Experimental Performance and Modeling of a New Cooled-Wall

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Ind. Eng. Chem. Res. 2009, 48, 6262–6272

Experimental Performance and Modeling of a New Cooled-Wall Reactor for the Supercritical Water Oxidation M. Dolores Bermejo,* Daniel Rincon, Angel Martin, and M. Jose´ Cocero High Pressure Process Group, Department of Chemical Engineering and EnVironmental Technology, UniVersity of Valladolid, Prado de la Magdalena s/n 47011, Valladolid, Spain

For the commercial application of the supercritical water oxidation (SCWO), new reactors that are able to withstand the harsh operational conditions of the process must be developed. In this work, a new scaled-up design of the cooled-wall reactor (CWR) is presented, which isolates the temperature and pressure stresses caused by SCWO processes. The new reactor consists of a series of chambers designed as a heat exchanger to maximize the residence time and minimize the temperature gradient inside the reactor, allowing a better thermal protection of the pressure shell. This reactor has been implemented in a demonstration scale plant located in the industrial site of a waste management company. Experimental results obtained in this plant with five prototypes of the reactor are presented, explaining the strong and weak points of the designs and the modifications proposed to overcome the operational problems. Experimental results obtained working with oxygen and air as an oxidant and feeds containing isopropyl alcohol and mixtures of isopropyl alcohol/ ammonia are thoroughly examined and compared. To have a better understanding of the behavior of the reactor a mathematical model specifically developed for the new reactor is presented. Introduction Supercritical water oxidation (SCWO) is an advantageous process for destroying organic wastes. Its commercial development has been delayed by serious problems of corrosion, salt deposition, and high energy consumption. First patented by M. Model in 1980,1 the technology has been extensively studied. More than 15 commercial pilot and full-scale plants have been built in several locations around the world for treating a variety of wastes. At least five of them have been shut down permanently. It is known that mechanical or operational problems have been an important factor in at least two of these shutdowns:2,3(1) The first plant for the destruction of sludge, built by HydroProcessing that began its operation in April 2001 in Harlingen (Texas, USA), is inactive due to corrosion issues.4,5(2) The plant developed by Foster & Wheeler for the U.S. Army, using the patented design of the transpiring wall reactor, is inactive owing to linear mechanical issues and limited funding.2,4,6,7 In addition several of the companies that have commercialized this technology are no longer in the SCWO field,2 as for example Foster and Wheeler and HydroProcessing. Other companies were sold, as MODAR, which was acquired by General Atomics, and EcoWaste Technologies, which was acquired by Chematur and later sold as a part of its Supercritical Fluids Division to SCFI (Supercritical Fluids International). The company SCFI currently commercializes the SCWO AquaCritox Process, claiming that the process has been enhanced by producing electricity from the excess energy.8 To favor the commercialization of the process, research must focus on the design of new reactors and the use of appropriate construction materials. The design of these new reactors must consider some key aspects: (i) Reaction temperatures in the range of 650 to 700 °C are necessary to achieve complete elimination efficiencies, but only a few materials can withstand these conditions (Ni alloys, alumina). New reactor designs are evolving in order to have * To whom correspondence should be addressed. E-mail: mdbermejo@ iq.uva.es.

the high temperature area confined by a cool protecting film, for instance created with clean water (transpiring wall reactor, reverse flow reactor with a brine pool). This film would also protect the wall against salt deposition. On the other hand, filmcooled reactors keep the wall refrigerated by coaxial introduction of large amounts of water. This allows having a reaction chamber constructed with a high temperature resistant material while the pressure vessel is maintained at a lower temperature and it can be constructed with a cheaper material. (ii) Plugging of the reactor inlet and of the preheaters must be avoided. Salt precipitation occurs when the mixture becomes supercritical (at about 350-375 °C, depending on the composition of the mixture). The simplest solution is to maintain subcritical conditions in these sections, reaching supercritical conditions in an area of the reactor where plugging is not probable (salt precipitation chamber, wider reaction chamber, area protected by a cool film or a transpiring flow). (iii) Heat recovery. A lot of energy is necessary for preheating and pressurizing the feed and the oxidant. It is convenient to recover at least part of this energy for preheating the feed or even for producing electricity. Reactors that can process feeds at relatively low temperatures and those in which preheating takes place inside the reactor are more efficient for the energy integration. The cooled-wall reactor is a film-cooled reactor developed by the High Pressure Processes Group (HPPG) of the University of Valladolid (UVa) (Spain). In this type of reactor, temperature and pressure stresses are isolated. This is achieved by using an external cooled-wall pressure vessel and an internal reaction chamber, where reactants are mixed and the reaction takes place. This reaction chamber is made of a special material to withstand the oxidant atmosphere of the reactants at a maximum temperature of 800°. It is enclosed in the pressure vessel, which is pressurized and cooled with the feed stream before entering the reaction chamber. Thus the pressure vessel works at about 400 °C and it is not in contact with the oxidant atmosphere. This pressure vessel can therefore be made of stainless steel with a relatively low thickness. Another additional advantage of this reactor design compared to other film-cooled reactor designs is

10.1021/ie900054e CCC: $40.75  2009 American Chemical Society Published on Web 06/03/2009

