical temperature. For this reason, power generation based on the Rankine cycle is not appropriate for the current plant design because the expected maximum attainable steam temperature is about 140-160 °C, which markedly reduces the steam cycle efficiency and therefore the energy recovery. Direct expansion of the supercritical fluid through a turbine is also unfeasible with the present technology. Two aspects should be taken into account in this respect: For typical industrial supercritical plant design with a fluid flow of 1000 kg/h before supercritical compression and for thermodynamic properties evaluated at the conditions of point 7, the expected turbine entrance area is markedly reduced. This would produce a very small blade height, and therefore the turbine performance would be dramatically reduced. Further research on this issue needs to be carried out. Extremes in pH, high concentrations of dissolved CO2 and O2, and ionic inorganic species13,46 would accelerate the corrosion of rotor and stator turbine blades. In addition, the presence of stable solid-matter particles can cause rotor and stator blade fouling and erosion, which would markedly reduce the expected turbine life.

782


Table 7. Flow Characteristics Considered for Simulation of an Industrial Plant Design with 1000 kg/h fluid composition

%

water nitrogen carbon dioxide oxygen

71.8 22.8 3.8 1.6

mass flow (kg/h) temperature (°C) 717.8 228.3 37.6 16.3

185

pressure (bar) 250

Despite the difficulties in power generation, other methods more appropriate for energy recovery at low temperature and high fluid pressure can be considered. The authors propose the following approaches: water heating and steam generation. The following analyses were performed for a typical industrial supercritical plant design with 1000 kg/h of fluid flow before supercritical compression and whose flow characteristics, corresponding to point 7 (Figure 1), are given in Table 7. The fluid temperature was estimated to be a little higher than that shown in Table 7, because for higher mass flow and similar fluid speed the pipe diameter increases and therefore heat loss decreases. 6.1. Atmospheric Pressure Water Heating. Domestic hot water should be stored at temperatures around 65 °C to avoid infectious disease caused by the aerobic bacterium Legionella, which thrives at temperatures between 25 and 45 °C, with an optimum around 35 °C. Heat recovery for the flow conditions shown in Table 7 was modeled with a countercurrent flow heat exchanger between supercritical fluid and domestic atmospheric pressure water. From Figure 11, the heat exchanger efficiency is defined by eq 12, with a typical design value of ε ) 0.7. ε)

T7 - T8 T7 - T1DW

(12)

From the energy balance and eq 12, the maximum domestic water flow that could be heated to 65 °C can be evaluated for different T1DW temperatures. The domestic water flow as a function of inlet water temperature T1DW is shown in Figure 12. The power extracted from the supercritical fluid goes from 118.3 to 98.8 kW for inlet water temperatures of 5 and 35 °C, respectively. 6.2. Steam Generation. Low steam pressure can be generated from the fluid conditions outlined in Table 7. Steam generation equipment is essentially a countercurrent flow heat exchanger composed of a series of superheater, boiler, and economizer sections positioned from the supercritical flow inlet to the outlet to maximize heat recovery. An alternative is the once-through design, which consists of one or more serpentine circuits encompassing the economizer, boiler, and superheater sections. The supercritical flow and steam evolution for typical design parameters for one pressure steam generation equipment is represented in Figure 13. The maximum steam temperature attainable results from a design temperature drop of 15 °C. The pinch point could be higher than 15 °C, which would reduce the generator cost, but as the pinch point temperature increases the recovered energy decreases. For

Figure 11. Domestic water heat exchanger.

Figure 12. Domestic water flow that could be heated to 65 °C as a function of the inlet temperature.

Figure 13. Supercritical flow and steam evolution for a water inlet temperature of 30 °C.

Figure 14. Maximum attainable mass steam flow for different design steam pressure.

each design steam pressure, there is a maximum attainable steam mass flow, and this is represented in Figure 14. As can be observed, the higher is the maximum required steam temperature, the lower the maximum steam flow is for each design steam pressure. In addition, the admissible


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Figure 15. Recovered energy (left) and outlet supercritical flow temperature (right).