Ind. Eng. Chem. Res., Vol. 48, No. 13, 2009

that it allows feed preheating inside the reactor, resulting in a compact unit very appropriate for mobile units. The cooled-wall reactor was developed in a 30 kg/h pilot plant and tested with industrial wastes such as cutting oils and PET effluents9-11 obtaining excellent results. Since 2001, a scaled up unit of the cooled-wall reactor is working in the demonstration plant located in the site of the firm CETRANSA in Santovenia de Pisuerga (Valladolid, Spain), with a treatment capacity of 200 kg/h of waste. In the new reactor, the bed reaction chamber has been substituted by a series of chambers designed as a heat exchanger to maximize the residence time and minimize the temperature gradient inside the reactor12 and at the same time improving the protection of the pressure shell from the high temperatures inside the reactor. In this work, experimental results of several prototypes of this new CWR are presented. Operation with the different prototypes is analyzed, focusing on the operational problems and proposing solutions. Additionally, a mathematical model specifically developed for the new cooled-wall reactor is presented. It consists of a steady-state model that considers the reactor divided into different parts with different simple flow patterns: a section with plug flow to represent the mixer, a perfect mixing zone that represents the outlet of the mixer, and several plug-flow sections in series that represent the annular chambers designed as heat exchangers. Mass and energy balances and heat transfer between the different chambers are considered. This model helps to gain an insight of the behavior of the reactor, paying special attention to reaction progress and heat transfer at different working conditions in order to complement experimental results, because this information is difficult to obtain experimentally owing to the complex design of the equipment. A number of models of SCWO have been developed in the last years.13,14 In previous works, our group has developed simple flow pattern models to describe the behavior of a transpiring-wall reactor15 and of a cooled-wall bed reactor.16 Experimental Section The new cooled-wall reactor consists of a mixer, three concentric tubes, a salt precipitation chamber, a lid covering the reaction chamber, and the pressure vessel with a flange. The inner tubes are made of Ni alloy 625 to prevent corrosion. The outer pressure vessel is made of SS 316. The feed inlet is located at the top of the reactor. Feed flows down through preheating chamber (chamber 4), whereby it is preheated by the heat released by the reaction taking place in the inner reaction chamber. Two holes are located at the bottom of the reactor. The feed enters in the salt precipitation chamber through them. The temperature in the chamber is around the critical temperature of water, thus most of the salts precipitate there. This effect is strongly wanted to prevent the salts from going into the mixer, where they may cause plugging. Then the fluid flows upward from the salt precipitation chamber and enters into the mixer, where it is mixed with oxygen. The reaction starts very quickly producing a high amount of heat (because it is strongly exothermal). At the outlet of the mixer, in chamber 1, a sharp increase of temperature is detected, indicating that it is there where most of the oxidation reaction proceeds. The reacting mixture flows downward through chamber 1, upward through chamber 2, and downward through chamber 3, being cooled by heat transfer to chamber 4. The effluent outlet is located at the bottom of the reactor, at the end of chamber 3. The flow directions in the different chambers of the reactor are presented in Scheme 1.

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Scheme 1. Scheme of the Performance of the New Cooled-Wall Reactor

The main feature of this reactor is that, in principle, it can be operated without any external energy supply, due to the use of the released energy in the reaction for preheating. Thus, the reactor can be defined as energetically self-sufficient. During the starting-up periods, an electrical heater is used. Five prototypes of the reactor have been tested. They are numbered as reactors 0 to 5. The schemes of these reactors are shown in Scheme 2. They are different in the kind of mixer used, in the configuration of the salt chamber, and in the thickness and addition of reinforcement rings in the inner wall. • Reactor 0. The mixer used in this reactor consists of an alloy 625 threaded bar with a diameter of 11 mm and a length of 1325 mm, introduced on an alumina tube of the same length and internal diameter. In this mixer, feed and air flows through the small channels left between the threaded bar and the alumina tube. • Reactor 1. The mixer of this reactor consists of an alloy 625 tube 25 mm external diameter and 5 mm internal diameter, 1325 mm length filled with alumina particles of 2-3 mm size and with a porosity of 0.6. A scheme of this mixer can be found in Scheme 3, part A. • Reactor 2. The difference between reactor 2 and reactor 1 is that in reactor 1 the salt precipitation chamber is empty and in reactor 2 the chamber is filled with alumina balls of about 5 mm to improve heat transmission. • Reactor 3. The mixer used in this reactor consists of an Ni-alloy 625 threaded bar with a diameter of 6 mm and a length of 1000 mm, introduced in an alloy 625 tube of the same length and 25 mm external diameter and 6 mm internal diameter. A scheme of this mixer can be found in Scheme 3, part B. The salt precipitation chamber of this reactor is again empty. The mechanical resistance of the outer reaction chamber has been increased by constructing it with an alloy 625 plate, 5 mm thick. The thickness of the lid in the upper part of the reaction chamber has been also doubled and in addition reinforcement rings have been added in the inner wall of the third chamber of the reactor in order to avoid deformations as can be observed in Figure 1, where all the tubes used for building the chambers of the reactor are shown. • Reactor 4. This reactor is exactly the same as reactor 3 with the difference that the length of the mixer has been increase up to 1325 mm and that the salt precipitation chamber has been

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Scheme 2. Schemes of the Five Prototypes of CWRs Tested

Scheme 3. Schemes of the Mixers Used: (a) Ni-Alloy Tube Filler with Alumina Particles; (b) Screwed Bar in an NI-Alloy Tube

with a metering diaphragm pump (DOSAPRO MAX ROYAL C) until work pressure (23 MPa) is attained. In this line also exists an electrical 10 kW preheater, used to preheat the feed in the start up of the reactor. (ii) Oxygen supply facility. Liquid oxygen is stored in a cryogenic deposit, from where it is pumped by a cryogenic metering pump until work pressure. The pump supplies a constant flow. After pumping, oxygen is vaporized in a finned tube unit. Then it is stored in four reservoirs before it is mixed with the aqueous waste stream. The flow of oxygen withdrawn from these reservoirs can be controlled with a control valve.

filled with pieces of alloy 625 1/4” tubing of lengths between 5 and 20 mm with a porosity of 0.76. Experiments of reactor 0 have been performed in the pilot plant located in the University of Valladolid that has been extensively described elsewhere.9-11 Experiments with reactors 1-4 have been performed working in the demonstration plant. The demonstration plant is located in the site of the firm CETRANSA in Santovenia de Pisuerga (Valladolid, Spain). It has a treatment capacity of 200 kg/h of waste and uses oxygen as oxidant. The flow diagram of the plant is shown in Scheme 4. This plant is extensively described elsewhere.12 The main items of this facility are as follows: (i) Feed streams conditioning equipment. It consists of a tank where the feed is prepared, with special attention given to the concentration of organic matter, because this concentration determines the operating temperature of the reactor. From this tank the feed is driven to the feed tank, from where it is pumped

Figure 1. Ni-alloy tubes using for the construction of the reaction chamber. In the third tube, reinforcement rings can be observed.


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Scheme 4. Flow Diagram of the Demonstration Plant of the University of Valladolid, Located in the Site of the Firm CETRANSA in Santovenia de Pisuerga (Valladolid, Spain)

(iii) The cooled-wall reactor described before. (iv) Cooling systems. They consist of two coolers, in which the hot product of reaction flows inside a titatium alloy (Ti3Al-2.5V) coil, and is refrigerated by cooling water.

Figure 2. Temperature profiles inside of the reactor for experimental points working with air as oxidant with feed flow 30 kg/h, air excess 5-10% above the stoichiometric concentration: (TR) reaction temperature profiles in reaction chamber 1; (T-Cooling) temperature profile in chamber 4 (cooling water); (T-wall) temperatures measured outside of the pressure vessel wall. Axial coordinate z was calculated considering z ) 0 at the top of the reactor: (a) reaction point at the bottom of the reaction chamber; (b) reaction point at the top of the reaction chamber.

(v) Back pressure regulator valve: A needle valve, SENTRY VREL11, is used. For security reasons, a second valve is placed in parallel to this valve. (vi) Flash chamber separator and sampling device that allows taking samples of the liquid and gaseous effluents. The experiments presented in this work have been performed using synthetic feeds prepared with isopropyl alcohol (IPA), normally used as a fuel in SCWO, and ammonia as a model nitrogen containing compound. IPA (99% purity) and NH3 (25% purity) is provided by COFARCAS (Spain). Temperatures were measured in several points of the reactor with type K thermocouples (temperature range from 0 to 1000 °C) with an accuracy of 1% of the measurement. Air and oxygen flows were measured with a Coriolis gas flow meter with a precision of 0.2%. Liquid flows were determined by measuring the time required to pump a certain liquid volume. IPA solutions were prepared volumetrically, measuring water volume with a precision of 1 L and IPA volume with a precision of 1 mL, resulting in an experimental error of 0.3% for 6.5% and 7% (w/w) IPA solutions, and of 0.4% for 7.5% and 8% (w/w) IPA solutions. Total organic carbon (TOC) was characterized using a Shimadzu 5050 TOC analyzer (detection limit ) 1 ppm).

Figure 3. Breakage of the wall of the alumina mixer in reactor 0.

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Figure 5. TOC removal and TOC concentrations versus maximum temperature in the reactor. For reactor 1 and 2: 6 (w/w) % IPA, feed ) 30 kg/h, tR ) 160 s, 0% O2 excess over the stoichiometric. For reactor 3 and 4: 8 (w/w) % IPA, feed ) 20-25 kg/h, tR ) 200 s, 0% O2 excess over the stoichiometric.

Figure 4. Example of the deterioration of the materials of the reactor after operation with feeds with high salt contents containing Cl -. (a) Corrosion in the outer part of the Ni-alloy mixer, the point where the maximum reaction temperatures are registered. (b) Salts deposited in the lower part of the reactor.

Ammonia concentrations were determined by analysis with an ion-specific electrode (Orion Research model 95-12) (detection limit ) 1 ppm). NOx and NH3 in the gas effluent were analyzed with Draeger tube detectors, Lab Safety Supply CH29401 and CH31001 (range of detection 0.5 -100 ppm).

Figure 6. TOC removal versus sample order in reactors 1 and 2: 6 (w/w) % IPA, feed ) 30 kg/h, tR ) 160 s, stoichiometric amount of O2.

Results Operation of the Reactor. Reactor 0 was first tested in the pilot plant using air as oxidant and feed flows between 20 and 40 kg/h. An extensive amount of experiments were performed. In every experiment, the reactor was first preheated electrically until the walls of the pressure vessel reached a temperature of 400 °C. Then the reaction was initiated using IPA concentrations as high as 8.5 (w/w) %. After reaching an appropriate reaction temperature, IPA concentration was reduced and the electrical heating of the wall of the reactor was turned off. Several stationary states were reached and samples were taken. Following, the fuel concentration was gradually decreased in order to evaluate the reactor performance at lower reaction temperature stationary states. Figure 2 shows the temperature profile in different sections of the reactor. TR is the temperature profiles in the reaction chamber 1, T-Cooling is the temperature of the preheating of the feed in the outer chamber of the reactor (chamber 4), and T-wall is the temperature registered outside of the pressure vessel wall. In the T-Cooling line it is observed that with the heat released by the reaction, the feed can be preheated from temperatures lower than 200 °C to temperatures above the critical point of water. As supercritical conditions were reached in the salt precipitation chamber, it can be expected that at least part of the salts precipitated there. It is also remarkable that with temperatures near 800 °C inside of the reaction chamber, the temperatures in the wall of the pressure vessel are never higher than 400 °C. This shows that the reactor design effectively isolates the pressure shell from the high temperatures achieved in the reaction chamber. It was expected that the reaction would occur inside the mixer and in the upper part of the reaction chamber 1, just at the exit

Figure 7. Damage caused in the outer wall of the reaction chamber because of the pressure increment caused by a blockage in the mixer due to salt precipitation in the mixer. Table 1. Kinetic Parameters in the Arrhenius Equation Used to Simulate the IPA Oxidation

kA kAB kB

A (s-1)

Ea (J/mol)

2.61 × 105 2.61 × 105 2.55 × 1011

64 000 64 000 172 700

of the mixer. Thus, it was expected that the highest temperature inside of the reactor would be registered in that point. But in these first experiments with reactor 0 two kinds of temperature profiles were found with similar operational conditions. These two types of temperature profiles are shown in Figure 2. In the temperature profile showed in part a of Figure 2 the maximum temperature was registered in the lower part of reaction chamber 1, while in the profile shown in part b the maximum temperature

Ind. Eng. Chem. Res., Vol. 48, No. 13, 2009 Table 2. Dimensions of the Mixer and the Reaction Chamber of the Reactor chamber

di (mm)

do (mm)

L (mm)

mixer 1 2 3 4 (preheating)