pressure range diminishes with the maximum steam flow temperature because of the saturation temperature associated with each steam pressure. The inlet water temperature also affects heat recovery. The heat transferred to steam, for a maximum steam temperature of 170 °C, is represented in Figure 15 (left) as a function of steam pressure and for two typical water inlet temperatures. From the energy recovery point of view, it is always advantageous to generate steam at as low a pressure as possible. However, the maximum steam temperature and design pressure are fixed by energy consumer equipment, and it is not always feasible to maximize energy recovery. Finally, a reduction in the thermal power transmitted to steam always leads to an increase in the supercritical flow outlet temperature from the steam generator, see Figure 15 (right). 7. Conclusions Supercritical water oxidation is an effective technology for the treatment of cutting fluid wastes, with more than 97% COD removal obtained for Biocut at 500 °C. If high feed concentrations are treated (>40 g O2/L), the heat released during oxidation produces a marked increase in temperature in the reaction medium, leading to higher reaction rates and better COD conversion. For the current experimental supercritical plant, it is very advisible to improve the thermal isolation system to reduce heat losses and increase organic compound conversion. The lengths of heat exchangers can be reduced accordingly, which will in turn reduce future plant cost. From the point of view of power generation, the current plant design is not appropriate, mainly because of the low temperature of the supercritical flow at the end of the air heat exchanger, which dramatically reduces the efficiency of any feasible Rankine cycle. The direct expansion of supercritical fluid in the current turbine design cannot be considered as an alternative. Heat production, hot water or steam, at current plant design is valid at low cost. For a supercritical flow of 1000 kg/h (water and air), the recovered energy goes from 118 kW (1700 m3/h of hot water at 65 °C) to 75 kW (100 kg/h of steam flow at 1.1 bar and 170 °C). If we consider a rate of 15.9 kg/h of cutting fluid (same proportion as BIOCUT 25) with a conversion of 95%, the waste heat energy recovered would be between 71% and 45.6% of the energy content of the organic compound, a level that can be considered as high thermal efficiency.

Acknowledgment We thank the Junta de Andalucıá for economic support through the Research Project P07-RNM-03276. Nomenclature COD ) chemical oxygen demand (mg O2/L) E.E.S ) engineering equation solver Bi ) thermo dynamical property i j ) pure chemical species mj ) mass flow of species j (kg/h) P ) pressure (bar) T ) temperature (°C) Bji ) thermo dynamical property i of the pure j chemical species HLreactor ) reactor heat losses (W) m ˙ inlet ) inlet mass flow (kg/h) hinlet ) inlet enthalpy (J/kg) Pinlet ) inlet pressure (bar) Tinlet ) inlet temperature (°C) m ˙ outlet ) outlet mass flow (kg/h) Houtlet ) outlet enthalpy (J/kg) Poutlet ) outlet pressure (bar) Toutlet ) outlet temperature (°C) Hp ) calorific value (J/kg) m ˙ organic compound ) organic compound mass flow (kg/h) m ˙ water inlet ) water inlet mass flow (kg/h) m ˙ air inlet ) air inlet mass flow (kg/h) rCOD ) reaction rate in terms of COD removal (mg O2/L · s) τ ) residence time (s) A ) pre-exponential factor of Arrhenius law ((mg O2/L)-0.579/s) Ea ) activation energy (J/mol) R ) universal gas constant (J/mol K) IAPWS ) International Association for the Properties of Water and Steam F ) mixture local density (kg/m3) u ) mixture velocity (m/s) x ) lineal direction (m) ∆ξ ) pressure drop (bar) Rj ) total chemical reaction rate of j species (kg/m3 · s) yj ) mass fraction of species j wlost ) energy loss per time unit and volume (J/m3 · s) ri ) molar consumption of i species per unit time and volume (mol/ m3 · s) Qloss ) energy loss per unit time (J/s) R1,R4 ) convective heat transfer resistances (K/W) R2,R3 ) conductive heat transfer resistances (K/W) Tf ) fluid temperature (°C) Tw ) wall temperature (°C) Ts ) surface temperature (°C)

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TI ) insulating temperature (°C) T∞ ) surrounding air temperature (°C) r1, r2, r3 ) radius (m) k ) thermal conductivity (W/m · K) ∆x ) incremental reactor length (m) hair ) air convective heat transfer coefficient (W/m2 · K) hfluid ) fluid convective heat transfer coefficient (W/m2 · K) f ) Darcy friction factor Re ) Reynolds number Pr ) Prandtl number ε ) heat exchanger efficiency TDW ) domestic water temperature (°C)

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ReceiVed for reView May 27, 2010 ReVised manuscript receiVed September 23, 2010 Accepted November 11, 2010 IE101166J

Supercritical Water Oxidation of Oily Wastes at Pilot Plant: Simulation

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