5 55 79 125 143

25 59 90 135 145

1300 1338 1348 1348 1355

was registered in the upper part of the chamber 1 while the temperature registered in the lower part was lower than the critical temperature of water. That indicated the existence of a “liquid pool” in the bottom of reaction chamber 1. While the temperature of the reaction was decreasing (decreasing IPA concentration) the temperature in the middle of the reactor also decreased, and presumably the level of the “pool” was increasing until the moment in which the reaction was extinguished. This behavior was due to the alumina wall of the mixer not being properly fixed to the bottom of the reaction chamber. Thus, part of the reagents, instead of flowing up through the mixer, leaked out to reaction chamber 1 by the lower part of the mixer. Sometimes these reagents were at temperatures high enough to ignite and produce an additional reaction point where the reaction was happening at a temperature higher than in the top. When the temperature of the leaking reagents was lower at the start up of the reactor, they simply accumulated there, cooling down the reaction mixture that flows though them to the second reaction chamber. Both behaviors were observed in different and alternate experiments, so they were not due to a breakage in the reactor in some moment. It was observed that one or other behavior depended strongly on how high the temperature was in the bottom of chamber 1 in the moment of reaction ignition. Experiments with IPA using air as an oxidant led to TOC removal higher than 94% (effluents with less than 3000 ppm C) when the reaction was situated in the upper part of the reaction chamber and higher than 98% (TOC < 1000 ppm) when the reaction took place in the lower part. Most of the fuel was eliminated, but these results were not good enough in the SCWO technology in which eliminations higher than 99% are generally obtained. This behavior is easily explained: when the reaction was proceeding in the upper part of the chamber 1, very low temperatures (lower than the critical point of water) were registered in the lower part of chamber 1, thus the reagents passing to chamber 2 and 3 were at so low temperature that the oxidation reaction reached very low reaction rates. Experiments using oxygen as an oxidant were performed in the same reactor with flow rates as high as 70 kg/h. With these conditions, the reaction always took place in the lower part of the mixer, indicating that, as it can be expected, ignition is easier with oxygen than with air as an oxidant. In these experiments TOC removals were always higher than 99% (TOC < 500 ppm C). Figure 10 in the Supporting Information shows the TOC removal and TOC concentrations in the effluent versus the maximum temperature registered in the reactor. Experiments with different O2 excess over the stoichiometric were performed. The results are shown in Figure 11 of the Supporting Information. TOC removal was enhanced when increasing the O2 flow. No enhancement was found for O2 excesses higher than 30% over the stoichiometric. The fact that so high O2 excess was needed evidenced that the mixing of the O2 and the feed was not efficient. This confirms an improper performance of the reactor as was suspected. Experiments using increasing concentrations of NH3 of 1, 3, and 5 (w/w) % combined with IPA as a fuel (in order to increase

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the reaction temperature) were performed trying to reproduce the experiments performed with the old CWR in a previous work.17 TOC removals were very high for all the NH3 concentrations, the ones corresponding to feeds with 5 (w/w) % in ammonia being slightly lower. Ammonia removal varied from 55 to 95%, being enhanced with higher temperatures. Even when temperature was as high as the ones reached in the previous work, this time total ammonia removal was not obtained. This fact gives further evidence that the reactor was not working properly (existence of leaking, dead volumes, etc). Concentrations of NH3 and NOx was also investigated in gas samples using Draeger tubes and in all cases NH3 concentration was under 5 ppm and NOx concentration was under 0.5 ppm. TOC and ammonia removals for ammonia concentrations of 1, 3, and 5 (w/w) % and different IPA concentrations are shown in Figure 12 of Supporting Information. Operation with high concentrations of ammonia was highly unstable, with sudden increases in pressure. Damage in the reactor was suspected and it was dismantled, with breakage in the alumina wall found in the lower section of the mixer, as shown in Figure 3. Apparently a plug in the mixer caused a sudden increase in pressure and the breakage of the alumina. It was demonstrated that the mixer with the wall of alumina was not appropriate for this reaction design. Without a proper union between the metal and the alumina, leaking and finally breakage of the material were produced. Thus, in subsequent designs, mixers were made using Ni-alloy tubes welded to the lower part of the reactor. Experiments with reactors 1-4 were performed using O2 as oxidant in the industrial site of the company Cetransa, in Santovenia de Pisuerga (Valladolid). The rest of the experiments were performed with synthetic wastes in water and IPA. The water available in the process plant was a nontreated water containing a high amount of mineral salts (pH ) 8, conductivity ) 1640 µS/cm; Na ) 325 mg/L; Cl- ) 275 mg/L, SO42- ) 650 mg/L). Thus, salt precipitation caused a great number of operational problems inside the reactor. The high content of salts, especially chlorides in the feed lead to corrosion in several parts of the reactor, especially at the outlet of the mixer, where the maximum temperatures were registered as shown in Figure 4, part a. Even when salt precipitation was produced mainly, as expected, in the salt precipitation chamber, salts were also precipitated in other zones of the reactor. Salts deposited inside the reaction chamber after a few weeks of experiments are shown in Figure 4, part b. The other big challenge in the operation with the demonstration plant was the oxygen supply. The oxygen was stored at a pressure between 23 and 25 MPa in the storing vessel and from there flowed to the reactor by pressure difference. When the fowling caused a higher pressure drop along the mixer, oxygen was not able to enter inside the reactor steadily, causing great instability inside of the reactor, and forcing the reaction to stop. It has to be noted that using reactor 0, experiments as long as 6 h were performed (limited by the maximum O2 storage capacity of the Dewar vessel available for operation in the pilot plant). In this case the same oxygen supply system was used but the feeds were prepared with tap water with a much lower salt content. In the demonstration plant using the well water described before, maximum operating time was 78 min. TOC removal in reactors 1-4 is shown in Figure 5. In general TOC removals were good in reactors 1 and 2 working at temperatures higher than 578 °C. But a few points were observed in which the TOC contents of the samples was really high despite the high reaction temperatures. These data normally

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Scheme 6. Strategy of Resolution in the Modela

a i ) the number of iterations; R1, R2, R3, and R4 ) the reaction chambers 1-4; TR1-TR4 ) the temperature profiles in chamber 1-4; Tin ) the temperature of the reaction mixture in the inlet of the mixer; and Tout ) the temperature of the feed at the end of chamber 4.

Figure 8. Comparison of experimental and calculated temperature profiles. IPA conversions and acetic acid flow for different working conditions: (T) temperature, (XIPA) IPA conversion, (FHAc) acetic acid molar flow (mol/h), (m) mixer, (R) reaction chamber, (1) chamber 1, (2) chamber 2, (3) chamber 3, (4) chamber 4 (preheating of the feed), (Exp) experimental values. Note that in the mixer and in chamber 2 the sense of flow is opposite. Table 3. Predictions of the Model for the Reactor Working at Different Feed Flow Rates feed (kg/h)

Tmax (°C)

T inlet mixer (°C)

T outlet (°C)

50 100 150 200

614 662 687 702

376 371 367 365

397 425 453 475

were taken after a time of operation, as can be shown in Figure 6. It was observed that for a number of experiments TOC

removal in the liquid samples decreases drastically with operational time. Thus, reactors 1 and 2 were operating correctly for short periods, in which extremely high eliminations were obtained. But the design of the mixer was not appropriate for feeds with high salt concentrations. The mixer was frequently obstructed and even completely plugged. When this happened, the feed pressure over the wall of the reaction chamber increased sharply. The wall had not been designed for standing high pressure and


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Figure 9. Predictions of the model for the reactor working with a feed flow of 200 kg/h, Tinlet ) 300 °C 6 (w/w)% IPA, 11% excess stoichiometric. (T) temperature, (XIPA) IPA conversion, (FHAC) acetic acid molar flow (mol/h), (m) mixer, (R) reaction chamber, (1) chamber 1, (2) chamber 2, (3) chamber 3, (4) chamber 4 (preheating of the feed). Note that in the mixer and in chamber 2 the sense of flow is opposite. Table 4. Predictions of the Model for the Reactor Working at Different Feed Inlet Temperatures in the Shell T inlet shell (°C) T max (°C) T inlet mixer (°C) 300 250 200

736 705 314

T outlet (°C) TOC (ppm)

365 358 259

476 473 308

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Table 5. Predictions of the Model for the Reactor Working at Different Fuel (Ipa) Concentrations CIPA(w/w)

T max (°C)

T inlet mixer (°C)

T outlet (°C)

TOC (ppm)

6.0% 5.0% 4.5% 4.0%

706 669 599 460

365 364 359 354

475 464 445 418

1999

collapsed, as shown in Figure 7. In this case the main advantage of this cooling-wall reactor, that is, feed preheating in the reactor and cooling of the pressure vessel, became a very serious weak point of the design, being the cause of a serious breakage. Thus, the possibility of a pressure increase in a nonpressure resisting chamber must be considered carefully when designing a cooledwall reactor and in general every kind of film-cooled reactors. Reactor 3 and 4 were constructed with a thicker wall and some reinforcement rings in order to minimize the collapse of the wall in the case of sudden plug. The mixer was designed as the mixer of reactor 0, but in this case using a tube of alloy 625 to contain the screwed bar. The pressure drop of this mixer was not as high as in the previous one, but a proper mixing was not produced. Figure 6 shows that in most of the experiments with this mixer reaction temperatures were lower than 500 °C (even with working with IPA concentrations as high as 8.5 (w/w) % IPA), being in most cases the conversions very low. In summary, five experimental prototypes of the new cooledwall reactor were evaluated. In the first one a mixer with the wall of alumina was tested. The alumina wall could not be properly fixed to the metal parts and some leaking was produced, causing recirculations and shortcuts in the flow inside the reaction chamber inside of the reaction. In the subsequent designs the wall of the mixer was made of Ni-alloy welded to the reaction chamber to avoid leaking. The second mixer consisting of a tube filled with alumina particles resulted to be the most efficient (TOC removals > 99%), but the salt precipitation inside it deteriorated the reactor performance after 1 h of reaction and caused serious breakage problems in the walls of the reaction chambers. The salt precipitation chamber worked properly but salt precipitation was observed in other

areas of the reactor. The presence of salts and chlorides in the feed also produced corrosion especially at the outlet of the mixer. Modeling A model of the new CWR has been developed following the main assumptions of other SCWO reactor modeling done by the High Pressure Process Group of the University of Valladolid.15,16 A stationary model is considered. Only mass and energy balances have been considered while the momentum equation has been neglected. The pressure along the reactor has been considered constant and equal to 23 MPa. The kinetics pathways used are those reported by Li et al.,18 considering the acetic acid (HAc) as the only intermediate in the oxidation of the isopropyl alcohol. The reactions considered are the following: kA 9 C3H7OH + O2 98 3CO2 + 4H2O 2

(R1)

kAB 3 3 C3H7OH + O2 98 CH3COOH + H2O 2 2

(R2)

kB

CH3COOH + 2O2 98 2CO2 + 2H2O

(R3)

where A is isopropyl alcohol (IPA) and B is acetic acid. Reaction kinetics is considered as first order with respect to the concentration of fuel. Kinetic constants are the same used in the previous models.15,16 They are listed in Table 1. The heats of reaction have been calculated for different temperatures using the PENG-ROBINSON EoS with BOSTONMATHIAS15 alpha function. Heat capacities have been calculated using the EoS of Anderko and Pitzer,19 which is an EoS specifically developed for reproducing volumetric properties and equilibrium data of aqueous systems at high temperature and pressure conditions. In a previous work, the parameters of the Anderko and Pitzer EoS were fitted for the air-water system.20 Using this EoS temperature could be predicted with an average deviation of 10.2% generally in excess. Thus, an overprediction of between 40-60 °C is also expected in the model presented here. The density of the mixture has been calculated using the Peng-Robinson EoS21 with the translated volume correction considering the exact composition of the mixture. Global transmission coefficients (U) have been estimated from experimental data. The coefficient has resulted of 150 W/m2 K between the mixer and the first chamber and of 350 W/m2 K the others. The coefficient for the heat transmission of the outer wall with the environment is taken as 0.5 W/m2 K. For the modeling, the reactor has been divided into several parts, with different flow patterns, as it can be seen in Scheme

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As the temperature profile should be calculated iteratively, for the first iteration TR is considered as 400 °C, and for the second iteration, this value is taken at the calculated temperature in the previous iteration for each z. 2. Top of the Reactor (CSTR). The mass balance for the isopropyl alcohol (A) and to the acetic acid (B) is shown in eqs 4-6.

Scheme 5. Scheme of the new CWR and Structure of the Modeling

CACSTR )

m0 FCSTR

CBCSTR )

FA0(1 - XAm) (kA(TR) + kAB(TR))3600V + 1000

FBm + 3600V(

∑β k

A AB(TR)CACSTR)

2000[(mo / FCSTR) + 3600VkB(TR)] FA0 - CACSTR

5. Results obtained with reactors 1 and 2 have been taken to validate the predictions of the model with the experimental temperature profiles measured in these reactors. The dimensions of these reactors, used to calculate the volumes of each section for application of the model, are presented in Table 2. • The mixer consists of a tube of 5 mm of diameter and 1.3 m length filled with alumina particles, as the one described for reactor 1 and 2. It has been modeled using plug flow and considering a porosity of 0.6. • Upper part of the reaction chamber. The reagents leave the mixer flowing upward, and then, they change the direction of flow, circulating downward through chamber 1. This is the reason why this part of the reactor is considered as the perfect mixing stage; thus, it is modeled as a CSTR. • Concentric reaction chambers. In this part the reagents flow in the annular gaps of the reaction chambers 1, 2, and 3 is described. Plug flow is considered. The dimensions of the concentric chambers are shown in Table 2. The governing equations of the modeling are described below. 1. Mixer. The mass balance for the isopropyl alcohol (A) and to the acetic acid (B) is shown in eq 1 and eq 2. -FA0

dXA ) εSm(RArA + βArAB) dz

dFB ) εSm(βiBriB - γBrB) dz

(1)

(2)

where the subindex A refers to IPA and the subindex B to acetic acid; ε is the porosity of the bed inside of the mixer; Sm is the section of the mixer; R, β, and γ are the stoichiometric coefficients of the reactions R1, R2, and R3, respectively; rij is the reaction rate; FA0 is the initial molar flow of IPA; X is the conversion; and z is the height of the reactor. The energy balance of the mixer is calculated as shown in eq 3.

( )

dTm ) εSm cpm0 dz

∑

XACSTR )

(3)

where cp is the specific heat of the mixture; m0 is the inlet mass flow (feed + oxygen); do is the external diameter of the mixer; Rij are the stoichiometric coefficients of reactions R1, R2, and R3; (-∆Hrij) are the heats of reaction of reaction R1,R2, and R3; TR1 is the temperature in the first reaction chamber; and Tm is the temperature in the mixer

(5)

m0 FCSTR

FA0

(6)

where C is concentration in mol/L and the subindex CSTR refers to the values inside the top part of the reactor; FA0 is the molar flow of IPA in the inlet of the reactor in mol/h; XAm is the conversion IPA at the outlet of the mixer; FCSRT is the density in inside the CSTR; βA is the stoichiometric coefficient of reaction R2; k is the kinetic constants; FBm is the molar flow of acetic acid in the outlet of the mixer in mol/h; m0 is the inlet mass flow (feed + oxygen); V is the volume of the CSTR in m3. It can be calculated with eq 7: V ) SLCSTR

(7)

where S is the section of the reactor and LCSTR is the length of the top part of the reactor considered as a CSTR. Experimentally, the CSTR length was fixed at 40 mm.15 The energy balance of the CSTR is shown in eq 8: [m0(-∆HT)in + V(

∑ k (T )C ij

R

) [m0(-∆HT)out]

iCSTR(-∆HR)ij)]

(8)

where ∆HT is the enthalpy of the total mixing at the inlet conditions (in) and at the outlet condition (out) in J/kg. Reaction Chambers. For a reaction chamber Rj (j ) 1, 2, and 3), the mass balance of the IPA (A) and of the acetic acid molar flows are calculated in every step using eqs 9 and 10. -FA0

dXA ) SRj(RArA + βArAB) dz

dFB ) SRj(βiBriB - γBrB) dz

(9)

(10)

The energy balance is expressed in eq 11.

(

.

cp m0 Rij(-∆Hrij)rij - Uπdo(Tm - TR1)

(4)

)

dTRj ) SΣ(-∆Hri)ri - Uπdoj(TRj - TR(j+1)) + dz Uπdo(j-1)(TR(j-1) - TRj)

(11)

As the temperature profile should be calculated iteratively, for the first iteration TR(j+1) is considered as 400 °C, and for the second iteration, this value is taken at the calculated temperature in the previous iteration for each z. For chamber 4, the chamber where the feed is preheated, only energy balance is considered, as there is no reaction in the chamber. It is shown in eq 12.


(

.

cp mfeed

)

dTR4 ) Uπdo3(TR3 - TR4) dz

(12)

where mfeed is the flow rate of the feed in kg/h, as the oxygen is still not present in this chamber. The strategy of resolution is shown in Scheme 6. First the balances in the mixer are solved. After that, balances in the CSTR section and in chambers 1 to 4 are solved. In the first iteration (i ) 1) a constant temperature of 400 °C inside the reactor is considered for calculating the heat transfer between reactor chambers. In subsequent iterations, the temperature profile of the previous iteration is used, that is, in the iteration i, the temperature profile calculated in iteration i-1 is used. The procedure is repeated until the differences between the temperature profiles calculated in two successive iterations are small enough. A minimum of three iterations is necessary to guarantee convergence. Comparison with Experimental Results. In the first place the predictions of the model for two experimental temperature profiles in stationary conditions measured for reactors 1 and 2 have been calculated, and the predicted temperature was compared to the experimental temperature profile in the first reaction chamber and to the outlet temperature of the reactor (final temperature of chamber 3). Temperatures were only measured in the salt chamber (outlet of chamber 4), reaction chamber 1, and at the outlet of the reactor (end of chamber 3). For making the comparison easier, the experimental temperature measured in the inlet of the mixer was fixed for simulating the performance of the reactor, instead of considering the outlet temperature of chamber 4 predicted by the model. Temperatures, conversions, and acetic acid flow (HAc) predicted by the model are shown in Figure 8 and compared to the experimental data. It was observed that the reaction began in the mixer where the oxidation of IPA produced a slight temperature increase, while a small amount of intermediate (acetic acid) was formed. Most of the reaction took place in the upper part of the chamber 1 (CSTR) where the complete oxidation of the IPA and the acetic acid caused the temperature to increase up to near 700 °C. In the rest of the chamber 1 and chambers 2 and 3 there was no reaction, and they were working only as a heat exchanger. Products cooled down very fast in chamber 1, but the temperature profile in chamber 2 and 3 was flat, so no heat exchange is expected between them. The existence of these two chambers reduces the efficiency of the heat exchange for feed preheating in chamber 4, but on the other hand contributes to isolate the temperature peaks in reaction chamber 1, protecting the pressure vessel from high temperatures, as observed in the experimental temperature profiles shown in Figure 2. In chamber 4 the feed was heated up to a temperature slightly above that of the entrance of the reaction chamber. The predicted temperature profiles matched well with the experimental ones, with the exception that the cooling predicted in chamber 1 is slower than that registered experimentally. Study of the Influence of the Feed Flow in the Behavior of the Reactor. A simulation of the reactor working with increasing feed flows was performed. Results for different feed flows are shown in Table 3. It was observed that when the feed flow rate was increased the maximum reaction temperature and reactor outlet temperature (final temperature in chamber 3) were increased too, while the preheating temperature of the feed (T inlet mixer) was lower. This is because a higher amount of organic matter was burnt inside the reactor, but as the exchange surface was the same, this made the preheating and the heat dissipation inside the reactor less efficient. It should be noted

6271

that in all the cases the concentration of pollutants at the end of the first chamber is zero, so chambers 2 and 3 are only working as a heat exchanger. In the case of feed flow of 200 kg/h (Figure 9) it was observed that the reaction was not completely finished in the upper part of the reactor, but it proceeded also for a few centimeters along chamber 1, where a peak in the concentration of the intermediate was observed. For this, feed flow temperature profiles in chambers 2 and 3 were not so similar. For higher flows the design is more efficient, increasing the surface for heat transmission and isolating the pressure vessel of the temperature peaks in reaction chamber 1. Study of the Influence of the Inlet Temperature. The influence of the inlet temperature in the outer preheating chamber (chamber 4) was studied when the reactor was working at the maximum flow rate, 200 kg/h, as shown in Table 4. For doing so, the inlet temperature in the mixer was considered as the outlet temperature of chamber 4. The simulation showed that the pollutants were completely destroyed for inlet temperatures of 300 and 250 °C, but at inlet temperatures of 200 °C, temperatures inside the reactor were too low and the acetic acid could not react even in chambers 2 or 3. Thus, according to the model the lower inlet temperature for the reaction to completely oxidize a feed of 6 (w/w) % IPA is 200 °C. It was observed in Figure 2 that the reactor can operate autothermally with inlet temperatures lower than 200 °C with feed flows of 30 kg/h using air as the oxidant. It was shown as well that when the flow was increased the heat transfer became less efficient in the reactor. It thus seems reasonable that for higher feed flows it will be needed to introduce the feed in the reactor at higher temperatures. Study of the Influence of the Fuel Concentration on the Behavior of the Reactor. Model results showed that the elimination of pollutants in the feed depended on the temperature reached in the reactor. If the inlet temperature was reduced the reaction temperature was also reduced and the pollutants were not eliminated. But the reaction temperature does not depend exclusively on the inlet temperature, but also depends strongly on the fuel concentration. Thus, in theory the inlet temperature can be reduced if the fuel concentration is higher and the fuel concentration can be reduced if the inlet temperature is high enough. To illustrate this effect some simulations were made reducing the fuel (IPA) concentration and keeping the inlet temperature at 300 °C. The results of the simulations are shown in Table 5. It is observed that when the IPA concentration was reduced, the reaction temperature and outlet temperature were also reduced, and the reagents needed more reactor length to be oxidized. With a 4 (w/w) % IPA the reaction temperature attained was so low that the reaction was still proceeding in chambers 2 and 3, and the residence time inside the reactor was not enough for completing the reaction. Experimentally, working at feed flows up to 40 kg/h the lower IPA concentrations used before the reaction was extinguished was 5-5.5 (w/ w) %. Even when the model predicted a stationary state with a reaction temperature of 460 °C and an inlet temperature of 354 °C, experimentally if the temperature attained in the salt chamber is lower than 350 °C and the reaction temperature is lower than 600 °C it is very difficult to keep the reactor working autothermally. Summarizing, the reactor was modeled using a steady-state model that considers simple flow patterns. Mass and energy balance was solved and Anderko-Pitzer EoS was used to calculate the heat capacities of the oxygen-water mixture. The model is able to describe satisfactorily the behavior of the

6272


reactor, reproducing some experimental data, and was used to extrapolate the reactor performance in other operational conditions: the reactor was predicted to be able to get a complete TOC removal of a feed consisting of 6 (w/w) % IPA, even for the highest flow of the demonstration plant (200 kg/h). Even at this high feed flow, the reaction took place in the most internal concentric chamber, the rest of the reactor acting as a heat exchanger. The model showed concentrations of 4 (w/w) % IPA or inlet temperatures of 200 °C as the limiting operational conditions of the reactor at this maximum flow. The model can predict the most appropriate concentration-inlet temperature combination for the destruction of a given waste. Conclusions The new cooled-wall reactor has been designed as a heat exchanger, with several concentric chambers maximizing the heat transfer surface and obtaining a further isolation of the pressure vessel from the high temperatures in the reaction chamber compared to the previous CWR. Five prototypes of this reactor have been constructed and tested using isopropyl alcohol and ammonia as a fuel and air and oxygen as an oxidant. When correctly operated the reactor was able to preheat feeds from 200 °C to supercritical temperature. The effects of temperature were isolated inside of the reaction chamber, not registering temperatures higher than 400 °C in the pressure shell. TOC conversions as high as 99.8% were obtained in some of the prototypes. The salt precipitation chamber was effective in collecting a certain amount of salts during steady operation. Unfortunately for high salt concentrations salt precipitation caused fowling in the inner reactions chamber and mixer. Obstructions in the mixer caused high pressure drops that made O2 delivery difficult, and in some cases a complete plug of the mixer formed that led to the collapse of internal components of the reactor. The present reactor design is in need of some modifications to be appropriate for the destruction of feeds with high salt contents. The reactor was modeled using a steady-state model that considers simple flow patterns. Despite its simplicity the model is able to describe satisfactorily the behavior of the reactor, reproducing some experimental data. Some simulations are presented to illustrate the effect of the operation variables. According to the results predicted by the simulation the reactor was able to get a complete TOC removal of a feed consisting of 6 (w/w)% IPA, even for the highest flow of the demonstration plant (200 kg/h). All the reactions took place in the most internal concentric chamber, the rest of the reactor acting as a heat exchanger. If the concentration was reduced down to 4 (w/w) % IPA or the inlet temperature was reduced down to 200 °C, the temperature reached inside of the reactor was not high enough and the reaction was not completed. The model can be used to choose the most appropriate concentration-inlet temperature combination for the destruction of a given waste. Acknowledgment The authors thank CETRANSA for providing technical and financial support and the Spanish Ministry of Science and Innovation project, PET 2006-0376.

Supporting Information Available: Figures 10, 11, and 12 showing pollutants elimination. This material is available free of charge via the Internet at http://pubs.acs.org. Literature Cited (1) Modell, M. Processing Methods for the Oxidation of Organics in Supercritical Water, U.S. Patent. No. 4338199 (1982). (2) Marrone, P. A., Hong, G. T. Supercritical Water Oxidation. In EnVironmentally Conscious Materials and Chemicals Processing; Kutz, M., Ed.; John Wiley & Sons Inc: Hoboken, NJ, 2007. (3) Bermejo, M. D.; Cocero, M. J. Supercritical Water Oxidation: A Technical Review. AIChE J. 2006, 52, 3933. (4) Marrone, P. A.; Cantwell, S. D.; Dalton, D. W. SCWO System Designs for Waste TreatmentsApplication to Chemical Weapons Destruction. Ind. Eng. Chem. Res. 2005, 44, 9030. (5) Griffith, J. W.; Raymond, D. H. The First Commercial Supercritical Water Oxidation Sludge Processing Plant. Waste Manage. 2002, 22, 453. (6) Crooker, P. I.; Ahluwalia, K. S.; Fan, Z.; Prince, J. Operating Results from Supercritical Water Oxidation Plants. Ind. Eng. Chem. Res. 2000, 39, 4865. (7) Rice, S. F., Wu, B. J., Winters, W. S. Engineering Modeling of the Pine Bluff Arsenal Supercritical Water Oxidation Reactor, Proceedings of 5th International Symposium on Supercritical Fluids, Atlanta, GA, April 8-12, 2000. (8) http://www.scfi.eu/aquacritox.html (accessed May 5, 2009). (9) Cocero, M. J.; Alonso, E.; Torıó, R.; Vallelado, D.; Fdz-Polanco, F. Supercritical Water Oxidation in Pilot Plant of Nitrogenous CompoundsIsopropanol Mixtures in the Temperature Range 500-750°C. J. Ind. Eng. Chem. Res. 2000, 39, 3707. (10) Cocero, M. J.; Alonso, E.; Vallelado, D.; Torıó, R.; Fdz-Polanco, F. Optimization of Operational Variables of a Supercritical Water Oxidation SCWO Process. Water Sci. Technol. 2000, 42, 107. (11) Cocero, M. J.; Alonso, E.; Torıó, R.; Vallelado, D.; Sanz, T.; FdzPolanco, F. Supercritical Water Oxidation (SCWO) for Polyethylene Terephthalate (PET) Industry Effluents. Ind. Eng. Chem. Res. 2000, 39, 4652. (12) Cocero, M. J.; Martıń, A., Bermejo, M. D.; Santos, M.; Rincoń, D.; Alonso, E. ; Fdez-Polanco, F. Supercritical Water Oxidation of Industrial Waste from Pilot to Demonstration Scale, 6th International Symposium on Supercritical Fluids, Versailles, France, 28-30th April, 2003. (13) Oh, C. H.; Kochan, R. J.; Charlton, T. R.; Bourhis, A. L. ThermalHydraulic Modeling of Supercritical Water Oxidation of Phenol. Energy Fuels 1996, 10, 326. (14) Vielcazals, S.; Mercadier, J.; Marias, F.; Mateos, D.; Bottreau, M.; Cansell, F.; Marraud, C. Modeling and Simulation of Hydrothermal Oxidation of Organic Compounds. AIChE J. 2006, 52, 818. (15) Bermejo, M. D.; Fdz-Polanco, F.; Cocero, M. J. Modeling of a Transpiring Wall Reactor for the Supercritical Water Oxidation Using Simple Flow Patterns: Comparison to Experimental Results. Ind. Eng. Chem. Res. 2005, 44, 3835. (16) Cocero, M. J.; Martinez, J. L. Cool-Wall Reactor for Supercritical Water Oxidation: Modelling and Operation Results. J. Supercrit. Fluids 2004, 31, 41. (17) Bermejo, M. D.; Cantero, F; Cocero, M. J. Supercritical Water Oxidation of Feeds with High Ammonia Concentrations: Pilot Plant Experimental Results and Modeling. Chem. Eng. J. 2008, 137, 542. (18) Li, L.; Chen, P.; Gloyna, E. F. Generalized Kinetic Model for Wet Oxidation of Organic Compounds. AIChE J. 1991, 37, 1687. (19) Anderko, A.; Pitzer, K. S. EOS Representation of Phase Equilibria and Volumetric Properties of the System NaCl-H2O above 573 K. Geochim. Cosmochim. Acta 1993, 57, 1657. (20) Bermejo, M. D.; Martıń, A.; Cocero, M. J. Application of the Anderko-Pitzer EoS to the Calculation of Thermodynamical Properties of Systems Involved in the Supercritical Water Oxidation Process. J. Supercrit. Fluids 2007, 42, 27. (21) Magoulas, C.; Tassios, D. Thermophysical Properties of n-Alkanes from C1 to C20 and Their Prediction for Higher ones. Fluid Phas. Equilib. 1990, 56, 119.

ReceiVed for reView October 29, 2008 ReVised manuscript receiVed May 8, 2009 Accepted May 11, 2009 IE900054E

Experimental Performance and Modeling of a New Cooled-Wall

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